E L E C T R O N I C A C C E S S http ://mit. imt. s i ISSN 1580-2949 M IN M ATER. TEH N O L. LETN IK VO LU M E [TEV. N O . STR. P. LJU BLJA N A SLO VEN IJA SEP.-O KT. SE P.-O C T. 51 5 707-879 ISSN: 1580-2949 UDK: 669+666+678+53 2017 5 MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY The journal MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY so znanstvena serijska publikacija, ki objavlja izvirne in tudi pregledne znanstvene ~lanke ter tehni~ne novice, ki obravnavajo teoreti~na in prakti~na vpra{anja naravoslovnih ved in tehnologije na podro~jih kovinskih in anorganskih materialov, polimerov, vakuumske tehnike in v zadnjem ~asu tudi nano- materialov. is a scientific journal, devoted to original scientific papers, reviewed scientific papers and technical news concerned with the areas of fundamental and applied science and technology. Topics of particular interest include metallic materials, inorganic materials, polymers, vacuum technique and lately nanomaterials. IMT), Jaka Burja (IMT), Monika Jenko, Varu`an Kevorkijan (IMPOL), Aleksandra Kocijan (IMT), Andra` Legat (ZAG), Vojteh Leskov{ek (IMT), Matja` Godec (IMT), Paul McGuiness (IMT), Djordje Mandrino (IMT), Bo{tjan Markoli (NTF), Jo`ef Medved (NTF), Peter Panjan (IJS), Irena Paulin (IMT), Danijela A. 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Po bazi podatkov JCR16 ima Materiali in Tehnologije dejavnik vpliva 0,436. / In JCR16 Database Mater.Tehnol. has an impact factor of 0.436. ^lanki objavljeni v periodi~ni publikaciji MATERIALI IN TEHNOLOGIJE so indeksirani v mednarodnih sekundarnih virih: (Articles published in journal are indexed in international secondary periodicals and databases): Izhajanje: Published: Naro~nina / Subscription: tujina / abroad: Naslov uredni{tva (Editorial Address): Glavni in odgovorni urednik (Editor-in-Chief): Pomo~nik glavnega urednika (Associate Editor-in-Chief): ^astna glavna urednika (Honorary Editors-in-Chief): Souredniki (Co-Editors): Tehni~ni urednik (Technical Editor): Lektorji (Linguistic Advisers): Mednarodni pridru`eni ~lani uredni{kega odbora (International Advisory Board): Paul John McGuiness Matja` Godec Franc Vodopivec, Matja` Torkar Igor Beli~ (IMT), Jaka Burja (IMT), Aleksandra Kocijan (IMT), Djordje Mandrino (IMT), Bo{tjan Markoli (NTF), Irena Paulin (IMT), Danijela A. Skobir Balanti~ (IMT), Darja Steiner Petrovi~ (IMT), Bojan Podgornik (IMT), Sre~o [kapin (IJS), Rok Zaplotnik (IJS), Ema @agar (KI) Erika Nared (IMT) Erika Nared (IMT) (slovenski jezik), Paul John McGuiness (IMT) (angle{ki jezik) Edvard Sla~ek ( Marko Drobni~ ( Andrej Gradi{nik ( – DOAJ (Directory of Open Access Journals) – Civil Engineering Abstracts – Google Scholar – Ceramic Abstracts/World Ceramic Abstracts – SCIRUS – Corrosion Abstracts – CA SEARCH® – Chemical Abstracts® – Mechanical & Transportation Engineering Abstracts – METADEX® – CSA Aerospace & High Technology Database – TEME – Technology and management – Solid State and Superconductivity Abstracts – Inside Conferences – Materials Business File – Engineered Materials Abstracts® – Referativnyj `urnal: Metallurgija – Aluminium Industry Abstracts – COBIB – SCOPUS ® Leonid B. Getsov, NPO CKT, St. Petersburg, Russia • Bo`o Smoljan, University of Rijeka, Croatia • David Nolan, Bluescope Steel Ltd. & University of Wollongong, Wollongong, Australia • Karlo T. Rai}, University of Belgrade, Faculty of Technology and Metallurgy, Belgrade, Serbia • Nicola Gargiulo, Engineering University of Naples, Naples, Italy • Francesco Colangelo, Parthenope University of Naples, Naples, Italy • Peter Jur~i, Faculty of Materials Science and Technology, STU, Trnava, Slovakia • Smilja Markovi}, Institute of Technical Sciences of the Serbian Academy of Sciences and Arts, Belgrade, Serbia • Stefan Zaefferer, Max-Planck Institute for Steel Research, Dusseldorff, Germany • Urban Wiklund, Uppsala University of Sweden, Sweden • Zdenka Zovko Brodarac, University of Zagreb, Metallurgical Faculty, Sisak, Croatia • Bojana Dolinar, Faculty of Civil Engineering, Transportation Engineering and Architecture, University of Maribor, Slovenia • Ivan Nazarenko, Kyiv National University of Construction and Architecture, Ukraine M EHNOLOGIJEIN AT E R IALI 5 M A T E R I A L S A N D T E C H N O L O G Y M ATER. TEH N O L. LETN IK VO LU M E [TEV. N O . STR. P. LJU BLJA N A SLO VEN IJA SEP.-O KT. SEP.-O C T. 51 5 707-879 ISSN : 1580-2949 U D K: 669+666+678+53 Na INTERNET-u je revija MATERIALI IN TEHNOLOGIJE dosegljiva na naslovu (ELECTRONIC ACCESS): http://mit.imt.si Elektronska po{ta (E-mail): Oblikovanje ovitka (Design): Oblikovanje plakata na naslovnici (Poster on the Cover): Ra~unalni{ki prelom in tisk (Prepress and Printed by): Naklada (Circulation): mit imt.si Ignac Kofol Ajda Schmidt NONPAREL grafi~ne storitve d.o.o., Medvode 400 izvodov/issues @ 2017 VSEBINA – CONTENTS PREGLEDNI ^LANEK – REVIEW ARTICLE Additive manufacturing: the future of manufacturing Dodajalna (3D) tehnologija: prihodnost proizvajanja S. A. Adekanye, R. M. Mahamood, E. T. Akinlabi, M. G. Owolabi . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 709 Pomembna obletnica revije Materiali in tehnologije: petdeset let izhajanja znanstvene periodi~ne publikacije An important anniversary of the Materials and Technology journal: fifty years of publication E. Nared . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 717 IZVIRNI ZNANSTVENI ^LANKI – ORIGINAL SCIENTIFIC ARTICLES Investigation of grain boundaries in Alloy 263 after special heat treatment Preiskava mej zrn v zlitini 263 po posebni toplotni obdelavi I. Slatkovský, M. Dománková, M. Sahul . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 721 Fracture toughness of ledeburitic Vanadis 6 steel after sub-zero treatment for 17 h and double tempering Lomna `ilavost ledeburitnega jekla Vanadis 6 po toplotni obdelavi s 17-urnim podhlajevanjem in dvojnim popu{~anjem J. Pta~inová, P. Jur~i, I. Dlouhý . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 729 Electronic and optical properties of the spinel oxides MgxZn1-xAl2O4 by first-principles calculations Elektronske in opti~ne lastnosti spinelnih oksidov MgxZn1-xAl2O4, izpeljane iz teoreti~nih osnov C. Xiang, J. X. Zhang, Y. Lu, D. Tian, C. Peng . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 735 Surface characteristics of Invar alloy according to micro-pulse electrochemical machining Karakteristike povr{ine Invar zlitine glede na mikropulzno elektrokemi~no obdelavo S.-H. Kim, S.-G. Choi, W.-K. Choi, E.-S. Lee . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 743 Durability of materials based on a polymer-silicate matrix and a lightweight aggregate exposed to aggressive influences combined with high temperatures Vzdr`ljivost materialov na osnovi iz polimer-silikatnih matric in lahkega dodatka, izpostavljenih agresivnim vplivom v kombinaciji z visokimi temperaturami T. Melichar, J. Byd`ovský, Á. Dufka . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 751 Influence of thermomechanical treatment on the grain-growth behaviour of new Fe-Al based alloys with fine Al2O3 precipitates Vpliv termomehanske obdelave FeAl zlitin s finimi Al2O3 izlo~ki na rast zrn B. Ma{ek, O. Khalaj, H. Jirková, J. Svoboda, D. Bublíková . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 759 Analysis of precipitates in aluminium alloys with the use of high-resolution electron microscopy and computer simulation Raziskave oborin v aluminijevih zlitinah z visokoresolucijsko elektronsko mikroskopijo in ra~unalni{ko simulacijo K. Matus, A. Tomiczek, K. Go³ombek, M. Pawlyta . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 769 Microstructural evaluation of Ni-SDC cermet from a representative 2D image and/or a 3D reconstruction based on a stack of images Vrednotenje mikrostruktur Ni-SDC kermeta z 2D in/ali 3D metodo G. Kapun, M. Marin{ek, F. Merzel, S. [turm, M. Gaber{~ek, T. Skalar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 775 A facile method to prepare super-hydrophobic surfaces on silicone rubbers Preprosta metoda za pripravo superhidrofobnih povr{in pri silikonskih gumah H. Y. Jin, Y. F. Li, S. C. Nie, P. Z., N. K. Gao, W. Li. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 783 Investigation of the static icing property for super-hydrophobic coatings on aluminium Preiskava lastnosti stati~ne zaledenitve pri superhidrofobnih prevlekah na aluminiju H. Y. Jin, S. C. Nie, Y. F. Li, T. F. Xu, P. Zhang, W. Li . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 789 Effect of ball milling on the properties of the porous Ti–26Nb alloy for biomedical applications Vpliv krogli~nega mletja na lastnosti porozne zlitine Ti–26Nb za biomedicinske aplikacije G. Dercz, I. Matu³a . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 795 Effects of an addition of coir-pith particles on the mechanical properties and erosive-wear behavior of a wood-dust-particle-reinforced phenol formaldehyde composite Vplivi dodatka kokosovih vlaken fenol-formaldehidnemu kompozitu, oja~anem z lesnim prahom, na njegove mehanske lastnosti in erozijsko obrabo A. S. Jose, A. Athijayamani, K. Ramanathan, S. Sidhardhan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 805 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS ISSN 1580-2949 UDK 669+666+678+53 MTAEC9, 51(5)707–879(2017) MATER. TEHNOL. LETNIK VOLUME 51 [TEV. NO. 5 STR. P. 707–879 LJUBLJANA SLOVENIJA SEP.–OKT. SEP.–OCT. 2017 Optimum bushing length in thermal drilling of galvanized steel using artificial neural network coupled with genetic algorithm Optimalna dol`ina podpore ({ablone, vodila) pri termi~nem vrtanju galvaniziranega jekla z uporabo umetne nevronske mre`e in genetskega algoritma N. Rajesh J. Hynes, R. Kumar, J. A. J. Sujana . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 813 Gelling polysaccharide as the electrolyte matrix in a dye-sensitized solar cell @elirni polisaharid kot elektrolitna osnova v solarnih celicah, ob~utljivih na barvila J. P. Bantang, D. Camacho . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 823 Development of a heat treatment for increasing the mechanical properties and stress corrosion resistance of 7000 Al alloys Razvoj toplotne obdelave za izbolj{anje mehanskih lastnosti in napetostno korozijsko odpornost 7000 Al zlitin M. Shakouri, M. Esmailian, S. Shabestari . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 831 Corrosion resistance of as-plated and heat-treated electroless dublex Ni-P/Ni-B-W coatings Korozijska odpornost platiranih in neelektri~no topolotno obdelanih dupleks Ni-P/Ni-B-W prevlek B. Yüksel, G. Erdogan, F. E. Bastan, R. A. Yýldý . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 837 Short-term creep of P91 heat-resistant steels at low stresses and an instantaneous-stress-change testing Kratkotrajno lezenje toplotno odpornega jekla P91 pri nizkih napetostih in nenadni menjavi napetosti obremenjevanja J. Zhe, S. Junjie, Z. Pengshuo . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 843 Effect of severe plastic and heavy cold deformation on the structural and mechanical properties of commercially pure titanium U~inek plasti~nosti in deformacije pri podhlajevanju na strukturne in mehanske lastnosti ~istega komercialnega titana J. Palán, P. [utta, T. Kubina, M. Dománková . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 849 Effect of yttrium and zirconium microalloying on the structure and properties of weld joints of a two-phase titanium alloy U~inek mikrolegiranja itrija in cirkonija na strukturo in lastnosti na spoje zavrov dvofazne zlitine titana A. Illarionov, A. Popov, S. Illarionova, D. Gadeev . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 855 Microstructure evolution and statistical analysis of Al/Cu friction-stir spot welds Razvoj mikrostrukture in statisti~na analiza vrtilno-tornih to~kastih zvarov Al/Cu M. P. Mubiayi, E. T. Akinlabi, M. E. Makhatha . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 861 Synthesis of PMMA/ZnO nanoparticles composite used for resin teeth Sinteza PMMA/ZnO nanodelcev kompozitov za izdelavo zob iz umetnih smol D. Popovi}, R. Bobovnik, S. Bolka, M. Vukadinovi}, V. Lazi}, R. Rudolf . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 871 ERRATUM . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 879 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING 709–715 ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING DODAJALNA (3D) TEHNOLOGIJA: PRIHODNOST PROIZVAJANJA Sheriff Adefemi Adekanye1, Rasheedat Modupe Mahamood1,2, Esther Titilayo Akinlabi2, Moses Gbadebo Owolabi3 1University of Ilorin, Department of Mechanical Engineering, Nigeria 2University of Johannesburg, Department of Mechanical Engineering Science, South Africa 3Howard University, Department of Mechanical Engineering, Washington, USA adekanyefm@gmail.com Prejem rokopisa – received: 2016-08-26; sprejem za objavo – accepted for publication: 2017-03-16 doi:10.17222/mit.2016.261 Additive manufacturing is an advanced manufacturing method used to fabricate prototypes, tooling as well as functional products. Additive manufacturing helps us produce complex parts as single-unit objects, which was not possible with the traditional manufacturing methods. There are different types of additive-manufacturing technologies, including selective laser melting, laser-metal-deposition process, fused deposition modeling and electron-beam melting. All these additive-manu- facturing technologies produce three-dimensional (3D) objects by adding materials layer after layer. A 3D object is built directly from a 3D computer-aided-design (CAD) model of the object. Additive manufacturing is a very promising method for the aerospace industry, in particular because of its ability to reduce the buy-to-fly ratio. This technology is the technology of the future because it is going to change the way products are designed and manufactured. In this research, various additive-manu- facturing technologies are described in detail and some of the research works in this field are also presented. The future research directions are also highlighted. Keywords: additive manufacturing, fused deposition modelling, laser metal deposition, selective laser melting, selective laser sintering Dodajalni proizvodni proces je napredna tehnologija, ki se uporablja za izdelavo prototipov, orodij in funkcionalnih izdelkov. Z naprednim proizvodnim procesom lahko izdelamo izdelke ali dele kompliciranih oblik, ki jih ni mo`no izdelati s tradicionalnimi metodami proizvodnje. Obstajajo razli~ne vrste naprednih dodajalnih tehnologij, kot so na primer: selektivno lasersko nataljevanje, proces laserskega nana{anja (depozicije) kovin, modeliranje z nana{anjem staljenih kapljic in taljenje z elektronskim curkom. Vse te napredne dodajalne tehnologije omogo~ajo izdelavo tridimenzionalnih (3D) izdelkov z dodajanjem materialov plast za plastjo. Tridimenzionalni produkt je izgrajen neposredno s pomo~jo 3D-ra~unalni{ko podprtega modeliranja (angl. CAD). Napredna dodajalna proizvodnja je zelo obetavna proizvodna metoda, predvsem v letalski in vesoljski industriji, {e posebej zaradi sposobnosti zmanj{evanja stro{kov izdelave (zmanj{anje razmerja med maso materiala, potrebnega za izdelavo dolo~enega izdelka in njegovo dejansko maso). Klju~ne besede: dodajalna proizvodnja, modeliranje z nana{anjem staljenih materialov, laserska depozicija kovin, selektivno lasersko nataljevanje, selektivno lasersko sintranje 1 INTRODUCTION The additive-manufacturing process is an advanced manufacturing process that produces three-dimensional (3D) objects directly from the 3D computer-aided-design (CAD) digital information of the parts by adding ma- terials layer by layer.1–3 This is different from the tradi- tional manufacturing that involves material removal or an energy-intensive process such as machining, casting and forging. With additive manufacturing, a complex-shaped product can be produced as a single object, which was not possible in the past. For the traditional manufacturing process, complex parts have to be designed based on the ease of manufact- uring the parts. Complex parts are usually broken down into smaller parts because of the ease of manufacturing those parts and the pieces are assembled at the later stage through various joining processes. This practice does not only involve labor-intensive processes but also results in wastage of materials, in addition to the materials re- moved during shaping and cutting operations; the excess materials added for joining also contribute to material wastage. The net weight of a product is also very large because of the excess materials used during the assembly. The additive-manufacturing technology, on the other hand, can make a product directly from the 3D CAD model of the product by just adding materials layer by layer. Irrespective of the complexity, any part that can be drawn using any computer-aided-design software can be made and as a single-unit object. Additive-manu- facturing technologies are classified into seven main groups, namely: powder-bed fusion, directed energy deposition, sheet lamination, material extrusion, binder jetting, material jetting and vat photopolymerization.1 A designer does not need to worry about the manufactur- ability of a product; he/she is only concerned with the functionality of the part being designed. At the inception, the additive manufacturing process was used to make prototypes because of the limitation in Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 709 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 620.183.183.4:621.039.337:316.422.44 ISSN 1580-2949 Review article/Pregledni ~lanek MTAEC9, 51(5)709(2017) Optimum bushing length in thermal drilling of galvanized steel using artificial neural network coupled with genetic algorithm Optimalna dol`ina podpore ({ablone, vodila) pri termi~nem vrtanju galvaniziranega jekla z uporabo umetne nevronske mre`e in genetskega algoritma N. Rajesh J. Hynes, R. Kumar, J. A. J. Sujana . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 813 Gelling polysaccharide as the electrolyte matrix in a dye-sensitized solar cell @elirni polisaharid kot elektrolitna osnova v solarnih celicah, ob~utljivih na barvila J. P. Bantang, D. Camacho . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 823 Development of a heat treatment for increasing the mechanical properties and stress corrosion resistance of 7000 Al alloys Razvoj toplotne obdelave za izbolj{anje mehanskih lastnosti in napetostno korozijsko odpornost 7000 Al zlitin M. Shakouri, M. Esmailian, S. Shabestari . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 831 Corrosion resistance of as-plated and heat-treated electroless dublex Ni-P/Ni-B-W coatings Korozijska odpornost platiranih in neelektri~no topolotno obdelanih dupleks Ni-P/Ni-B-W prevlek B. Yüksel, G. Erdogan, F. E. Bastan, R. A. Yýldý . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 837 Short-term creep of P91 heat-resistant steels at low stresses and an instantaneous-stress-change testing Kratkotrajno lezenje toplotno odpornega jekla P91 pri nizkih napetostih in nenadni menjavi napetosti obremenjevanja J. Zhe, S. Junjie, Z. Pengshuo . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 843 Effect of severe plastic and heavy cold deformation on the structural and mechanical properties of commercially pure titanium U~inek plasti~nosti in deformacije pri podhlajevanju na strukturne in mehanske lastnosti ~istega komercialnega titana J. Palán, P. [utta, T. Kubina, M. Dománková . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 849 Effect of yttrium and zirconium microalloying on the structure and properties of weld joints of a two-phase titanium alloy U~inek mikrolegiranja itrija in cirkonija na strukturo in lastnosti na spoje zavrov dvofazne zlitine titana A. Illarionov, A. Popov, S. Illarionova, D. Gadeev . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 855 Microstructure evolution and statistical analysis of Al/Cu friction-stir spot welds Razvoj mikrostrukture in statisti~na analiza vrtilno-tornih to~kastih zvarov Al/Cu M. P. Mubiayi, E. T. Akinlabi, M. E. Makhatha . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 861 Synthesis of PMMA/ZnO nanoparticles composite used for resin teeth Sinteza PMMA/ZnO nanodelcev kompozitov za izdelavo zob iz umetnih smol D. Popovi}, R. Bobovnik, S. Bolka, M. Vukadinovi}, V. Lazi}, R. Rudolf . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 871 ERRATUM . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 879 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING 709–715 ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING DODAJALNA (3D) TEHNOLOGIJA: PRIHODNOST PROIZVAJANJA Sheriff Adefemi Adekanye1, Rasheedat Modupe Mahamood1,2, Esther Titilayo Akinlabi2, Moses Gbadebo Owolabi3 1University of Ilorin, Department of Mechanical Engineering, Nigeria 2University of Johannesburg, Department of Mechanical Engineering Science, South Africa 3Howard University, Department of Mechanical Engineering, Washington, USA adekanyefm@gmail.com Prejem rokopisa – received: 2016-08-26; sprejem za objavo – accepted for publication: 2017-03-16 doi:10.17222/mit.2016.261 Additive manufacturing is an advanced manufacturing method used to fabricate prototypes, tooling as well as functional products. Additive manufacturing helps us produce complex parts as single-unit objects, which was not possible with the traditional manufacturing methods. There are different types of additive-manufacturing technologies, including selective laser melting, laser-metal-deposition process, fused deposition modeling and electron-beam melting. All these additive-manu- facturing technologies produce three-dimensional (3D) objects by adding materials layer after layer. A 3D object is built directly from a 3D computer-aided-design (CAD) model of the object. Additive manufacturing is a very promising method for the aerospace industry, in particular because of its ability to reduce the buy-to-fly ratio. This technology is the technology of the future because it is going to change the way products are designed and manufactured. In this research, various additive-manu- facturing technologies are described in detail and some of the research works in this field are also presented. The future research directions are also highlighted. Keywords: additive manufacturing, fused deposition modelling, laser metal deposition, selective laser melting, selective laser sintering Dodajalni proizvodni proces je napredna tehnologija, ki se uporablja za izdelavo prototipov, orodij in funkcionalnih izdelkov. Z naprednim proizvodnim procesom lahko izdelamo izdelke ali dele kompliciranih oblik, ki jih ni mo`no izdelati s tradicionalnimi metodami proizvodnje. Obstajajo razli~ne vrste naprednih dodajalnih tehnologij, kot so na primer: selektivno lasersko nataljevanje, proces laserskega nana{anja (depozicije) kovin, modeliranje z nana{anjem staljenih kapljic in taljenje z elektronskim curkom. Vse te napredne dodajalne tehnologije omogo~ajo izdelavo tridimenzionalnih (3D) izdelkov z dodajanjem materialov plast za plastjo. Tridimenzionalni produkt je izgrajen neposredno s pomo~jo 3D-ra~unalni{ko podprtega modeliranja (angl. CAD). Napredna dodajalna proizvodnja je zelo obetavna proizvodna metoda, predvsem v letalski in vesoljski industriji, {e posebej zaradi sposobnosti zmanj{evanja stro{kov izdelave (zmanj{anje razmerja med maso materiala, potrebnega za izdelavo dolo~enega izdelka in njegovo dejansko maso). Klju~ne besede: dodajalna proizvodnja, modeliranje z nana{anjem staljenih materialov, laserska depozicija kovin, selektivno lasersko nataljevanje, selektivno lasersko sintranje 1 INTRODUCTION The additive-manufacturing process is an advanced manufacturing process that produces three-dimensional (3D) objects directly from the 3D computer-aided-design (CAD) digital information of the parts by adding ma- terials layer by layer.1–3 This is different from the tradi- tional manufacturing that involves material removal or an energy-intensive process such as machining, casting and forging. With additive manufacturing, a complex-shaped product can be produced as a single object, which was not possible in the past. For the traditional manufacturing process, complex parts have to be designed based on the ease of manufact- uring the parts. Complex parts are usually broken down into smaller parts because of the ease of manufacturing those parts and the pieces are assembled at the later stage through various joining processes. This practice does not only involve labor-intensive processes but also results in wastage of materials, in addition to the materials re- moved during shaping and cutting operations; the excess materials added for joining also contribute to material wastage. The net weight of a product is also very large because of the excess materials used during the assembly. The additive-manufacturing technology, on the other hand, can make a product directly from the 3D CAD model of the product by just adding materials layer by layer. Irrespective of the complexity, any part that can be drawn using any computer-aided-design software can be made and as a single-unit object. Additive-manu- facturing technologies are classified into seven main groups, namely: powder-bed fusion, directed energy deposition, sheet lamination, material extrusion, binder jetting, material jetting and vat photopolymerization.1 A designer does not need to worry about the manufactur- ability of a product; he/she is only concerned with the functionality of the part being designed. At the inception, the additive manufacturing process was used to make prototypes because of the limitation in Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 709 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 620.183.183.4:621.039.337:316.422.44 ISSN 1580-2949 Review article/Pregledni ~lanek MTAEC9, 51(5)709(2017) the choice of the materials that could be used on an additive-manufacturing machine at that time.2–4 There are a number of industries benefiting from this revolutionary manufacturing technology,5 such as the aerospace, auto- mobile and biomedical industries.6–9 In this study, some of these additive-manufacturing technologies are re- viewed with the aim of throwing light on these techno- logies and describing their systems of operation as well as their advantages, limitations and areas of applications. The rest of the paper is organized as follows: section 2 of the paper presents the powder-bed-based process. The extrusion-based process and sheet-lamination pro- cess are presented in sections 3 and 4, respectively. Di- rected-energy-deposition process is presented in section 5. Material development in additive manufacturing is presented in section 6, while conclusions are presented in section 7. 2 POWDER-BED-BASED PROCESS This class of additive manufacturing technology uses the energy from an electron beam or laser beam to fuse or melt powder materials for the manufacturing of parts.10,11 The additive-manufacturing technologies in this class include direct metal laser sintering (DMLS), electron-beam melting (EBM), selective heat sintering (SHS), selective laser melting (SLM) and selective laser sintering (SLS). Each of these technologies is described in this section. 2.1 Direct metal laser sintering (DMLS) A schematic diagram of direct laser sintering is shown in Figure 1 below. This process manufactures parts by compacting a metal powdered material with a power source such as laser, without melting but by binding the material together to create a solid structure defined by the 3D CAD model of the part.12 The working principle of this technology makes it capable of design- ing and producing complex geometry of both internal and external intricacies. In this process, to avoid the collision of the recoater blade when moving, the building and dispenser platforms are lowered by the layer thickness. The powder metal needed to create a layer of the material is supplied by the dispenser platform after it has been ensured that the recoater stands in the right position. The spreading of the metal powder from the dispenser to the building platform is done by the re- coater, moving from the right to the left position, and the excess metal powder falls into the excess-powder collector. The scan head moves the laser beam through the two-dimensional (2D) cross-section, which is switched on and off during the exposure of designated areas. The powdered metals are generated, cured and sintered by the solidified areas through the absorption of energy.13 This process continues to create layer after layer until the part is completed. Using this process, parts have been successfully made from the materials such as aluminium alloys – AlSi10Mg, stainless steel, titanium alloys and cobalt chrome superalloy.13 Some of the advantages of this additive-manufacturing process are highlighted below.12,13 High speed: parts are produced easily within hours since special tools are not required. Complex geometry: this technology allows designs of both internal and external features. High quality: it creates parts with high accuracy and detailed resolution. It provides parts with better mechanical strength. In spite of all these advantages, some of the limiting factors are as follows: the process is expensive and power intensive, the surface needs to be polished and the metal support structure removed, and the thermal post-processing is time consuming.14,15 2.2 Electron-beam melting Figure 2 shows a schematic presentation of an electron-beam process. Parts are developed by adding material layer by layer, following the path described by the 3D CAD data program.16 Electron beam is generated within an electron-beam gun; the tungsten filament is heated to an extremely high temperature so that it releases electrons; the electrons are accelerated with an electric field and then focused with electromagnetic coils.17,18 The electron beam melts each layer of the S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING 710 Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 1: Schematic diagram of the direct metal laser sintering system14 Figure 2: Schematic drawing of an electron-beam additive-manu- facturing machine18 metal powder to the desired geometry in a vacuum; the process in the vacuum eliminates impurities and yields high-strength properties of the material under processing.16 The materials used for electron-beam melting are mainly steel and titanium alloy. The following advantages are characteristic of this technology:17 • an increased efficiency of the raw-material use, • the vacuum eliminates impurities and provides a good thermal environment for a freeform fabrication, • it significantly reduces the amount of finishing operations, • freedom in design – design for functionality, • processing of high-melting-point and highly reactive materials, • decreased lead times for design and fabrication, • a high degree of component customization. Some of the disadvantages of the electron-beam technology include the following:19 • It requires a vacuum, another system to the machine, which must be maintained, leading to a high cost. • It produces X-rays while in operation, which is hazardous. 2.3 Selective laser sintering Selective laser sintering is also a powder-bed-based technology that involves the spreading of powder and the subsequent compaction. A schematic diagram of the selective laser sintering is shown in Figure 3. The set-up is made up of a laser, an automatic powder-layering apparatus, and a computer system for process control.20 Selective laser sintering uses a substrate for the part fabrication, which is fixed onto the building platform and leveled. The sealed building chamber containing oxygen has a reduced amount of oxygen due to the protective inert gas such as argon, which is fed into the chamber. A thin layer of a loose powder is deposited on the substrate by the layering mechanism. The laser beam scans the powder-bed surface through the CAD program forming layers of the material to be produced. The powder spreading and laser-treatment process are repeated and the parts are built layer by layer until completion.21–23 This technology allows a variety of materials, the most common are: wax, paraffin, polymer-metal powders or various types of steel alloys, polymers, nylon and carbo- nates.24 The advantages of using selective laser sintering include: • parts can be created out of a wide selection of materials, • complex geometry is made easy as long as the non-sintered powder can be removed easily, • when printing overhanging, unsupported structures, supports are not needed because the unused powder provides the necessary support. Disadvantages of selective laser sintering are: • the equipment is very expensive, • the most common problem of this technology is that the fabricated parts are porous and the surface could be rough, depending on the materials used, • thermal distortion occurs on polymer parts and can cause shrinking and warping of the fabricated parts. 2.4 Selective heat sintering Figure 4 shows a schematic presentation of the selective-heat-sintering process, which operates by selec- tively fusing a thin layer of polymer powder through a thermal print-head assembly. This assembly operates bi-directionally and incorporates thermal print heads at point A as shown in the diagram. The part labelled as B is the powder-deposition mechanism, C is the layer heater, D is the chamber for the material built up in an internal build volume, E is the floor that is a vertically movable building platform; at F, fresh powder is supplied via scoops to the powder-deposition mechanism from the powder containers.25 The heated building platform builds up the parts, using the layer technology, with the roller spreading across the plastic powders in layers. The thermal print head of the apparatus forms a part in its full cross-section with the already laid powder. The heat at the building platform sinters the top layer of the powder; once completed, the process is repeated until a complete 3D object is produced. The complex geometries pro- S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 711 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 4: Schematic presentation of selective heat sintering25 Figure 3: Schematic presentation of a selective-laser-sintering apparatus21 the choice of the materials that could be used on an additive-manufacturing machine at that time.2–4 There are a number of industries benefiting from this revolutionary manufacturing technology,5 such as the aerospace, auto- mobile and biomedical industries.6–9 In this study, some of these additive-manufacturing technologies are re- viewed with the aim of throwing light on these techno- logies and describing their systems of operation as well as their advantages, limitations and areas of applications. The rest of the paper is organized as follows: section 2 of the paper presents the powder-bed-based process. The extrusion-based process and sheet-lamination pro- cess are presented in sections 3 and 4, respectively. Di- rected-energy-deposition process is presented in section 5. Material development in additive manufacturing is presented in section 6, while conclusions are presented in section 7. 2 POWDER-BED-BASED PROCESS This class of additive manufacturing technology uses the energy from an electron beam or laser beam to fuse or melt powder materials for the manufacturing of parts.10,11 The additive-manufacturing technologies in this class include direct metal laser sintering (DMLS), electron-beam melting (EBM), selective heat sintering (SHS), selective laser melting (SLM) and selective laser sintering (SLS). Each of these technologies is described in this section. 2.1 Direct metal laser sintering (DMLS) A schematic diagram of direct laser sintering is shown in Figure 1 below. This process manufactures parts by compacting a metal powdered material with a power source such as laser, without melting but by binding the material together to create a solid structure defined by the 3D CAD model of the part.12 The working principle of this technology makes it capable of design- ing and producing complex geometry of both internal and external intricacies. In this process, to avoid the collision of the recoater blade when moving, the building and dispenser platforms are lowered by the layer thickness. The powder metal needed to create a layer of the material is supplied by the dispenser platform after it has been ensured that the recoater stands in the right position. The spreading of the metal powder from the dispenser to the building platform is done by the re- coater, moving from the right to the left position, and the excess metal powder falls into the excess-powder collector. The scan head moves the laser beam through the two-dimensional (2D) cross-section, which is switched on and off during the exposure of designated areas. The powdered metals are generated, cured and sintered by the solidified areas through the absorption of energy.13 This process continues to create layer after layer until the part is completed. Using this process, parts have been successfully made from the materials such as aluminium alloys – AlSi10Mg, stainless steel, titanium alloys and cobalt chrome superalloy.13 Some of the advantages of this additive-manufacturing process are highlighted below.12,13 High speed: parts are produced easily within hours since special tools are not required. Complex geometry: this technology allows designs of both internal and external features. High quality: it creates parts with high accuracy and detailed resolution. It provides parts with better mechanical strength. In spite of all these advantages, some of the limiting factors are as follows: the process is expensive and power intensive, the surface needs to be polished and the metal support structure removed, and the thermal post-processing is time consuming.14,15 2.2 Electron-beam melting Figure 2 shows a schematic presentation of an electron-beam process. Parts are developed by adding material layer by layer, following the path described by the 3D CAD data program.16 Electron beam is generated within an electron-beam gun; the tungsten filament is heated to an extremely high temperature so that it releases electrons; the electrons are accelerated with an electric field and then focused with electromagnetic coils.17,18 The electron beam melts each layer of the S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING 710 Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 1: Schematic diagram of the direct metal laser sintering system14 Figure 2: Schematic drawing of an electron-beam additive-manu- facturing machine18 metal powder to the desired geometry in a vacuum; the process in the vacuum eliminates impurities and yields high-strength properties of the material under processing.16 The materials used for electron-beam melting are mainly steel and titanium alloy. The following advantages are characteristic of this technology:17 • an increased efficiency of the raw-material use, • the vacuum eliminates impurities and provides a good thermal environment for a freeform fabrication, • it significantly reduces the amount of finishing operations, • freedom in design – design for functionality, • processing of high-melting-point and highly reactive materials, • decreased lead times for design and fabrication, • a high degree of component customization. Some of the disadvantages of the electron-beam technology include the following:19 • It requires a vacuum, another system to the machine, which must be maintained, leading to a high cost. • It produces X-rays while in operation, which is hazardous. 2.3 Selective laser sintering Selective laser sintering is also a powder-bed-based technology that involves the spreading of powder and the subsequent compaction. A schematic diagram of the selective laser sintering is shown in Figure 3. The set-up is made up of a laser, an automatic powder-layering apparatus, and a computer system for process control.20 Selective laser sintering uses a substrate for the part fabrication, which is fixed onto the building platform and leveled. The sealed building chamber containing oxygen has a reduced amount of oxygen due to the protective inert gas such as argon, which is fed into the chamber. A thin layer of a loose powder is deposited on the substrate by the layering mechanism. The laser beam scans the powder-bed surface through the CAD program forming layers of the material to be produced. The powder spreading and laser-treatment process are repeated and the parts are built layer by layer until completion.21–23 This technology allows a variety of materials, the most common are: wax, paraffin, polymer-metal powders or various types of steel alloys, polymers, nylon and carbo- nates.24 The advantages of using selective laser sintering include: • parts can be created out of a wide selection of materials, • complex geometry is made easy as long as the non-sintered powder can be removed easily, • when printing overhanging, unsupported structures, supports are not needed because the unused powder provides the necessary support. Disadvantages of selective laser sintering are: • the equipment is very expensive, • the most common problem of this technology is that the fabricated parts are porous and the surface could be rough, depending on the materials used, • thermal distortion occurs on polymer parts and can cause shrinking and warping of the fabricated parts. 2.4 Selective heat sintering Figure 4 shows a schematic presentation of the selective-heat-sintering process, which operates by selec- tively fusing a thin layer of polymer powder through a thermal print-head assembly. This assembly operates bi-directionally and incorporates thermal print heads at point A as shown in the diagram. The part labelled as B is the powder-deposition mechanism, C is the layer heater, D is the chamber for the material built up in an internal build volume, E is the floor that is a vertically movable building platform; at F, fresh powder is supplied via scoops to the powder-deposition mechanism from the powder containers.25 The heated building platform builds up the parts, using the layer technology, with the roller spreading across the plastic powders in layers. The thermal print head of the apparatus forms a part in its full cross-section with the already laid powder. The heat at the building platform sinters the top layer of the powder; once completed, the process is repeated until a complete 3D object is produced. The complex geometries pro- S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 711 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 4: Schematic presentation of selective heat sintering25 Figure 3: Schematic presentation of a selective-laser-sintering apparatus21 duced are supported by the excess material surrounding the object.26 This technology uses thermoplastic powder as its working material. The benefits of this technology are as follows: it allows inexpensive manufacturing and a good concept evaluation; the thermal print heads are less expensive and the overall cost is affordable. 2.5 Selective laser melting Selective laser melting is similar to the selective- laser-sintering process, with the main difference being between full melting and fusing in selective laser sinter- ing. The process begins with designing a 3D CAD model; the model is broken down into parts to form a number of finite layers. For each layer, the laser scan path is calculated, defining both the boundary contour and the form of the fill sequence called the raster pattern. Layers are formed by spreading powder on the build platform and then the laser melts the powder with the scanning laser beam; the melted powder solidifies to form a 3D solid component.27 A schematic diagram of the process is shown in Figure 5.28 Different materials used with this technology include steel, titanium, aluminium, cobalt-chromium and nickel alloys. The benefits of using the selective-laser-melting technology are as follows: • it minimizes material wastage and saves costs, • improved production-development cycle, • it allows for complex geometry to be made, • ideal process for a low-volume production, • improved buy-to-fly ratio, • functionally graded parts can be produced, • it allows for fully customized parts that suit indi- viduals. The disadvantages of selective laser melting are as follows: • it is an expensive and a very slow process, • tolerances and surface finishes are limited because of the sticking of the unused powder on the surface of the produced part. 3 EXTRUSION-BASED SYSTEM An extrusion-based system uses material extrusion where the material is heated to the molten state and then extruded through a nozzle to form parts layer by layer. The most common technology is fused deposition modeling (FDM). A schematic diagram of the process is shown in Figure 6. Fused deposition modeling produces components by depositing extruded material in layers through a nozzle. Fused deposition modeling uses a 3D CAD program to prepare and build a model; the material is heated to the molten state and forced through the nozzle under pressure to build up a component.29 The advantages of fused deposition modeling are as follows: • the technology supports complex geometries and cavities, • the technology produces high-grade parts, • it is simple to use. The limitations of fused deposition modeling are: • limitation on the materials that can be used, • limited size, • the cost of the actual machine. This technology uses different materials for compo- nent fabrication, including acrylonitrile butadiene styrene (ABS), polylactic acid (PLA), polycarbonate (PC), polyamide (PA), polystyrene (PS), lignin, rubber, nylon and others.29 4 SHEET-LAMINATION PROCESS This is an additive-manufacturing technology, which mainly uses metal sheets or paper for the manufacturing process. This technology is characterized by a low cost of production, high strength of the model, possibilities to S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING 712 Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: Presentation of an extrusion-based system29Figure 5: Schematic presentation of selective laser melting28 control the range of speed of the process to create the outcome desired. Apart from the benefits achieved when using this technology, there are also limitations: the pro- perties of the outcome product depends on the material used and the process of fusing the material together needs improved techniques.30 There are two variants of this process, namely: ultra- sonic additive manufacturing (UAM) and laminated object manufacturing (LOM). Ultrasonic additive manu- facturing (UAM) uses metal as its working material and employs the layer-by-layer techniques, whereby the layers are welded together ultrasonically. This process uses metals like aluminum, copper, stainless steel and titanium alloy. Laminated object manufacturing (LOM) incorporates the layer-by-layer lamination technique, using sheets of paper as the working material and CO2 laser for cutting a sheet that has been joined together by an adhesive to form layers of CAD-model parts. The excess material that is not part of the component being built is sliced into cubes, using a cross-hutch cutting operation.31 5 DIRECTED ENERGY DEPOSITION Directed energy deposition is a type of additive manufacturing that uses the energy from a laser or elec- tron beam to create a melt pool on the surface of the substrate, where coaxial powder or wire is deposited, following the path described by a CAD profile. An example of this process is laser metal deposition, also known as laser engineered net shaping (LENS), a pro- cess with high material-utilization efficiency and cost efficiency.32–41 The laser, or electron beam, leaves a solid track of the melted powder. A schematic diagram of this process is shown in Figure 7. The advantages of directed energy deposition include the ability to carry out a high-value repair on a part that was cost prohibitive in the past; it can also be used to manufacture parts with functionally graded materials or functionally graded coatings.39–41 The materials that can be used for this tech- nology include metals, alloys, ceramics and composites. Other additive-manufacturing processes are material jetting, binder jetting and the vat photopolymerization process. These types of additive manufacturing use non- metallic based materials and their applications are limited. For further reading about these and other additive-manufacturing processes, the reader can consult references.42–52 A new 3D printing technology was developed and it will help us create metallic, free-standing 3D structures. This technology was developed by the researchers at Harvard University.53 The existing direct ink writing 3D printing technology is combined with laser annealing to ensure accurate and stable free-standing 3D structures. This technology is very useful in the applications where free-standing 3D microstructures are required, e.g., in sensors, electronics and also in biomedical applications. 6 MATERIAL DEVELOPMENT FOR ADDITIVE MANUFACTURING There is constant demand for light materials with improved properties, especially in the automobile and aerospace industries, with the aim of reducing the global warming. The advent of additive manufacturing is the beginning of a new dawn in the manufacturing indus- tries. A number of restrictions were placed on material development in the past because of thermodynamic limitations as well as the difficulties in processing some materials in the bulk form. With additive-manufacturing technologies, new materials are now being developed without any restriction. A number of complex metallic alloys are now being developed and used for the fabri- cation of important light-weight structures with im- pressive properties, using the additive-manufacturing process.54 Complex metallic alloys are intermetallic compounds with unique thermal and transport properties that cannot be found in any metal system.55 Complex metallic alloys are needed in a number of applications because of their excellent properties, but they cannot be used with the conventional manufacturing processes because of their brittleness that makes them difficult to process. These materials have low coeffi- cients of friction, good corrosion resistance, good wear- resistance properties and can now be processed using additive manufacturing. New composite materials with improved properties are introduced and used in additive- manufacturing processes, with the functional-part pro- duction now being commercialized.54 The traditional manufacturing process presents a lot of challenges when processing these materials because of the segregation of constituent materials due to the thermodynamic proper- ties of individual materials, but additive-manufacturing technologies can be used to process these materials without such problems. S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 713 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 7: Schematic diagram of directed energy deposition41 duced are supported by the excess material surrounding the object.26 This technology uses thermoplastic powder as its working material. The benefits of this technology are as follows: it allows inexpensive manufacturing and a good concept evaluation; the thermal print heads are less expensive and the overall cost is affordable. 2.5 Selective laser melting Selective laser melting is similar to the selective- laser-sintering process, with the main difference being between full melting and fusing in selective laser sinter- ing. The process begins with designing a 3D CAD model; the model is broken down into parts to form a number of finite layers. For each layer, the laser scan path is calculated, defining both the boundary contour and the form of the fill sequence called the raster pattern. Layers are formed by spreading powder on the build platform and then the laser melts the powder with the scanning laser beam; the melted powder solidifies to form a 3D solid component.27 A schematic diagram of the process is shown in Figure 5.28 Different materials used with this technology include steel, titanium, aluminium, cobalt-chromium and nickel alloys. The benefits of using the selective-laser-melting technology are as follows: • it minimizes material wastage and saves costs, • improved production-development cycle, • it allows for complex geometry to be made, • ideal process for a low-volume production, • improved buy-to-fly ratio, • functionally graded parts can be produced, • it allows for fully customized parts that suit indi- viduals. The disadvantages of selective laser melting are as follows: • it is an expensive and a very slow process, • tolerances and surface finishes are limited because of the sticking of the unused powder on the surface of the produced part. 3 EXTRUSION-BASED SYSTEM An extrusion-based system uses material extrusion where the material is heated to the molten state and then extruded through a nozzle to form parts layer by layer. The most common technology is fused deposition modeling (FDM). A schematic diagram of the process is shown in Figure 6. Fused deposition modeling produces components by depositing extruded material in layers through a nozzle. Fused deposition modeling uses a 3D CAD program to prepare and build a model; the material is heated to the molten state and forced through the nozzle under pressure to build up a component.29 The advantages of fused deposition modeling are as follows: • the technology supports complex geometries and cavities, • the technology produces high-grade parts, • it is simple to use. The limitations of fused deposition modeling are: • limitation on the materials that can be used, • limited size, • the cost of the actual machine. This technology uses different materials for compo- nent fabrication, including acrylonitrile butadiene styrene (ABS), polylactic acid (PLA), polycarbonate (PC), polyamide (PA), polystyrene (PS), lignin, rubber, nylon and others.29 4 SHEET-LAMINATION PROCESS This is an additive-manufacturing technology, which mainly uses metal sheets or paper for the manufacturing process. This technology is characterized by a low cost of production, high strength of the model, possibilities to S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING 712 Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: Presentation of an extrusion-based system29Figure 5: Schematic presentation of selective laser melting28 control the range of speed of the process to create the outcome desired. Apart from the benefits achieved when using this technology, there are also limitations: the pro- perties of the outcome product depends on the material used and the process of fusing the material together needs improved techniques.30 There are two variants of this process, namely: ultra- sonic additive manufacturing (UAM) and laminated object manufacturing (LOM). Ultrasonic additive manu- facturing (UAM) uses metal as its working material and employs the layer-by-layer techniques, whereby the layers are welded together ultrasonically. This process uses metals like aluminum, copper, stainless steel and titanium alloy. Laminated object manufacturing (LOM) incorporates the layer-by-layer lamination technique, using sheets of paper as the working material and CO2 laser for cutting a sheet that has been joined together by an adhesive to form layers of CAD-model parts. The excess material that is not part of the component being built is sliced into cubes, using a cross-hutch cutting operation.31 5 DIRECTED ENERGY DEPOSITION Directed energy deposition is a type of additive manufacturing that uses the energy from a laser or elec- tron beam to create a melt pool on the surface of the substrate, where coaxial powder or wire is deposited, following the path described by a CAD profile. An example of this process is laser metal deposition, also known as laser engineered net shaping (LENS), a pro- cess with high material-utilization efficiency and cost efficiency.32–41 The laser, or electron beam, leaves a solid track of the melted powder. A schematic diagram of this process is shown in Figure 7. The advantages of directed energy deposition include the ability to carry out a high-value repair on a part that was cost prohibitive in the past; it can also be used to manufacture parts with functionally graded materials or functionally graded coatings.39–41 The materials that can be used for this tech- nology include metals, alloys, ceramics and composites. Other additive-manufacturing processes are material jetting, binder jetting and the vat photopolymerization process. These types of additive manufacturing use non- metallic based materials and their applications are limited. For further reading about these and other additive-manufacturing processes, the reader can consult references.42–52 A new 3D printing technology was developed and it will help us create metallic, free-standing 3D structures. This technology was developed by the researchers at Harvard University.53 The existing direct ink writing 3D printing technology is combined with laser annealing to ensure accurate and stable free-standing 3D structures. This technology is very useful in the applications where free-standing 3D microstructures are required, e.g., in sensors, electronics and also in biomedical applications. 6 MATERIAL DEVELOPMENT FOR ADDITIVE MANUFACTURING There is constant demand for light materials with improved properties, especially in the automobile and aerospace industries, with the aim of reducing the global warming. The advent of additive manufacturing is the beginning of a new dawn in the manufacturing indus- tries. A number of restrictions were placed on material development in the past because of thermodynamic limitations as well as the difficulties in processing some materials in the bulk form. With additive-manufacturing technologies, new materials are now being developed without any restriction. A number of complex metallic alloys are now being developed and used for the fabri- cation of important light-weight structures with im- pressive properties, using the additive-manufacturing process.54 Complex metallic alloys are intermetallic compounds with unique thermal and transport properties that cannot be found in any metal system.55 Complex metallic alloys are needed in a number of applications because of their excellent properties, but they cannot be used with the conventional manufacturing processes because of their brittleness that makes them difficult to process. These materials have low coeffi- cients of friction, good corrosion resistance, good wear- resistance properties and can now be processed using additive manufacturing. New composite materials with improved properties are introduced and used in additive- manufacturing processes, with the functional-part pro- duction now being commercialized.54 The traditional manufacturing process presents a lot of challenges when processing these materials because of the segregation of constituent materials due to the thermodynamic proper- ties of individual materials, but additive-manufacturing technologies can be used to process these materials without such problems. S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 713 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 7: Schematic diagram of directed energy deposition41 A number of new materials with improved properties and performances are now being developed for the additive-manufacturing technologies. These new mate- rials include the high-temperature copper alloy, deve- loped by NASA, with excellent properties such as retention of high strength at elevated temperatures, used in rocket engines. Other materials that present challenges with the conventional manufacturing processes and are now successfully processed using the additive-manu- facturing technologies include titanium aluminide, used in the fabrication of high-speed gas-turbine blades, reactor-grade pure niobium and NiTi. 7 CONCLUSIONS The concept of additive manufacturing has provided solutions to most engineering problems, especially the problems faced by product designers in the past when products needed to be designed and redesigned because of the limitations imposed by the available manu- facturing process. Designers can now freely design parts without having to worry about how they will be made. This research paper presents a review on various additive-manufacturing technologies. The aim of this paper is to provide a clear description of this novel manufacturing process – the additive-manufacturing process. Each of the technologies is described in this paper, including the materials suitable for use in each technology, advantages, disadvantages. Application areas and the new material development relating to the addi- tive-manufacturing technologies are presented. In summary, additive manufacturing is a reliable techno- logy that can be used to manufacture prototypes, tooling and functional parts, as it is less time consuming, allow- ing an easy fabrication of a complex geometry, an easy production of customized and personalized parts, with- out any need for special tools, and low material wastage that reduces the overall cost of the production. Im- provements in the computing power and the reduction in mass-storage costs paved the way for processing the large amounts of data typical of modern 3D CAD models within reasonable time frames. A new 3D printing technology has recently been developed at Harvard University. This new additive- manufacturing technology combines the existing direct ink writing 3D printing technology and the annealing process. The innovation was born out of the necessity for flexible and wearable electronic devices such as sensors and in biomedical applications. The technology uses the 3D printing technology to create complex architectures, while a printed structure is simultaneously annealed in midair. This helps to increase the accuracy of the printed material, also allowing the printed material to be created in free air without the need for any support structure. This technology has opened up possibilities of creating microscopic metallic free-standing 3D structures without an auxiliary support and in a single manufacturing run. 8 REFERENCES 1 J. Scott, N. Gupta, C. Wember, S. Newsom, T. Wohlers, T. Caffrey, Additive manufacturing: status and opportunities, Science and Tech- nology Policy Institute, 2012, https://www.ida.org/stpi/occasional- papers/papers/AM3D_33012_Final.pdf 2 J. P. Kruth, M. C. Leu, T. Nakagawa, Progress in additive manu- facturing and rapid prototyping, CIRP Annals, 47 (1998), 2, doi:10.1016/S0007-8506(07)63240-5 3 T. Wohlers, T. Caffrey, Additive manufacturing: going mainstream, manufacturing engineering, http://advancedmanufacturing.org/ additive-manufacturing-going-mainstream/, 2013 4 F. Liou, K. Slattery, M. Kinsella, J. Newkirk, H. N. Chou, R. Lan- ders, Applications of a hybrid manufacturing process for fabrication of metallic structures, Rapid Prototyping Journal, 13 (2007), 4, doi:10.1108/13552540710776188 5 R. M. Mahamood, E. T. Akinlabi, M. Shukla, S. Pityana, Revolution- ary additive manufacturing: an overview, Lasers in Engineering, 27 (2014), 3–4, 161–178 6 N. Hopkinson, P. Dickens, Rapid prototyping for direct manufacture, Rapid Prototyping Journal, 7 (2001), 4, doi:10.1108/EUM0000- 000005753 7 Y. L. Hou, T. T. Zhao, C. H. Li, Y. C. Ding, The manufacture of rapid tooling by stereo lithography, Advanced Materials Research, 102–104 (2010), 578–582 8 L. Hao, S. Dadbakhsh, Materials and process aspects of selective laser melting of metals and metal matrix composites: a review, Chinese Journal of Lasers, 36 (2009), 12, doi:10.3788/CJL2009- 3612.3192 9 V. Petrovic, J. V. H. Gonzalez, O. J. Ferranda, J. D. Gordillo, J. R. B. Puchades, L. P. Grinan, Additive layered manufacturing: sectors of industrial application shown through case studies, International Journal of Production Research, 49 (2011), 4, doi:10.1080/00207- 540903479786 10 S. F. S. Shirazi, S. Ghanehkhani, M. Mehrali, H. Yarmand, H. S. C. Metselaar, N. A. Kadri, N. A. A. Usman, A review on powder-based additive manufacturing for tissue engineering: selective laser sintering and inkjet 3D printing, Science and Technology of Advance Materials, 16 (2015), 3, doi:10.1088/1468-6996/16/3/033502 11 M. Dickson, Soft strain sensors fabricated through additive manu- facturing, MRS Bulletin, 40 (2015), 6, doi:10.1557/mrs.2015.124 12 A. R. R. Bineli, A. P. G. Peres, L. F. Bernardes, A. L. Jardini, R. Filho, Design of microreactor by integration of reverse engineering and direct metal laser sintering process, Proceedings of the 5th international workshop on hydrogen and fuel cells, Campinas, 2010 13 O. Nyrhila, Characterization of Process Parameter for Direct Metal Laser Sintering, Ph.D. Thesis, Nagoya Institute of Technology, 2005 14 A. Bineli, A. Peres, A. Jardini, R. Filho, Direct metal laser sintering: technology for design and construction of microreactors, Proceed- ings of the 6th Brazilian Conference on Manufacturing Engineering, 2011 15 R. Aulus, R. Maciel, Direct metal laser sintering (DMLS): tech- nology for design and construction of microreactors, Science and Technology of Advance Materials, 7 (2015), 2 16 L. Morgan, L. Ulf, H. Ola, Rapid manufacturing with electron beam melting (EBM) – a manufacturing revolution, 2006, 433–438, https://sffsymposium.engr.utexas.edu/Manuscripts/2003/2003-41-Lar sson.pdf 17 L. Ladani, L. Roy, Mechanical behavior of TI-6AL-4V manufactured by electron beam additive fabrication, Proceedings of the ASME International Manufacturing Science and Engineering Conference, Madison, 2013 18 S. Biamino, A. Penna, U. Ackelid, S. Sabbadini, O. Tassa, P. Fino, M. Pavese, P. Gennaro, C. Badini, Electron beam melting of Ti-48Al-2Cr-2Nb alloy: microstructure and mechanical properties investigation, Intermetallics, 19 (2010), 6, doi:10.1016/j.intermet. 2010.11.017 S. A. 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Pityana, Scanning Velocity Influence on Microstructure, Microhardness and Wear Resistance Performance on Laser Deposited Ti6Al4V/TiC Com- posite, Materials and Design, 50 (2013), 656–666 38 R. M. Mahamood, E. T. Akinlabi, M. Shukla, S. Pityana, Charac- terizing the Effect of Laser Power Density on Microstructure, Microhardness and Surface Finish of Laser Deposited Titanium Alloy, Journal of Manufacturing Science and Engineering, 135 (2013), 6, doi:10.1115/1.4025737 39 J. Scott, N. Gupta, C. Weber, S. Newsome, T. Wohlers, T. Caffrey, Additive Manufacturing: Status and Opportunities, https://cgsr. llnl.gov/content/assets/docs/IDA_AdditiveM3D_33012_Final.pdf 40 R. M. Mahamood, E. T. Akinlabi, Laser metal deposition of func- tionally graded Ti6Al4V/TiC, Materials & Design, 84 (2015), 402–410 41 R. M. Mahamood, E. T. Akinlabi, M. Shukla, S. Pityana, Func- tionally graded material: An overview, Proceedings of the World Congress on Engineering, London, UK, 2012, 1593–1597 42 A. D. 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Nakagawa, Progress in additive manu- facturing and rapid prototyping, CIRP Annals – Manufacturing Technology, 47 (1998) 2, 525–540 49 J. Kietzman, L. Pitt, P. Berthon, Disruptions, decisions, and desti- nations: Enter the age of 3-D printing and additive manufacturing, Business Horizons, 58 (2015) 2, 201–210 50 K. Chan, Metal Additive Manufacturing: A Review, Journal of Ma- terials Engineering and Performance, 31 (2016) 6, 1917–1928 51 R. M. Mahamood, E. T. Akinlabi, Influence of scanning speed intermetallic produced in-situ in laser metal deposited TiC/Ti6Al4V composite, Mater. Tehnol., 51 (2017) 3, 473–478, doi:10.17222/ mit.2016.096 52 Y. Zhai, D. A. Lados, J. L. Lagoy, Additive Manufacturing: Making Imagination the Major Limitation, JOM, 66 (2014) 5, 808–816 53 M. A. Skylar-Scott, S. Gunasekaran, J. A. Lewis, Laser-assisted direct ink writing of planar and 3D metal architectures, PNAS, 113 (2016) 22, doi:10.1073/pnas.1525131113 54 S. Kenzari, D. Bonina, J. M. Dubois, V. Fournée, Complex metallic alloys as new materials for additive manufacturing, Science and Technology of Advanced Materials, 15 (2014) 2, 1–9 55 J-M. Dubois, Properties and applications of quasicrystals and com- plex metallic alloys, Chem. Soc. Rev., 41 (2012), 6760–6777 S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 715 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS A number of new materials with improved properties and performances are now being developed for the additive-manufacturing technologies. These new mate- rials include the high-temperature copper alloy, deve- loped by NASA, with excellent properties such as retention of high strength at elevated temperatures, used in rocket engines. Other materials that present challenges with the conventional manufacturing processes and are now successfully processed using the additive-manu- facturing technologies include titanium aluminide, used in the fabrication of high-speed gas-turbine blades, reactor-grade pure niobium and NiTi. 7 CONCLUSIONS The concept of additive manufacturing has provided solutions to most engineering problems, especially the problems faced by product designers in the past when products needed to be designed and redesigned because of the limitations imposed by the available manu- facturing process. Designers can now freely design parts without having to worry about how they will be made. This research paper presents a review on various additive-manufacturing technologies. The aim of this paper is to provide a clear description of this novel manufacturing process – the additive-manufacturing process. Each of the technologies is described in this paper, including the materials suitable for use in each technology, advantages, disadvantages. Application areas and the new material development relating to the addi- tive-manufacturing technologies are presented. In summary, additive manufacturing is a reliable techno- logy that can be used to manufacture prototypes, tooling and functional parts, as it is less time consuming, allow- ing an easy fabrication of a complex geometry, an easy production of customized and personalized parts, with- out any need for special tools, and low material wastage that reduces the overall cost of the production. Im- provements in the computing power and the reduction in mass-storage costs paved the way for processing the large amounts of data typical of modern 3D CAD models within reasonable time frames. A new 3D printing technology has recently been developed at Harvard University. This new additive- manufacturing technology combines the existing direct ink writing 3D printing technology and the annealing process. The innovation was born out of the necessity for flexible and wearable electronic devices such as sensors and in biomedical applications. The technology uses the 3D printing technology to create complex architectures, while a printed structure is simultaneously annealed in midair. This helps to increase the accuracy of the printed material, also allowing the printed material to be created in free air without the need for any support structure. This technology has opened up possibilities of creating microscopic metallic free-standing 3D structures without an auxiliary support and in a single manufacturing run. 8 REFERENCES 1 J. Scott, N. Gupta, C. Wember, S. Newsom, T. Wohlers, T. Caffrey, Additive manufacturing: status and opportunities, Science and Tech- nology Policy Institute, 2012, https://www.ida.org/stpi/occasional- papers/papers/AM3D_33012_Final.pdf 2 J. P. Kruth, M. C. Leu, T. Nakagawa, Progress in additive manu- facturing and rapid prototyping, CIRP Annals, 47 (1998), 2, doi:10.1016/S0007-8506(07)63240-5 3 T. Wohlers, T. Caffrey, Additive manufacturing: going mainstream, manufacturing engineering, http://advancedmanufacturing.org/ additive-manufacturing-going-mainstream/, 2013 4 F. Liou, K. Slattery, M. Kinsella, J. Newkirk, H. N. Chou, R. Lan- ders, Applications of a hybrid manufacturing process for fabrication of metallic structures, Rapid Prototyping Journal, 13 (2007), 4, doi:10.1108/13552540710776188 5 R. M. Mahamood, E. T. Akinlabi, M. Shukla, S. Pityana, Revolution- ary additive manufacturing: an overview, Lasers in Engineering, 27 (2014), 3–4, 161–178 6 N. Hopkinson, P. Dickens, Rapid prototyping for direct manufacture, Rapid Prototyping Journal, 7 (2001), 4, doi:10.1108/EUM0000- 000005753 7 Y. L. Hou, T. T. Zhao, C. H. Li, Y. C. Ding, The manufacture of rapid tooling by stereo lithography, Advanced Materials Research, 102–104 (2010), 578–582 8 L. Hao, S. Dadbakhsh, Materials and process aspects of selective laser melting of metals and metal matrix composites: a review, Chinese Journal of Lasers, 36 (2009), 12, doi:10.3788/CJL2009- 3612.3192 9 V. Petrovic, J. V. H. Gonzalez, O. J. Ferranda, J. D. Gordillo, J. R. B. Puchades, L. P. Grinan, Additive layered manufacturing: sectors of industrial application shown through case studies, International Journal of Production Research, 49 (2011), 4, doi:10.1080/00207- 540903479786 10 S. F. S. Shirazi, S. Ghanehkhani, M. Mehrali, H. Yarmand, H. S. C. Metselaar, N. A. Kadri, N. A. A. Usman, A review on powder-based additive manufacturing for tissue engineering: selective laser sintering and inkjet 3D printing, Science and Technology of Advance Materials, 16 (2015), 3, doi:10.1088/1468-6996/16/3/033502 11 M. Dickson, Soft strain sensors fabricated through additive manu- facturing, MRS Bulletin, 40 (2015), 6, doi:10.1557/mrs.2015.124 12 A. R. R. Bineli, A. P. G. Peres, L. F. Bernardes, A. L. Jardini, R. Filho, Design of microreactor by integration of reverse engineering and direct metal laser sintering process, Proceedings of the 5th international workshop on hydrogen and fuel cells, Campinas, 2010 13 O. Nyrhila, Characterization of Process Parameter for Direct Metal Laser Sintering, Ph.D. Thesis, Nagoya Institute of Technology, 2005 14 A. Bineli, A. Peres, A. Jardini, R. Filho, Direct metal laser sintering: technology for design and construction of microreactors, Proceed- ings of the 6th Brazilian Conference on Manufacturing Engineering, 2011 15 R. Aulus, R. Maciel, Direct metal laser sintering (DMLS): tech- nology for design and construction of microreactors, Science and Technology of Advance Materials, 7 (2015), 2 16 L. Morgan, L. Ulf, H. Ola, Rapid manufacturing with electron beam melting (EBM) – a manufacturing revolution, 2006, 433–438, https://sffsymposium.engr.utexas.edu/Manuscripts/2003/2003-41-Lar sson.pdf 17 L. Ladani, L. Roy, Mechanical behavior of TI-6AL-4V manufactured by electron beam additive fabrication, Proceedings of the ASME International Manufacturing Science and Engineering Conference, Madison, 2013 18 S. Biamino, A. Penna, U. Ackelid, S. Sabbadini, O. Tassa, P. Fino, M. Pavese, P. Gennaro, C. Badini, Electron beam melting of Ti-48Al-2Cr-2Nb alloy: microstructure and mechanical properties investigation, Intermetallics, 19 (2010), 6, doi:10.1016/j.intermet. 2010.11.017 S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING 714 Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 19 A. Neira, Thermal Modeling and Simulation of Electron Beam Melting for Rapid Prototyping on Ti6Al4V Alloys, Ph.D. Disser- tation, North Carolina State University, Raleigh, NC, 2012 20 D. Bouk, H. Kaal, Selective Laser Sintering of Metals, Proceedings of ASME, American Society of Mechanical Engineers, New York, 2015 21 H. J. Niu, I. T. H. Chang, Selective laser sintering of gas atomized M2 high speed steel powder, Journal of Materials Science, 35 (2000), 1, doi:10.1023/A:1004720011671 22 M. Peterson, A. Kalz, Characterization and comparison of materials of selective laser sintering, Journal of Materials Processing Technology, 21 (2013) 3, 21–22 23 C. Telenko, C. C. Seepersad, Assessing energy requirements and material flows of selective laser sintering of nylon parts, Proceedings of the Solid Freeform Fabrication Symposium, Austin, USA, 2010, 289–297 24 S. Caroly, Design rules for selective laser sintering, https://www.me.utexas.edu/~ppmdlab/files/designers.guide.sls.pdf 25 M. Baumers, C. Tuck, R. Hague, Selective heat sintering versus laser sintering: comparison of deposition rate, process energy consump- tion and cost performance, https://sffsymposium.engr.utexas.edu/ sites/default/files/2015/2015-9-Baumers.pdf 26 R. M. Mahamood, E. T. Akinlabi, M. Shukla, S. Pityana, Material efficiency of laser metal deposited TI6AL4V: effect of laser power, Engineering Letters, 21 (2013), 1, http://www.engineeringletters. com/issues_v21/issue_1/EL_21_1_03.pdf 27 A. Diatlov, D. Buchbinder, W. Meiners. K. Wissenbach, Towards surface topography: Quantification of Selective Laser Melting (SLM) built parts, Review of selected measurement methods and ongoing report on development of measurement specifications, 2011 28 Additive Manufacturing Processes, www.slidehot.com/9673714743 29 A. K. Sood, R. K. Ohdar, S. S. Mahapatra, Parametric appraisal of mechanical property of fused deposition modeling processed parts, Materials and Design, 31 (2010), 201–246 30 L. Gibson, D. W. Rosen, B. Stucker, Sheet lamination processes: additive manufacturing technologies, Journal of Materials and Manu- facturing, 23 (2012) 3, 223–252 31 C. M. Cheah, Rapid sheet metal manufacturing, International Journal of Advanced Manufacturing Technology, 19 (2002), 510–515 32 R. M. Mahamood, E. T. Akinlabi, Effect of processing parameters on wear resistance property of laser material deposited titanium-alloy composite, Journal of Optoelectronics and Advanced Materials, 17 (2015) 9–10, 1348–1360 33 R. M. Mahamood, E. T. Akinlabi, Process parameters optimization for material deposition efficiency in laser metal deposited titanium alloy, Lasers in Manufacturing and Materials Processing, 3 (2016), 1, doi:10.1007/s40516-015-0020-5 34 R. M. Mahamood, E. T. Akinlabi, Effect of laser power and powder flow rate on the wear resistance behaviour of laser metal deposited TiC/Ti6Al4V composites, Materials Today: Proceedings, 2 (2015) 4–5, 2679–2686 35 R. M. Mahamood, E. T. Akinlabi, M. Shukla, S. Pityana, Charac- terization of Laser Deposited Ti6A4V/TiC Composite, Lasers in Engineering, 29 (2014) 3–4, 197–213 36 R. M. Mahamood, E. T. Akinlabi, S. A. Akinlabi, Laser power and scanning speed influence on the mechanical property of laser metal deposited titanium-alloy, Lasers in Manufacturing and Materials Processing, 2 (2014) 1, 43–55 37 R. M. Mahamood, E. T. Akinlabi, M. Shukla, S. Pityana, Scanning Velocity Influence on Microstructure, Microhardness and Wear Resistance Performance on Laser Deposited Ti6Al4V/TiC Com- posite, Materials and Design, 50 (2013), 656–666 38 R. M. Mahamood, E. T. Akinlabi, M. Shukla, S. Pityana, Charac- terizing the Effect of Laser Power Density on Microstructure, Microhardness and Surface Finish of Laser Deposited Titanium Alloy, Journal of Manufacturing Science and Engineering, 135 (2013), 6, doi:10.1115/1.4025737 39 J. Scott, N. Gupta, C. Weber, S. Newsome, T. Wohlers, T. Caffrey, Additive Manufacturing: Status and Opportunities, https://cgsr. llnl.gov/content/assets/docs/IDA_AdditiveM3D_33012_Final.pdf 40 R. M. Mahamood, E. T. Akinlabi, Laser metal deposition of func- tionally graded Ti6Al4V/TiC, Materials & Design, 84 (2015), 402–410 41 R. M. Mahamood, E. T. Akinlabi, M. Shukla, S. Pityana, Func- tionally graded material: An overview, Proceedings of the World Congress on Engineering, London, UK, 2012, 1593–1597 42 A. D. Halvorsen, P. Vaidya, M. Robinson, D. L. Schulz, Transform- ing a laser micromachine into a direct-write tool for electronic materials, Journal of Micro-Electronics and Electronic Packaging, 5 (2008), 116–121 43 K. K. B. Hon, L. Li, I. M. Hutchings, Direct writing technology – advances and developments, CIRP Annals, 57 (2008) 2, 601–620 44 A. Lutfurakhmanov, G. K. Loken, D. L. Schulz, I. S. Akhatov, Capillary-based liquid microdroplet deposition, Applied Physics Letters, 97 (2010) 12, 1–3 45 D. B. Chrisey, A. Pique, R. Modi, H. D. Wu, R. C. Y. Auyeung, H. D. Young, Direct writing of conformal microscopic electronic device by MAPLE DW, Applied Surface Science, 168 (2000) 1–4, 345–352 46 M. Burns, Automated Fabrication: Improving Productivity in Manu- facturing, Prentice Hall, Eaglewood Cliffs, NJ, 1993 47 Wohlers Report, Additive manufacturing and 3D printing state of the industry, https://wohlersassociates.com/state-of-the-industry-re- ports.html 48 J. P. Kruth, M. C. Leu, T. Nakagawa, Progress in additive manu- facturing and rapid prototyping, CIRP Annals – Manufacturing Technology, 47 (1998) 2, 525–540 49 J. Kietzman, L. Pitt, P. Berthon, Disruptions, decisions, and desti- nations: Enter the age of 3-D printing and additive manufacturing, Business Horizons, 58 (2015) 2, 201–210 50 K. Chan, Metal Additive Manufacturing: A Review, Journal of Ma- terials Engineering and Performance, 31 (2016) 6, 1917–1928 51 R. M. Mahamood, E. T. Akinlabi, Influence of scanning speed intermetallic produced in-situ in laser metal deposited TiC/Ti6Al4V composite, Mater. Tehnol., 51 (2017) 3, 473–478, doi:10.17222/ mit.2016.096 52 Y. Zhai, D. A. Lados, J. L. Lagoy, Additive Manufacturing: Making Imagination the Major Limitation, JOM, 66 (2014) 5, 808–816 53 M. A. Skylar-Scott, S. Gunasekaran, J. A. Lewis, Laser-assisted direct ink writing of planar and 3D metal architectures, PNAS, 113 (2016) 22, doi:10.1073/pnas.1525131113 54 S. Kenzari, D. Bonina, J. M. Dubois, V. Fournée, Complex metallic alloys as new materials for additive manufacturing, Science and Technology of Advanced Materials, 15 (2014) 2, 1–9 55 J-M. Dubois, Properties and applications of quasicrystals and com- plex metallic alloys, Chem. Soc. Rev., 41 (2012), 6760–6777 S. A. ADEKANYE et al.: ADDITIVE MANUFACTURING: THE FUTURE OF MANUFACTURING Materiali in tehnologije / Materials and technology 51 (2017) 5, 709–715 715 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS E. NARED: POMEMBNA OBLETNICA REVIJE MATERIALI IN TEHNOLOGIJE: PETDESET LET IZHAJANJA ... 717–720 POMEMBNA OBLETNICA REVIJE MATERIALI IN TEHNOLOGIJE: PETDESET LET IZHAJANJA ZNANSTVENE PERIODI^NE PUBLIKACIJE AN IMPORTANT ANNIVERSARY OF THE MATERIALS AND TECHNOLOGY JOURNAL: FIFTY YEARS OF PUBLICATION Erika Nared In{titut za kovinske materiale in tehnologije, Lepi pot 11, 1000 Ljubljana erika.nared@imt.si Prejem rokopisa – received: 2017-08-11; sprejem za objavo – accepted for publication: 2017-09-04 doi:10.17222/mit.2017.136 Revija Materiali in tehnologije (MIT), ki jo izdaja In{titut za kovinske materiale in tehnologije Ljubljana, v leto{njem letu praznuje petdeseto obletnico neprekinjenega izhajanja v tiskani obliki. Tudi v elektronski obliki je revija MIT, kot jo skraj{ano imenujemo, na voljo `e kar precej let in sicer od leta 1996. V ~lanku je predstavljen pregled izhajanja serijske publikacije Materiali in tehnologije (ISSN 1580-2949) v zadnjih desetih letih; od leta 2007 pa vse do danes. Opisane so nekatere spremembe v reviji v tem obdobju, njena dinamika izhajanja ter vizija njenega izhajanja v naslednjih letih. Klju~ne besede: znanstvena periodi~na publikacija, zgodovinski pregled, obletnica izhajanja, uredni{ka politika The Materials and Technology journal (MIT), published by the Institute of Metals and Technology Ljubljana, is celebrating 50 years of its printed version, with the online archive stretching back to 1996. This article provides an overwiev of the journal (ISSN 1580-2949), covering the past 10 years. Some of the new changes to the journal are described, together with its vision for the future. Keywords: scientific periodical publication, historical overview, anniversary of publication, editorial policy 1 PREDSTAVITEV REVIJE IN NJEN RAZVOJ Revija Materiali in tehnologije ima zelo dolgo zgo- dovino izhajanja, saj izhaja `e od leta 1967. @elezarski zbornik (ISSN 0372-8633), kakor se je imenovala pred petdesetimi leti, ko je za~ela izhajati, je bil strokovno glasilo Slovenskih `elezarn in Metalur{kega in{tituta Ljubljana. ^etrtletnik je takrat tehni~no urejal Edo @agar in sicer do leta 1980; nasledila sta ga Darko Brada{kja do leta 1987 in Jana Jamar do leta 1991. Glavni in odgo- vorni urednik @elezarskega zbornika pa je bil vse do konca njegovega izhajanja pod tem imenom mag. Jo`a Arh (Slika 1), torej kar 24 let (1967–1991). @elezarski zbornik je objavljal prispevke s podro~ja kovinskih in deloma nekovinskih materialov.1 Po dveh desetletjih je revija zaradi vsebinskih sprememb dobila novo ime: Kovine zlitine tehnologije (KZT) (ISSN 1318-0010). Preimenovala se je zaradi vse- binske raz{iritve, saj je po novem vsebovala prispevke tako s podro~ja kovinskih materialov kot tudi s podro~ja anorganskih materialov, polimerov in materialov za namen vakuumske tehnike. Skladno z menjavo naslova je dobila tudi novo {tevilko ISSN.1 Glavni in odgovorni urednik je bil do leta 1994 {e vedno mag. Jo`e Arh, nasledil ga je mag. Ale{ Lagoja, ki je oblikoval ured- ni{ko politiko revije v naslednjih 4 letih (Slika 1). MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Materiali in tehnologije / Materials and technology 51 (2017) 5, 717–720 717 UDK 050:669.1:6 (497) ISSN 1580-2949 Review article/Pregledni ~lanek MTAEC9, 51(5)717(2017) Slika 1 (od leve proti desni): Jo`e Arh – glavni urednik revije od 1967 do 1991, Ale{ Lagoja – glavni urednik revije od 1991 do 1995, Franc Vodopivec – glavni urednik revije od 1995 do 2011, Matja` Torkar – glavni urednik revije od 2012 do julija 2016 – Paul McGuiness, sedanji glavni urednik revije E. NARED: POMEMBNA OBLETNICA REVIJE MATERIALI IN TEHNOLOGIJE: PETDESET LET IZHAJANJA ... 717–720 POMEMBNA OBLETNICA REVIJE MATERIALI IN TEHNOLOGIJE: PETDESET LET IZHAJANJA ZNANSTVENE PERIODI^NE PUBLIKACIJE AN IMPORTANT ANNIVERSARY OF THE MATERIALS AND TECHNOLOGY JOURNAL: FIFTY YEARS OF PUBLICATION Erika Nared In{titut za kovinske materiale in tehnologije, Lepi pot 11, 1000 Ljubljana erika.nared@imt.si Prejem rokopisa – received: 2017-08-11; sprejem za objavo – accepted for publication: 2017-09-04 doi:10.17222/mit.2017.136 Revija Materiali in tehnologije (MIT), ki jo izdaja In{titut za kovinske materiale in tehnologije Ljubljana, v leto{njem letu praznuje petdeseto obletnico neprekinjenega izhajanja v tiskani obliki. Tudi v elektronski obliki je revija MIT, kot jo skraj{ano imenujemo, na voljo `e kar precej let in sicer od leta 1996. V ~lanku je predstavljen pregled izhajanja serijske publikacije Materiali in tehnologije (ISSN 1580-2949) v zadnjih desetih letih; od leta 2007 pa vse do danes. Opisane so nekatere spremembe v reviji v tem obdobju, njena dinamika izhajanja ter vizija njenega izhajanja v naslednjih letih. Klju~ne besede: znanstvena periodi~na publikacija, zgodovinski pregled, obletnica izhajanja, uredni{ka politika The Materials and Technology journal (MIT), published by the Institute of Metals and Technology Ljubljana, is celebrating 50 years of its printed version, with the online archive stretching back to 1996. This article provides an overwiev of the journal (ISSN 1580-2949), covering the past 10 years. Some of the new changes to the journal are described, together with its vision for the future. Keywords: scientific periodical publication, historical overview, anniversary of publication, editorial policy 1 PREDSTAVITEV REVIJE IN NJEN RAZVOJ Revija Materiali in tehnologije ima zelo dolgo zgo- dovino izhajanja, saj izhaja `e od leta 1967. @elezarski zbornik (ISSN 0372-8633), kakor se je imenovala pred petdesetimi leti, ko je za~ela izhajati, je bil strokovno glasilo Slovenskih `elezarn in Metalur{kega in{tituta Ljubljana. ^etrtletnik je takrat tehni~no urejal Edo @agar in sicer do leta 1980; nasledila sta ga Darko Brada{kja do leta 1987 in Jana Jamar do leta 1991. Glavni in odgo- vorni urednik @elezarskega zbornika pa je bil vse do konca njegovega izhajanja pod tem imenom mag. Jo`a Arh (Slika 1), torej kar 24 let (1967–1991). @elezarski zbornik je objavljal prispevke s podro~ja kovinskih in deloma nekovinskih materialov.1 Po dveh desetletjih je revija zaradi vsebinskih sprememb dobila novo ime: Kovine zlitine tehnologije (KZT) (ISSN 1318-0010). Preimenovala se je zaradi vse- binske raz{iritve, saj je po novem vsebovala prispevke tako s podro~ja kovinskih materialov kot tudi s podro~ja anorganskih materialov, polimerov in materialov za namen vakuumske tehnike. Skladno z menjavo naslova je dobila tudi novo {tevilko ISSN.1 Glavni in odgovorni urednik je bil do leta 1994 {e vedno mag. Jo`e Arh, nasledil ga je mag. Ale{ Lagoja, ki je oblikoval ured- ni{ko politiko revije v naslednjih 4 letih (Slika 1). MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Materiali in tehnologije / Materials and technology 51 (2017) 5, 717–720 717 UDK 050:669.1:6 (497) ISSN 1580-2949 Review article/Pregledni ~lanek MTAEC9, 51(5)717(2017) Slika 1 (od leve proti desni): Jo`e Arh – glavni urednik revije od 1967 do 1991, Ale{ Lagoja – glavni urednik revije od 1991 do 1995, Franc Vodopivec – glavni urednik revije od 1995 do 2011, Matja` Torkar – glavni urednik revije od 2012 do julija 2016 – Paul McGuiness, sedanji glavni urednik revije Do konca izhajanja revije z imenom Kovine zlitine tehnologije (1999), je uredni{ko politiko vodil prof. dr. Franc Vodopivec (Slika 1), ki je bil na mestu urednika 16 let in tako po letih urednikovanja na drugem mestu. Tehni~no uredni{tvo revije Kovine zlitine tehnologije je vodila Jana Jamar. Izdajatelj revije je bil In{titut za kovinske materiale in tehnologije Ljubljana skupaj s soizdajatelji: ACRONI Jesenice, IMPOL Slovenska Bistrica, Kemijski in{titut Ljubljana, Koncern Slovenske `elezarne, Metal Ravne, Talum Kidri~evo, Fakulteta za strojni{tvo Ljubljana, Institut Jo`ef Stefan in Slovensko dru{tvo za tribologijo.1 Revija z novim imenom Materiali in tehnologije (ISSN 1580-2949) je nasledila obe svoji predhodnici na prelomu tiso~letja in z letom 2000 za~ela izhajati kot znanstvena serijska publikacija, ki objavlja izvirne in pregledne znanstvene ~lanke ter strokovne ~lanke, ki obravnavajo teoreti~na in prakti~na vpra{anja naravo- slovnih ved in tehnologije na podro~jih kovinskih in anorganskih materialov, polimerov, vakuumske tehnike, kompozitnih in gradbenih materialov ter nanomate- rialov.2 Njen namen je bil raz{iriti njeno interesno podro~je, da ne bi delovala le na podro~ju kovin. Odgovorni urednik revije je bil vse do leta 2011 prof. dr. Franc Vodopivec. Do istega leta je tehni~no urednikovanje vodila Jana Jamar. Z letom 2012 so se v uredni{kem odboru zgodile spremembe predvsem na kadrovskem podro~ju. Uredni{ko delo je vse do svoje upokojitve v letu 2016 vodil dr. Matja` Torkar (Slika 1), tehni~na urednica v tem ~asu je bila dr. Danijela A. Skobir Balanti~. Leta 2016 je mesto odgovornega in glavnega urednika revije prevzel dr. Paul John McGuiness (Sli- ka 1), mesto pomo~nika glavnega in odgovornega ured- nika doc. dr. Matja` Godec, ~astna glavna urednika sta prof. dr. Franc Vodopivec in dr. Matja` Torkar. Tehni~no uredni{tvo vodi Erika Nared, ki na In{titutu za kovinske materiale in tehnologije vodi tudi specialno knji`nico. Souredniki revije Materiali in tehnologije so: Igor Beli~, Jaka Burja, Aleksandra Kocijan, Djordje Man- drino, Irena Paulin, Danijela A. Skobir Balanti~, Darja Steiner Petrovi~, Bojan Podgornik iz In{tituta za kovin- ske materiale in tehnologije (IMT), Bo{tjan Markoli iz Naravoslovnotehni{ke fakultete (NTF), Sre~o [kapin in Rok Zaplotnik iz In{tituta Jo`ef Stefan (IJS) ter Ema @agar iz Kemijskega in{tituta (KI).3 Osve`en je tudi uredni{ki odbor in zasedba ~lanov mednarodnih pri- dru`enih ~lanov uredni{kega odbora ter izdajateljskega sveta. ^lanki revije MIT so indeksirani v bazah podatkov, kot so: Science Citation Index Expanded, Materials Science Citation Index in Journal Citation Reports (Science Edition). Po slednji, bazi podatkov JCR, ima revija trenutno faktor vpliva 0.436. ^lanki, objavljeni v reviji, so indeksirani tudi v mednarodnih in doma~ih sekundarnih virih/bazah: DOAJ (Directory of Open Access Journals), Google Scholar, SCOPUS, WoS, COBIB in dLib.si (Digitalna knji`nica Slovenije). Revija si, tako kot `e ve~ino ~asa svojega izhajanja, {e vedno prizadeva za vi{jo citiranost, saj bi bila posledi~no vsekakor bolj zanimiva, tako za slovenske raziskovalce kot tudi tuje, in bi pritegnila uveljavljene doma~e in tuje raziskovalce, da bi v njej objavljali svoje znanstvene ~lanke.3 2 PREGLED IZHAJANJA REVIJE IN OPIS SPREMEMB V OBDOBJU 2007–2017 Revijo Materiali in tehnologije danes izdaja In{titut za kovinske materiale in tehnologije Ljubljana (IMT), skupaj s soizdajatelji: IMPOL - Industrija aluminija Slo- venska Bistrica, SIJ METAL Ravne in TALUM Kidri- ~evo. Izdajanje revije sofinancira Javna agencija za raziskovalno dejavnost Republike Slovenije (ARRS), prej Javna agencija za knjigo (JAK). In{titut za kovinske materiale in tehnologije sodeluje na javnih razpisih, ki jih razpi{e ARRS ter tako pridobi sofinanciranje izdajanja revije. Periodika izhajanja je 6 {tevilk letno. V vsaki {tevilki je povpre~no objavljenih od 20 do 25 ~lankov in pri- spevkov, ki so: ve~inoma izvirni znanstveni ~lanki, pregledni ~lanki in strokovni ~lanki ter uvodne besede urednika ob posebnih prilo`nostih. Tak{en obseg omo- go~a korekten in kvaliteten potek procesa objave ~lankov, od oddaje ~lanka v uredni{tvo, do recenzije, prevodov in kon~no objave. V letu 2016 so z novim vodstvom pri{le tudi spre- membe.4 S {tevilko 6/2016 smo po mnogih letih brezpla~nega objavljanja ~lankov, za~eli objavo ~lankov zara~unavati. Eden od razlogov je tudi kr~enje sredstev za sofinanciranje pri ARRS. Za obi~ajne ~lanke je tako cena objave 300 EUR, za tiste ~lanke, ki bodo pred- stavljeni na letni Mednarodni konferenci o materialih in tehnologijah, ki poteka vsako leto v Portoro`u, pa je cena za objavo 150 EUR. Druga sprememba je spreje- manje in objavljanje le izvirnih znanstvenih in pre- glednih ~lankov.4 Vzrok za odlo~itev uredni{tva, da uvede objavljanje le znanstvenih prispevkov, je ta, da gre pri strokovnih ~lankih zgolj za strokovna poro~ila brez znanstvenega pristopa, in kot tako ti bolj sodijo v strokovna ali druga interna glasila. V desetletnem obdobju (2007–2017) je {tevilo oddaje ~lankov nara{~alo in v zadnjih {tirih letih naraslo do skoraj 400 letno (Slika 2), zato so se ~asovni roki za objavo podalj{ali skoraj na obdobje enega leta ali ve~, kar je po mnenju uredni{tva absolutno predolgo. Zato smo, v pomo~ avtorjem, posodobili Navodila za avtorje, ki so objavljena tako v tiskani reviji (na zadnjih straneh vsake {tevilke) kot na spletni strani revije.5 Predstavili smo vzorec oz. predlogo, kako naj bo ~lanek napisan in, da poleg oddaje ~lanka `elimo od avtorjev prejeti tudi kontrolni seznam, s katerim avtor potrdi, da je seznanjen z veljavno politiko uredni{tva. Oddaja E. NARED: POMEMBNA OBLETNICA REVIJE MATERIALI IN TEHNOLOGIJE: PETDESET LET IZHAJANJA ... 718 Materiali in tehnologije / Materials and technology 51 (2017) 5, 717–720 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS ~lankov {e vedno poteka preko e-po{te: mit@imt.si, si pa uredni{tvo revije prizadeva, da bi v prihodnje v celoti uveljavili spletno oddajanje ~lankov. Glede na `elje in zahteve mnogih avtorjev, smo v zadnjem letu uspeli skraj{ati obdobje ~akanja na objavo ~lankov in ~akalne vrste iz enega leta na 2 do 3 mesece in tako stopiti naproti tako avtorjem, kot tudi bralcem oz. ciljni publiki, ki ima v kraj{em ~asu mo`nost slediti novim spoznanjem in novostim na podro~ju materialov in tehnologij. V obdobju od leta 2007 do 2016 je v reviji Materiali in tehnologije iz{lo 1.071 ~lankov, kar je veliko ve~ kot v desetletnem obdobju pred letom 2007. V zadnjih letih se je {tevilo objavljenih ~lankov pove~evalo, saj smo v uredni{tvo prejemali vedno ve~ ~lankov, katerih avtorji so `eleli objavo. Nemara je bil najpogostej{i razlog ravno brezpla~na objava ~lankov. Najve~ji dvig je bil leta 2008, ko je iz{la posebna {tevilka revije, v katero so bili vklju~eni ~lanki iz druge mednarodne konference o to- plotni obdelavi in povr{inski obdelavi (2nd International Conference on the Heat Treatment and Surface Engi- neering of Tools and Dies), ter ponoven porast med letoma 2013 in 2014, ko so {e vedno velik dele` ~lankov predstavljali prispevki udele`encev Mednarodne kon- ference o materialih in tehnologijah v Portoro`u, katerih ~lanki so bili objavljeni {e v reviji, ~etudi so nekateri vsebinsko predstavljali zgolj poro~ila o eksperimentih in ne toliko rezultatov raznih raziskav (Slika 2). Namen tovrstnih objav je bil med drugim dati prilo`nost tudi mladim, {e ne uveljavljenim raziskovalcem, da na ta na~in predstavijo svoja dela. Interne evidence od leta 2011 dalje ka`ejo, koliko ~lankov je vsakoletno prispelo na naslov uredni{tva revije v `elji za objavo. Ta {tevilka je dosegla vrh leta 2015. S spremembami v lanskem letu (2016), ko je ured- ni{ki odbor dolo~il zara~unavanje objave, se je {tevilo prispelih ~lankov, pri~akovano, postopoma za~elo ni`ati. 2.1 [tevilo objavljenih ~lankov V prihodnje si v uredni{tvu revije Materiali in tehnologije `elimo ve~ objav doma~ih raziskovalcev in strokovnjakov ter tudi tistih iz o`jega evropskega prostora. V obdobju med 2011 in 2016 so avtorji ~lankov najpogosteje iz naslednjih dr`av (Slika 3). V letu 2011 je najve~je {tevilo ~lankov avtorjev iz Slovenije, ta {tevilka se je v zadnjih letih zelo zni`ala, v zadnjem ~asu namre~ prevladujejo predvsem objave avtorjev iz ^e{ke, Tur~ije in Poljske. Pove~uje se torej {tevilo tujih avtorjev glede na doma~e. 2.2 Pregled revije po letih izhajanja v obdobju od 2007 do 2017 Temeljit pregled revije postopno po letih izhajanja, od za~etka pa do leta 2007, je mo~ najti v `e objavljeni literaturi1, zato se bomo na tem mestu dotaknili let, ki so sledila kasneje, torej v obdobju od 2007 do danes (Tabela 1). Statisti~na analiza ka`e, da {tevilo ~lankov na splo{no zelo nara{~a (2007:43 in 2015:159), {tevilo znanstvenih ~lankov je vi{je kot {tevilo strokovnih, ve~je je tudi {tevilo ~lankov v angle{kem jeziku. Z uvedbo sprememb v letu 2016 je uredni{tvo odlo- ~ilo tudi, da bo pri ~lankih {e vedno povzetek v sloven- skem in angle{kem jeziku, ravno tako tudi naslov ter klju~ne besede, ki ozna~ujejo vsebino ~lanka v obeh jezikih Od 4. {tevilke letnika 51 pa podnapisi k slikam in tabelam ne bodo ve~ v dveh jezikih, pa~ pa le enem – in sicer v angle{kem jeziku. Tabela 1: Pregled po letih izhajanja, 2007–2017 Letnik Leto [tevilo zvezkovin {tevilk Znanstveni in strokovni ~lanki 41 2007 6 (1–6) 43 42 2008 6 (1–6) 38 43 2009 6 (1–6) 46 44 2010 6 (1–6) 45 45 2011 6 (1–6) 76 46 2012 6 (1–6) 142 47 2013 6 (1–6) 138 48 2014 6 (1–6) 157 49 2015 6 (1–6) 159 50 2016 6 (1–6) 156 51 2017 6 (1–2) 72* SKUPAJ: 1.072 *{tevilo ~lankov do {tevilke 4 (2017) E. NARED: POMEMBNA OBLETNICA REVIJE MATERIALI IN TEHNOLOGIJE: PETDESET LET IZHAJANJA ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 717–720 719 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Slika 2: Grafikon s prikazom {tevila ~lankov prispelih za objavo po letih (rde~a barva) in {tevilo dejansko objavljenih (modra barva) Slika 3: Dr`ave avtorjev objavljenih ~lankov med leti 2011–2016 Do konca izhajanja revije z imenom Kovine zlitine tehnologije (1999), je uredni{ko politiko vodil prof. dr. Franc Vodopivec (Slika 1), ki je bil na mestu urednika 16 let in tako po letih urednikovanja na drugem mestu. Tehni~no uredni{tvo revije Kovine zlitine tehnologije je vodila Jana Jamar. Izdajatelj revije je bil In{titut za kovinske materiale in tehnologije Ljubljana skupaj s soizdajatelji: ACRONI Jesenice, IMPOL Slovenska Bistrica, Kemijski in{titut Ljubljana, Koncern Slovenske `elezarne, Metal Ravne, Talum Kidri~evo, Fakulteta za strojni{tvo Ljubljana, Institut Jo`ef Stefan in Slovensko dru{tvo za tribologijo.1 Revija z novim imenom Materiali in tehnologije (ISSN 1580-2949) je nasledila obe svoji predhodnici na prelomu tiso~letja in z letom 2000 za~ela izhajati kot znanstvena serijska publikacija, ki objavlja izvirne in pregledne znanstvene ~lanke ter strokovne ~lanke, ki obravnavajo teoreti~na in prakti~na vpra{anja naravo- slovnih ved in tehnologije na podro~jih kovinskih in anorganskih materialov, polimerov, vakuumske tehnike, kompozitnih in gradbenih materialov ter nanomate- rialov.2 Njen namen je bil raz{iriti njeno interesno podro~je, da ne bi delovala le na podro~ju kovin. Odgovorni urednik revije je bil vse do leta 2011 prof. dr. Franc Vodopivec. Do istega leta je tehni~no urednikovanje vodila Jana Jamar. Z letom 2012 so se v uredni{kem odboru zgodile spremembe predvsem na kadrovskem podro~ju. Uredni{ko delo je vse do svoje upokojitve v letu 2016 vodil dr. Matja` Torkar (Slika 1), tehni~na urednica v tem ~asu je bila dr. Danijela A. Skobir Balanti~. Leta 2016 je mesto odgovornega in glavnega urednika revije prevzel dr. Paul John McGuiness (Sli- ka 1), mesto pomo~nika glavnega in odgovornega ured- nika doc. dr. Matja` Godec, ~astna glavna urednika sta prof. dr. Franc Vodopivec in dr. Matja` Torkar. Tehni~no uredni{tvo vodi Erika Nared, ki na In{titutu za kovinske materiale in tehnologije vodi tudi specialno knji`nico. Souredniki revije Materiali in tehnologije so: Igor Beli~, Jaka Burja, Aleksandra Kocijan, Djordje Man- drino, Irena Paulin, Danijela A. Skobir Balanti~, Darja Steiner Petrovi~, Bojan Podgornik iz In{tituta za kovin- ske materiale in tehnologije (IMT), Bo{tjan Markoli iz Naravoslovnotehni{ke fakultete (NTF), Sre~o [kapin in Rok Zaplotnik iz In{tituta Jo`ef Stefan (IJS) ter Ema @agar iz Kemijskega in{tituta (KI).3 Osve`en je tudi uredni{ki odbor in zasedba ~lanov mednarodnih pri- dru`enih ~lanov uredni{kega odbora ter izdajateljskega sveta. ^lanki revije MIT so indeksirani v bazah podatkov, kot so: Science Citation Index Expanded, Materials Science Citation Index in Journal Citation Reports (Science Edition). Po slednji, bazi podatkov JCR, ima revija trenutno faktor vpliva 0.436. ^lanki, objavljeni v reviji, so indeksirani tudi v mednarodnih in doma~ih sekundarnih virih/bazah: DOAJ (Directory of Open Access Journals), Google Scholar, SCOPUS, WoS, COBIB in dLib.si (Digitalna knji`nica Slovenije). Revija si, tako kot `e ve~ino ~asa svojega izhajanja, {e vedno prizadeva za vi{jo citiranost, saj bi bila posledi~no vsekakor bolj zanimiva, tako za slovenske raziskovalce kot tudi tuje, in bi pritegnila uveljavljene doma~e in tuje raziskovalce, da bi v njej objavljali svoje znanstvene ~lanke.3 2 PREGLED IZHAJANJA REVIJE IN OPIS SPREMEMB V OBDOBJU 2007–2017 Revijo Materiali in tehnologije danes izdaja In{titut za kovinske materiale in tehnologije Ljubljana (IMT), skupaj s soizdajatelji: IMPOL - Industrija aluminija Slo- venska Bistrica, SIJ METAL Ravne in TALUM Kidri- ~evo. Izdajanje revije sofinancira Javna agencija za raziskovalno dejavnost Republike Slovenije (ARRS), prej Javna agencija za knjigo (JAK). In{titut za kovinske materiale in tehnologije sodeluje na javnih razpisih, ki jih razpi{e ARRS ter tako pridobi sofinanciranje izdajanja revije. Periodika izhajanja je 6 {tevilk letno. V vsaki {tevilki je povpre~no objavljenih od 20 do 25 ~lankov in pri- spevkov, ki so: ve~inoma izvirni znanstveni ~lanki, pregledni ~lanki in strokovni ~lanki ter uvodne besede urednika ob posebnih prilo`nostih. Tak{en obseg omo- go~a korekten in kvaliteten potek procesa objave ~lankov, od oddaje ~lanka v uredni{tvo, do recenzije, prevodov in kon~no objave. V letu 2016 so z novim vodstvom pri{le tudi spre- membe.4 S {tevilko 6/2016 smo po mnogih letih brezpla~nega objavljanja ~lankov, za~eli objavo ~lankov zara~unavati. Eden od razlogov je tudi kr~enje sredstev za sofinanciranje pri ARRS. Za obi~ajne ~lanke je tako cena objave 300 EUR, za tiste ~lanke, ki bodo pred- stavljeni na letni Mednarodni konferenci o materialih in tehnologijah, ki poteka vsako leto v Portoro`u, pa je cena za objavo 150 EUR. Druga sprememba je spreje- manje in objavljanje le izvirnih znanstvenih in pre- glednih ~lankov.4 Vzrok za odlo~itev uredni{tva, da uvede objavljanje le znanstvenih prispevkov, je ta, da gre pri strokovnih ~lankih zgolj za strokovna poro~ila brez znanstvenega pristopa, in kot tako ti bolj sodijo v strokovna ali druga interna glasila. V desetletnem obdobju (2007–2017) je {tevilo oddaje ~lankov nara{~alo in v zadnjih {tirih letih naraslo do skoraj 400 letno (Slika 2), zato so se ~asovni roki za objavo podalj{ali skoraj na obdobje enega leta ali ve~, kar je po mnenju uredni{tva absolutno predolgo. Zato smo, v pomo~ avtorjem, posodobili Navodila za avtorje, ki so objavljena tako v tiskani reviji (na zadnjih straneh vsake {tevilke) kot na spletni strani revije.5 Predstavili smo vzorec oz. predlogo, kako naj bo ~lanek napisan in, da poleg oddaje ~lanka `elimo od avtorjev prejeti tudi kontrolni seznam, s katerim avtor potrdi, da je seznanjen z veljavno politiko uredni{tva. Oddaja E. NARED: POMEMBNA OBLETNICA REVIJE MATERIALI IN TEHNOLOGIJE: PETDESET LET IZHAJANJA ... 718 Materiali in tehnologije / Materials and technology 51 (2017) 5, 717–720 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS ~lankov {e vedno poteka preko e-po{te: mit@imt.si, si pa uredni{tvo revije prizadeva, da bi v prihodnje v celoti uveljavili spletno oddajanje ~lankov. Glede na `elje in zahteve mnogih avtorjev, smo v zadnjem letu uspeli skraj{ati obdobje ~akanja na objavo ~lankov in ~akalne vrste iz enega leta na 2 do 3 mesece in tako stopiti naproti tako avtorjem, kot tudi bralcem oz. ciljni publiki, ki ima v kraj{em ~asu mo`nost slediti novim spoznanjem in novostim na podro~ju materialov in tehnologij. V obdobju od leta 2007 do 2016 je v reviji Materiali in tehnologije iz{lo 1.071 ~lankov, kar je veliko ve~ kot v desetletnem obdobju pred letom 2007. V zadnjih letih se je {tevilo objavljenih ~lankov pove~evalo, saj smo v uredni{tvo prejemali vedno ve~ ~lankov, katerih avtorji so `eleli objavo. Nemara je bil najpogostej{i razlog ravno brezpla~na objava ~lankov. Najve~ji dvig je bil leta 2008, ko je iz{la posebna {tevilka revije, v katero so bili vklju~eni ~lanki iz druge mednarodne konference o to- plotni obdelavi in povr{inski obdelavi (2nd International Conference on the Heat Treatment and Surface Engi- neering of Tools and Dies), ter ponoven porast med letoma 2013 in 2014, ko so {e vedno velik dele` ~lankov predstavljali prispevki udele`encev Mednarodne kon- ference o materialih in tehnologijah v Portoro`u, katerih ~lanki so bili objavljeni {e v reviji, ~etudi so nekateri vsebinsko predstavljali zgolj poro~ila o eksperimentih in ne toliko rezultatov raznih raziskav (Slika 2). Namen tovrstnih objav je bil med drugim dati prilo`nost tudi mladim, {e ne uveljavljenim raziskovalcem, da na ta na~in predstavijo svoja dela. Interne evidence od leta 2011 dalje ka`ejo, koliko ~lankov je vsakoletno prispelo na naslov uredni{tva revije v `elji za objavo. Ta {tevilka je dosegla vrh leta 2015. S spremembami v lanskem letu (2016), ko je ured- ni{ki odbor dolo~il zara~unavanje objave, se je {tevilo prispelih ~lankov, pri~akovano, postopoma za~elo ni`ati. 2.1 [tevilo objavljenih ~lankov V prihodnje si v uredni{tvu revije Materiali in tehnologije `elimo ve~ objav doma~ih raziskovalcev in strokovnjakov ter tudi tistih iz o`jega evropskega prostora. V obdobju med 2011 in 2016 so avtorji ~lankov najpogosteje iz naslednjih dr`av (Slika 3). V letu 2011 je najve~je {tevilo ~lankov avtorjev iz Slovenije, ta {tevilka se je v zadnjih letih zelo zni`ala, v zadnjem ~asu namre~ prevladujejo predvsem objave avtorjev iz ^e{ke, Tur~ije in Poljske. Pove~uje se torej {tevilo tujih avtorjev glede na doma~e. 2.2 Pregled revije po letih izhajanja v obdobju od 2007 do 2017 Temeljit pregled revije postopno po letih izhajanja, od za~etka pa do leta 2007, je mo~ najti v `e objavljeni literaturi1, zato se bomo na tem mestu dotaknili let, ki so sledila kasneje, torej v obdobju od 2007 do danes (Tabela 1). Statisti~na analiza ka`e, da {tevilo ~lankov na splo{no zelo nara{~a (2007:43 in 2015:159), {tevilo znanstvenih ~lankov je vi{je kot {tevilo strokovnih, ve~je je tudi {tevilo ~lankov v angle{kem jeziku. Z uvedbo sprememb v letu 2016 je uredni{tvo odlo- ~ilo tudi, da bo pri ~lankih {e vedno povzetek v sloven- skem in angle{kem jeziku, ravno tako tudi naslov ter klju~ne besede, ki ozna~ujejo vsebino ~lanka v obeh jezikih Od 4. {tevilke letnika 51 pa podnapisi k slikam in tabelam ne bodo ve~ v dveh jezikih, pa~ pa le enem – in sicer v angle{kem jeziku. Tabela 1: Pregled po letih izhajanja, 2007–2017 Letnik Leto [tevilo zvezkovin {tevilk Znanstveni in strokovni ~lanki 41 2007 6 (1–6) 43 42 2008 6 (1–6) 38 43 2009 6 (1–6) 46 44 2010 6 (1–6) 45 45 2011 6 (1–6) 76 46 2012 6 (1–6) 142 47 2013 6 (1–6) 138 48 2014 6 (1–6) 157 49 2015 6 (1–6) 159 50 2016 6 (1–6) 156 51 2017 6 (1–2) 72* SKUPAJ: 1.072 *{tevilo ~lankov do {tevilke 4 (2017) E. NARED: POMEMBNA OBLETNICA REVIJE MATERIALI IN TEHNOLOGIJE: PETDESET LET IZHAJANJA ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 717–720 719 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Slika 2: Grafikon s prikazom {tevila ~lankov prispelih za objavo po letih (rde~a barva) in {tevilo dejansko objavljenih (modra barva) Slika 3: Dr`ave avtorjev objavljenih ~lankov med leti 2011–2016 2.3 Pregled revije po tipologiji [tevilo ~lankov, ki spadajo v skupino izvirnih znan- stvenih ~lankov se je z leti vi{alo, najve~ jih je bilo v letu 2016, najmanj pa v letu 2007. [tevilo strokovnih ~lan- kov, ki so z letom 2016, po odlo~itvi uredni{tva ukinjeni, je obdr`alo konstantno {tevilko, od 30 do 40 letno. Tabela 2: Pregled prispevkov v letih od 2007 do 2017 po tipologiji Leto Tipologija 1.01 (izvirni znan- stveni ~lanek) 1.02 (pregledni znanstveni ~lanek) 1.04 (strokovni ~lanek) 1.20 (predgovor, spremna beseda) 2007 39 4 1 0 2008 30 6 8 0 2009 38 3 9 0 2010 43 8 7 0 2011 79 4 14 0 2012 81 0 30 3 2013 98 6 30 0 2014 118 2 39 0 2015 114 6 35 0 2016 125 2 30 2 2017 (do {t. 3)* 46 2 10 1 *podatki pridobljeni pred izidom 4. {tevilke 3 DOSTOPNOST REVIJE Revija Materiali in tehnologije je `e v devetdesetih letih prej{njega stoletja, ko se je imenovala {e Kovine zlitine tehnologije, stopila naproti ideji, da bi bila bolj odprta in la`je dostopna {ir{i javnosti. V skladu z `eljami in dejanji uredni{tva in navsezadnje te`njami {ir{e javnosti ter splo{nih trendov tistega ~asa, je v letu 1998 postala dostopna na svetovnem spletu v celotnem bese- dilu in jo je bilo sprva mo~ najti na spletni strani Centralne tehni{ke knji`nice (repozitorij CTK), kasneje pa na spletni strani www.imt.si, kjer je revija dostopna {e danes. Revija (njeni predhodnici @elezarski zbornik in Ko- vine zlitine tehnologije) je bila v Narodni in univerzitetni knji`nici (NUK) vklju~ena v projekt digitalizacije in je danes na voljo v digitalni obliki tako na spletnem portalu pri Digitalni knji`nici Slovenije pri NUK, kot tudi na spletni strani revije MIT. Ob petdeseti obletnici izhajanja revije smo se v uredni{tvu odlo~ili, da na spletni strani revije uredimo celoten arhiv vseh {tevilk revije, tudi {te- vilk njenih predhodnic, ki so doslej manjkale: @elezarski zbornik (ISSN 0372-8633) in Kovine zlitine tehnologije (ISSN 1318-0010). Revijo Materiali in tehnologije je mogo~e prelistati kadarkoli in od koderkoli (~e le imate v bli`ini internet), in sicer od prve {tevilke @elezarskega zbornika iz leta 1967 pa do vseh ~lankov revije MIT do danes. Ve~ino {tevilk revije v tiskani obliki hranimo tudi v knji`nici IMT. Skupno je bilo tako pripravljenih, in v spletni arhiv dodanih, 123 datotek; 102 {tevilki @elezar- skega zbornika in 21 {tevilk revije KZT.6 4 ZAKLJU^EK Za~etki revije in objava ~lankov samo iz podro~ja `elezarstva in metalurgije pred petdesetimi leti, ko je industrija jekla in `elezarstva predstavljala najve~ji dele` tovrstne dejavnosti v gospodarstvu takratne Jugoslavije in kasneje dr`ave Slovenije, so bili med drugim ravno tako podlaga za nadaljnje raziskovanje in uveljavljanje metalurgije kot panoge, ki je do danes po~asi, a vztrajno pre{la v skoraj vse pore industrije in je v dana{njem ~asu razvoja tehnologije na izredno visoki ravni. Danes tako z visoko tehnolo{kimi principi in najnovej{imi tehnolo- gijami na podro~ju tako kovinskih, kot tudi drugih materialov, dosegamo zavidljive rezultate. Napredni ma- teriali, napredne proizvodne tehnologije, nanotehnologije in biotehnologije so klju~na podro~ja, identificirana v strategiji pametne specializacije, ki bodo omogo~ala evropski in slovenski industriji ohraniti mednarodno globalno konkuren~nost in izkoristiti nove trge. Z njimi sta metalurgija in kemijska industrija nelo~ljivo povezani in sta del opredeljenih klju~nih tehnologij. Pri tem razvoj novih materialov in tehnologij pomeni vstop novih, do sedaj neznanih mo`nosti, na tr`i{~e, kjer imajo ra- ziskave, razvoj in inovacije zelo pomembno vlogo.7 Del raziskav in preizkusov, je predstavljen tudi skozi prispev- ke in ~lanke v reviji MIT in tudi v drugih medijih,8 ki s tem {iri nova dognanja, spoznanja in znanja na {ir{i krog raziskovalcev, znanstvenikov in drugo potencialno publi- ko. Revija Materiali in tehnologije je danes sicer pri- znana tudi izven meja Slovenije, vendar pa si v prihodnje {e vedno `elimo, da bi v reviji objavljalo ve~je {tevilo strokovnjakov in raziskovalcev iz Slovenije in evrop- skega prostora. 5 LITERATURA 1 N. Jamar, J. Jamar, Zgodovina znanstvene serijske publikacije Materiali in tehnologije/Materials and Technology = Historical over- view of the scientific journal Materiali in tehnologije/Materials and Technology, Mater. Tehnol., 41 (2007) 1, 13–19 2 http://mit.imt.si/Revija/index-slo.html, 28.7.2017 3 http://mit.imt.si/Revija/information-slo.html, 28.7.2017 4 P. McGuiness, Predgovor urednika=Editor’s preface, Mater. Tehnol., 50 (2016) 5, str. 639–640 5 http://mit.imt.si/Revija/authors-slo.html, 28.07.2017 6 http://mit.imt.si/Revija/archive.html, 29.07.2017 7 https://www.gzs.si/Novice/ArticleId/58642/srip-matpro, 29.7.2017 8 M. Godec, Najve~ji hit je 3D tisk, Glas gospodarstva, Naj materiali, pano`na {tevilka, april (2017), 30–31 E. NARED: POMEMBNA OBLETNICA REVIJE MATERIALI IN TEHNOLOGIJE: PETDESET LET IZHAJANJA ... 720 Materiali in tehnologije / Materials and technology 51 (2017) 5, 717–720 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT 721–727 INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT PREISKAVA MEJ ZRN V ZLITINI 263 PO POSEBNI TOPLOTNI OBDELAVI Ivan Slatkovský, Mária Dománková, Martin Sahul Slovak University of Technology Bratislava, Faculty of Materials Science and Technology, Institute of Materials Science, Bottová 25, 917 24, Trnava, Slovak Republic ivan.slatkovsky@stuba.sk Prejem rokopisa – received: 2016-06-20; sprejem za objavo – accepted for publication: 2017-01-24 doi:10.17222/mit.2016.115 Alloy 263 is well known for its very good creep resistance and also for its weldability. These kinds of properties are appreciated in the power-plant industry where Alloy 263 is used for shafts in a high-pressure circle. One of the possible ways to improve the properties of superalloys, including Alloy 263, is through the effect of the grain-boundary serration (GBS) which, as research indicates, is associated with the improvement of the creep resistance that can lead to an increased efficiency of coal power plants. Grain-boundary serration was observed in different kinds of superalloys although the formation mechanism of serration has not been clearly explained yet. Some researchers reported that the formation of serration is associated with the change in the character of the precipitates at grain boundaries. This paper deals with an investigation of the grain boundaries in Alloy 263 using two different kinds of heat treatment. To form serrated grain boundaries in Material A (MA), slow controlled cooling from the temperature of solution annealing to 800 °C was carried out. Standard heat treatment of Alloy 263 was performed on material B (MB). Experimental techniques of scanning electron microscopy (SEM) and transmission electron microscopy (TEM), including electron diffraction, were used to analyze the microstructure, determine the character of the grain boundaries and identify the secondary particles at the grain boundaries. Keywords: Alloy 263, grain-boundary serration, precipitates Zlitina 263 je znana po zelo dobri odpornosti proti lezenju in tudi po dobri varivosti. Te vrste lastnosti so zelo cenjene v termoelektrarnah, kjer se zlitina 263 uporablja za gredi v visokotla~nem delu turbin. Eden od mo`nih na~inov za izbolj{anje lastnosti superzlitin, vklju~no z zlitino 263, je tvorba (nastanek) nazob~anih kristalnih mej (angl. GBS). Preiskava je pokazala, da je ta fenomen povezan z izbolj{anjem odpornosti proti lezenju, kar lahko vodi k ve~ji u~inkovitosti termoelektrarn na premog. U~inek nazob~anosti mej kristalnih zrn so opazili pri razli~nih vrstah superzlitin, ~eprav mehanizem tvorbe {e ni popolnoma pojasnjen. Nekateri raziskovalci so ugotovili, da je tvorba nazob~anosti povezana s spremembo lastnosti izlo~kov na mejah med zrni. Prispevek se ukvarja s preiskavo mej med kristalnimi zrni v zlitini 263 z dvema razli~nima vrstama toplotne obdelave. Nastanek nazob~anih mej kristalnih zrn materiala A (MA), je bil povzro~en s po~asnim kontroliranim ohlajanjem iz temperature raztopnega `arjenja, ki je bila 800 ° C. Standardna toplotna obdelava se je izvajala za zlitino 263 - material B (MB). Za analizo mikrostrukture so uporabili vrsti~ni elektronski mikroskop (SEM) in presevno elektronsko mikroskopijo (TEM), vklju~no z elektronsko difrakcijo. Na ta na~in so dolo~ili lastnosti mej kristalnih zrn in sekundarnih delcev na mejah zrn. Klju~ne besede: zlitina 263, nazob~anost mej kristalnih zrn, izlo~ki 1 INTRODUCTION It is a well-known fact that a reduction in the CO2 emissions produced by coal power plants is one of the major goals for the countries all over the world. A possible way to reduce the CO2 emissions is to improve thermal efficiencies through super critical and ultra- super critical technologies in power plants, where efficiencies above 40 % could be reached. Hence, to achieve this kind of efficiency, the materials like nickel- based superalloys are considered to be used (Fig- ure 1).1–3 To prolong the life time of parts when the tempera- ture exceeds 700 °C, scientists are looking for new paths of material processing. In the case of heat treatment, one of the possible ways is a modification of grain boun- daries (GBs) when the character of the boundaries changes from straight to zigzag. The phenomenon of grain-boundary serration (GBS) was observed in the nickel-based superalloys and in austenitic stainless steels. The formation of serrations on grain boundaries (GBs) has not been fully described yet. However, it is a known fact that in the case of nickel- based superalloys, the GBS is closely related to the slow controlled cooling from the temperature of the solution annealing. Early studies of serration in the nickel-based alloys were focused on the interaction between the grain boundary and ’ phase.4,5 Until now, researchers have found that the serration is associated not only with the presence of the ’ phase, but also with other precipitates like M23C6 carbide, the  phase or the ’’ phase on the GBs that were observed in different kinds of superalloys. The formation of these secondary particles could be highly related to the serrated grain boundaries and their formation.6–13 Recently, contemporary authors have Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 721 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 669.018:620.1:620.193 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)721(2017) 2.3 Pregled revije po tipologiji [tevilo ~lankov, ki spadajo v skupino izvirnih znan- stvenih ~lankov se je z leti vi{alo, najve~ jih je bilo v letu 2016, najmanj pa v letu 2007. [tevilo strokovnih ~lan- kov, ki so z letom 2016, po odlo~itvi uredni{tva ukinjeni, je obdr`alo konstantno {tevilko, od 30 do 40 letno. Tabela 2: Pregled prispevkov v letih od 2007 do 2017 po tipologiji Leto Tipologija 1.01 (izvirni znan- stveni ~lanek) 1.02 (pregledni znanstveni ~lanek) 1.04 (strokovni ~lanek) 1.20 (predgovor, spremna beseda) 2007 39 4 1 0 2008 30 6 8 0 2009 38 3 9 0 2010 43 8 7 0 2011 79 4 14 0 2012 81 0 30 3 2013 98 6 30 0 2014 118 2 39 0 2015 114 6 35 0 2016 125 2 30 2 2017 (do {t. 3)* 46 2 10 1 *podatki pridobljeni pred izidom 4. {tevilke 3 DOSTOPNOST REVIJE Revija Materiali in tehnologije je `e v devetdesetih letih prej{njega stoletja, ko se je imenovala {e Kovine zlitine tehnologije, stopila naproti ideji, da bi bila bolj odprta in la`je dostopna {ir{i javnosti. V skladu z `eljami in dejanji uredni{tva in navsezadnje te`njami {ir{e javnosti ter splo{nih trendov tistega ~asa, je v letu 1998 postala dostopna na svetovnem spletu v celotnem bese- dilu in jo je bilo sprva mo~ najti na spletni strani Centralne tehni{ke knji`nice (repozitorij CTK), kasneje pa na spletni strani www.imt.si, kjer je revija dostopna {e danes. Revija (njeni predhodnici @elezarski zbornik in Ko- vine zlitine tehnologije) je bila v Narodni in univerzitetni knji`nici (NUK) vklju~ena v projekt digitalizacije in je danes na voljo v digitalni obliki tako na spletnem portalu pri Digitalni knji`nici Slovenije pri NUK, kot tudi na spletni strani revije MIT. Ob petdeseti obletnici izhajanja revije smo se v uredni{tvu odlo~ili, da na spletni strani revije uredimo celoten arhiv vseh {tevilk revije, tudi {te- vilk njenih predhodnic, ki so doslej manjkale: @elezarski zbornik (ISSN 0372-8633) in Kovine zlitine tehnologije (ISSN 1318-0010). Revijo Materiali in tehnologije je mogo~e prelistati kadarkoli in od koderkoli (~e le imate v bli`ini internet), in sicer od prve {tevilke @elezarskega zbornika iz leta 1967 pa do vseh ~lankov revije MIT do danes. Ve~ino {tevilk revije v tiskani obliki hranimo tudi v knji`nici IMT. Skupno je bilo tako pripravljenih, in v spletni arhiv dodanih, 123 datotek; 102 {tevilki @elezar- skega zbornika in 21 {tevilk revije KZT.6 4 ZAKLJU^EK Za~etki revije in objava ~lankov samo iz podro~ja `elezarstva in metalurgije pred petdesetimi leti, ko je industrija jekla in `elezarstva predstavljala najve~ji dele` tovrstne dejavnosti v gospodarstvu takratne Jugoslavije in kasneje dr`ave Slovenije, so bili med drugim ravno tako podlaga za nadaljnje raziskovanje in uveljavljanje metalurgije kot panoge, ki je do danes po~asi, a vztrajno pre{la v skoraj vse pore industrije in je v dana{njem ~asu razvoja tehnologije na izredno visoki ravni. Danes tako z visoko tehnolo{kimi principi in najnovej{imi tehnolo- gijami na podro~ju tako kovinskih, kot tudi drugih materialov, dosegamo zavidljive rezultate. Napredni ma- teriali, napredne proizvodne tehnologije, nanotehnologije in biotehnologije so klju~na podro~ja, identificirana v strategiji pametne specializacije, ki bodo omogo~ala evropski in slovenski industriji ohraniti mednarodno globalno konkuren~nost in izkoristiti nove trge. Z njimi sta metalurgija in kemijska industrija nelo~ljivo povezani in sta del opredeljenih klju~nih tehnologij. Pri tem razvoj novih materialov in tehnologij pomeni vstop novih, do sedaj neznanih mo`nosti, na tr`i{~e, kjer imajo ra- ziskave, razvoj in inovacije zelo pomembno vlogo.7 Del raziskav in preizkusov, je predstavljen tudi skozi prispev- ke in ~lanke v reviji MIT in tudi v drugih medijih,8 ki s tem {iri nova dognanja, spoznanja in znanja na {ir{i krog raziskovalcev, znanstvenikov in drugo potencialno publi- ko. Revija Materiali in tehnologije je danes sicer pri- znana tudi izven meja Slovenije, vendar pa si v prihodnje {e vedno `elimo, da bi v reviji objavljalo ve~je {tevilo strokovnjakov in raziskovalcev iz Slovenije in evrop- skega prostora. 5 LITERATURA 1 N. Jamar, J. Jamar, Zgodovina znanstvene serijske publikacije Materiali in tehnologije/Materials and Technology = Historical over- view of the scientific journal Materiali in tehnologije/Materials and Technology, Mater. Tehnol., 41 (2007) 1, 13–19 2 http://mit.imt.si/Revija/index-slo.html, 28.7.2017 3 http://mit.imt.si/Revija/information-slo.html, 28.7.2017 4 P. McGuiness, Predgovor urednika=Editor’s preface, Mater. Tehnol., 50 (2016) 5, str. 639–640 5 http://mit.imt.si/Revija/authors-slo.html, 28.07.2017 6 http://mit.imt.si/Revija/archive.html, 29.07.2017 7 https://www.gzs.si/Novice/ArticleId/58642/srip-matpro, 29.7.2017 8 M. Godec, Najve~ji hit je 3D tisk, Glas gospodarstva, Naj materiali, pano`na {tevilka, april (2017), 30–31 E. NARED: POMEMBNA OBLETNICA REVIJE MATERIALI IN TEHNOLOGIJE: PETDESET LET IZHAJANJA ... 720 Materiali in tehnologije / Materials and technology 51 (2017) 5, 717–720 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT 721–727 INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT PREISKAVA MEJ ZRN V ZLITINI 263 PO POSEBNI TOPLOTNI OBDELAVI Ivan Slatkovský, Mária Dománková, Martin Sahul Slovak University of Technology Bratislava, Faculty of Materials Science and Technology, Institute of Materials Science, Bottová 25, 917 24, Trnava, Slovak Republic ivan.slatkovsky@stuba.sk Prejem rokopisa – received: 2016-06-20; sprejem za objavo – accepted for publication: 2017-01-24 doi:10.17222/mit.2016.115 Alloy 263 is well known for its very good creep resistance and also for its weldability. These kinds of properties are appreciated in the power-plant industry where Alloy 263 is used for shafts in a high-pressure circle. One of the possible ways to improve the properties of superalloys, including Alloy 263, is through the effect of the grain-boundary serration (GBS) which, as research indicates, is associated with the improvement of the creep resistance that can lead to an increased efficiency of coal power plants. Grain-boundary serration was observed in different kinds of superalloys although the formation mechanism of serration has not been clearly explained yet. Some researchers reported that the formation of serration is associated with the change in the character of the precipitates at grain boundaries. This paper deals with an investigation of the grain boundaries in Alloy 263 using two different kinds of heat treatment. To form serrated grain boundaries in Material A (MA), slow controlled cooling from the temperature of solution annealing to 800 °C was carried out. Standard heat treatment of Alloy 263 was performed on material B (MB). Experimental techniques of scanning electron microscopy (SEM) and transmission electron microscopy (TEM), including electron diffraction, were used to analyze the microstructure, determine the character of the grain boundaries and identify the secondary particles at the grain boundaries. Keywords: Alloy 263, grain-boundary serration, precipitates Zlitina 263 je znana po zelo dobri odpornosti proti lezenju in tudi po dobri varivosti. Te vrste lastnosti so zelo cenjene v termoelektrarnah, kjer se zlitina 263 uporablja za gredi v visokotla~nem delu turbin. Eden od mo`nih na~inov za izbolj{anje lastnosti superzlitin, vklju~no z zlitino 263, je tvorba (nastanek) nazob~anih kristalnih mej (angl. GBS). Preiskava je pokazala, da je ta fenomen povezan z izbolj{anjem odpornosti proti lezenju, kar lahko vodi k ve~ji u~inkovitosti termoelektrarn na premog. U~inek nazob~anosti mej kristalnih zrn so opazili pri razli~nih vrstah superzlitin, ~eprav mehanizem tvorbe {e ni popolnoma pojasnjen. Nekateri raziskovalci so ugotovili, da je tvorba nazob~anosti povezana s spremembo lastnosti izlo~kov na mejah med zrni. Prispevek se ukvarja s preiskavo mej med kristalnimi zrni v zlitini 263 z dvema razli~nima vrstama toplotne obdelave. Nastanek nazob~anih mej kristalnih zrn materiala A (MA), je bil povzro~en s po~asnim kontroliranim ohlajanjem iz temperature raztopnega `arjenja, ki je bila 800 ° C. Standardna toplotna obdelava se je izvajala za zlitino 263 - material B (MB). Za analizo mikrostrukture so uporabili vrsti~ni elektronski mikroskop (SEM) in presevno elektronsko mikroskopijo (TEM), vklju~no z elektronsko difrakcijo. Na ta na~in so dolo~ili lastnosti mej kristalnih zrn in sekundarnih delcev na mejah zrn. Klju~ne besede: zlitina 263, nazob~anost mej kristalnih zrn, izlo~ki 1 INTRODUCTION It is a well-known fact that a reduction in the CO2 emissions produced by coal power plants is one of the major goals for the countries all over the world. A possible way to reduce the CO2 emissions is to improve thermal efficiencies through super critical and ultra- super critical technologies in power plants, where efficiencies above 40 % could be reached. Hence, to achieve this kind of efficiency, the materials like nickel- based superalloys are considered to be used (Fig- ure 1).1–3 To prolong the life time of parts when the tempera- ture exceeds 700 °C, scientists are looking for new paths of material processing. In the case of heat treatment, one of the possible ways is a modification of grain boun- daries (GBs) when the character of the boundaries changes from straight to zigzag. The phenomenon of grain-boundary serration (GBS) was observed in the nickel-based superalloys and in austenitic stainless steels. The formation of serrations on grain boundaries (GBs) has not been fully described yet. However, it is a known fact that in the case of nickel- based superalloys, the GBS is closely related to the slow controlled cooling from the temperature of the solution annealing. Early studies of serration in the nickel-based alloys were focused on the interaction between the grain boundary and ’ phase.4,5 Until now, researchers have found that the serration is associated not only with the presence of the ’ phase, but also with other precipitates like M23C6 carbide, the  phase or the ’’ phase on the GBs that were observed in different kinds of superalloys. The formation of these secondary particles could be highly related to the serrated grain boundaries and their formation.6–13 Recently, contemporary authors have Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 721 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 669.018:620.1:620.193 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)721(2017) found that in some alloys, in the early stages of the formation, GBS occurred in the absence of the adjacent coarse ’ particle or M23C6 carbide.12–14 Furthermore, H. U. Hong et al.15 observed in their research that the serrations on GBs were formed in the material after a long solution annealing (2000 min) without any slow controlled cooling. Also, no secondary particles were observed. As reported by the authors of references,10–13 GBS leads to a morphological change of the carbides on GBs, from granular to planar. A growth of secondary particles on the serrated grain boundaries was observed mostly at a low angle and on special GBs. The authors of referen- ces10,11 reported that the formation of planar M23C6 carbides is orientated in the {111} plane. Also, the orientation relationship between the carbides and the matrix was observed. Therefore, the purpose of this study is to investigate the effect of special heat treatment on GBs and identify the secondary phases formed at the GBs in Alloy 263. 2 EXPERIMENTAL PART In order to investigate GBS, commercially available hot-rolled nickel-based superalloy 263 was used in the study. The chemical composition of the examined alloy is given in Table 1. A two-stage special heat-treatment method was designed to form a GBS in material A, based on the previous studies.6–13 The solution heat treatment followed by slow cooling at a carefully controlled rate (until the aging temperature was reached) was found to be necessary to generate serrated GB structures. Material A was solution annealed at 1150 °C for 80 min and cooled slowly to 800 °C at a cooling rate of 3 °C/min, followed by water quenching. After that, precipitation hardening at 800 °C for 4 h was performed and followed by air cooling. On material B, the standard two-stage heat treatment of Alloy 263 (solution annealing at 1150 °C for 80 min followed by water quenching and precipitation hardening at 800 °C for 4 h, cooled in the air) was applied. For a SEM analysis, the samples were sliced, mecha- nically grinded and polished. The etching solution for the SEM observation was a solution containing 5 g FeCl3, 15 mL HCl, 2 mL HNO3 and 60 mL ethanol used for 10–15 s in order to reveal the GB configuration and carbides. A JEOL 7600 F scanning electron microscope was used for the observation of the GBs. Thin foils were prepared for a detailed grain-boun- dary observation using mechanical grinding to a thick- ness of about 0.1 mm, and then electrolytically etched. Etching was done on TENUPOL 5 in a solution of perchloric acid and methanol (1:9). The temperature during etching was -30 °C and the voltage was 25 V. For the detailed observation, a transmission electron micro- scope, JEOL 200 CX with an accelerating voltage of 200 kV, was used. To identify secondary particles at the GBs, an electron-diffraction analysis was applied. Extraction replicas were prepared in the following route: samples were first prepared using standard metallographic techniques. After that, they were etched with the above-mentioned solution for 15 s. A thin carbon film was sputtered onto the etched surface of the samples. The carbon film (replica) was subsequently electrolytically extracted, using 8 % perchloric acid, at a bias of 10 V. 3 RESULTS Figure 2 shows the microstructure of material A after the special heat treatment. Light microscopy revealed a polyhedral grain, the presence of twins as well as the grain-size heterogeneity in material A. Two types of grain-boundary morphology were observed with light microscopy. Straight boundaries with a percentage share of approximately 38 % of the total of 254 observed grain boundaries were located in the structure of material A. Serrated boundaries, as the second type, covered approximately 39 % of the noticed boundaries. In some cases, the type of boundary could not be determined owing to the state of the boundary. As a reference sample, material B (Figure 3) was processed with the conventional heat treatment. In material B, we also observed the base microstructure I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT 722 Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Table 1: Nominal chemical composition of Alloy 263 Chemical-composition limits Weight % Ni Cr Co Mo C Al Ti Al+Ti Mn 263 Bal 19–21 19–21 5.6–6.1 0.04–0.08 0.60 max 1.90–2.40 2.40–2.80 0.60 max Weight % Si Fe Cu B Pb S Ag Bi 263 0.40 max 0.07 max 0.020 max 0.005 max 0.0020 0.007 max 0.005 max 0.0001 max Figure 1: Net efficiency development in the case of hard coal-fired power plants considering different structural alloys3 with polyhedral grains with the heterogeneity of the grain size and twins similar to those in material A. As expected, no serrated boundaries were observed in material B after the conventional heat treatment. The grain-boundary details taken with SEM indicate the presence of precipitates on the serrated and straight grain boundaries in material A as well as in material B (Figures 4 and 5). Results of the EDX analysis taken from material A (Figure 6) reveal the eventual presence of carbon-rich secondary particles, which could possibly indicate the presence of carbides at the grain boundaries. The occurrence of the other elements at the grain boun- daries was not significant. Comparable results were also noticed for material B. To confirm the presence of the secondary particles formed at the boundaries and to identify the chemical nature of these particles, ED and EDX using the TEM were taken for materials A and B, and the results are shown below. The authors of reference16 predicted and identified typical kinds of the secondary particles for this alloy, allowing us to expect mainly the presence of the MC, M6C and M23C6 carbides. Figures 7 to 9 summarize the identified phases on the replicas for material A. Electron diffraction spectra as well as the EDX analysis of the extracted precipitates I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 723 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: SEM, material B: a) straight grain-boundary triple point and b) detail of precipitatesFigure 3: Material B Figure 4: SEM, material A: a) serrated grain-boundary triple point and b) detail of precipitates Figure 2: Material A found that in some alloys, in the early stages of the formation, GBS occurred in the absence of the adjacent coarse ’ particle or M23C6 carbide.12–14 Furthermore, H. U. Hong et al.15 observed in their research that the serrations on GBs were formed in the material after a long solution annealing (2000 min) without any slow controlled cooling. Also, no secondary particles were observed. As reported by the authors of references,10–13 GBS leads to a morphological change of the carbides on GBs, from granular to planar. A growth of secondary particles on the serrated grain boundaries was observed mostly at a low angle and on special GBs. The authors of referen- ces10,11 reported that the formation of planar M23C6 carbides is orientated in the {111} plane. Also, the orientation relationship between the carbides and the matrix was observed. Therefore, the purpose of this study is to investigate the effect of special heat treatment on GBs and identify the secondary phases formed at the GBs in Alloy 263. 2 EXPERIMENTAL PART In order to investigate GBS, commercially available hot-rolled nickel-based superalloy 263 was used in the study. The chemical composition of the examined alloy is given in Table 1. A two-stage special heat-treatment method was designed to form a GBS in material A, based on the previous studies.6–13 The solution heat treatment followed by slow cooling at a carefully controlled rate (until the aging temperature was reached) was found to be necessary to generate serrated GB structures. Material A was solution annealed at 1150 °C for 80 min and cooled slowly to 800 °C at a cooling rate of 3 °C/min, followed by water quenching. After that, precipitation hardening at 800 °C for 4 h was performed and followed by air cooling. On material B, the standard two-stage heat treatment of Alloy 263 (solution annealing at 1150 °C for 80 min followed by water quenching and precipitation hardening at 800 °C for 4 h, cooled in the air) was applied. For a SEM analysis, the samples were sliced, mecha- nically grinded and polished. The etching solution for the SEM observation was a solution containing 5 g FeCl3, 15 mL HCl, 2 mL HNO3 and 60 mL ethanol used for 10–15 s in order to reveal the GB configuration and carbides. A JEOL 7600 F scanning electron microscope was used for the observation of the GBs. Thin foils were prepared for a detailed grain-boun- dary observation using mechanical grinding to a thick- ness of about 0.1 mm, and then electrolytically etched. Etching was done on TENUPOL 5 in a solution of perchloric acid and methanol (1:9). The temperature during etching was -30 °C and the voltage was 25 V. For the detailed observation, a transmission electron micro- scope, JEOL 200 CX with an accelerating voltage of 200 kV, was used. To identify secondary particles at the GBs, an electron-diffraction analysis was applied. Extraction replicas were prepared in the following route: samples were first prepared using standard metallographic techniques. After that, they were etched with the above-mentioned solution for 15 s. A thin carbon film was sputtered onto the etched surface of the samples. The carbon film (replica) was subsequently electrolytically extracted, using 8 % perchloric acid, at a bias of 10 V. 3 RESULTS Figure 2 shows the microstructure of material A after the special heat treatment. Light microscopy revealed a polyhedral grain, the presence of twins as well as the grain-size heterogeneity in material A. Two types of grain-boundary morphology were observed with light microscopy. Straight boundaries with a percentage share of approximately 38 % of the total of 254 observed grain boundaries were located in the structure of material A. Serrated boundaries, as the second type, covered approximately 39 % of the noticed boundaries. In some cases, the type of boundary could not be determined owing to the state of the boundary. As a reference sample, material B (Figure 3) was processed with the conventional heat treatment. In material B, we also observed the base microstructure I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT 722 Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Table 1: Nominal chemical composition of Alloy 263 Chemical-composition limits Weight % Ni Cr Co Mo C Al Ti Al+Ti Mn 263 Bal 19–21 19–21 5.6–6.1 0.04–0.08 0.60 max 1.90–2.40 2.40–2.80 0.60 max Weight % Si Fe Cu B Pb S Ag Bi 263 0.40 max 0.07 max 0.020 max 0.005 max 0.0020 0.007 max 0.005 max 0.0001 max Figure 1: Net efficiency development in the case of hard coal-fired power plants considering different structural alloys3 with polyhedral grains with the heterogeneity of the grain size and twins similar to those in material A. As expected, no serrated boundaries were observed in material B after the conventional heat treatment. The grain-boundary details taken with SEM indicate the presence of precipitates on the serrated and straight grain boundaries in material A as well as in material B (Figures 4 and 5). Results of the EDX analysis taken from material A (Figure 6) reveal the eventual presence of carbon-rich secondary particles, which could possibly indicate the presence of carbides at the grain boundaries. The occurrence of the other elements at the grain boun- daries was not significant. Comparable results were also noticed for material B. To confirm the presence of the secondary particles formed at the boundaries and to identify the chemical nature of these particles, ED and EDX using the TEM were taken for materials A and B, and the results are shown below. The authors of reference16 predicted and identified typical kinds of the secondary particles for this alloy, allowing us to expect mainly the presence of the MC, M6C and M23C6 carbides. Figures 7 to 9 summarize the identified phases on the replicas for material A. Electron diffraction spectra as well as the EDX analysis of the extracted precipitates I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 723 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: SEM, material B: a) straight grain-boundary triple point and b) detail of precipitatesFigure 3: Material B Figure 4: SEM, material A: a) serrated grain-boundary triple point and b) detail of precipitates Figure 2: Material A extracted at the grain boundaries confirmed the presence of carbides. The carbides contained minor elements such as Mo, Cr and Ni (Table 2). The M6C ((Co,Ni)3Mo3C) carbide was not confirmed by electron diffraction (confirmed only by the EDX anal- I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT 724 Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 8: a) Extracted small particles of irregular shapes at the grain boundary, material B and b) point-diffraction spectra of MC carbide Figure 6: EDX map, carbon-rich grain boundary, material A Figure 9: Typical EDX spectra of precipitates in material A: a) carbide MC and b) carbide M6C Figure 7: a) Detail of a precipitate at the grain boundary, material A and b) point-diffraction spectra of M23C6 carbide ysis); a slight difference between the lattice parameters of the M23C6 and M6C (aM23C6 = 0.1065 nm aM6C = 0.1085 nm) carbides makes the identification difficult. Table 2: Approximate chemical composition of carbides in Material A M23C6 MC M6C Avr. (w/%)  Avr. (w/%)  Avr. (w/%)  Mo 28.6 ±6.1 64.7 ±6.3 22.9 ±5.0 Ti 4.6 ±3.6 26.0 ±5.0 3.3 ±1.3 Cr 66.8 ±9.6 8.8 ±4.3 27.5 ±3.6 Ni 46.3 ±6.8 Details of the serrated boundaries and precipitates in Material A, observed with TEM on the foils, can be seen in Figures 10a to 10b. As shown in the figures, the asymmetry of the serration was observed on some parts of the serrated boundaries when the irregularity of the serration exhibited a high difference in  at the observed parts of the boundary (on the edges,  = 800–900 nm; in the middle of the boundary,  = 150–200 nm). On the serrated grain boundary documented in Figure 10b, secondary particles of different sizes were noticed. The shape of the particles copied the serration of the boundary. Using electron diffraction, the particles were identified as the M23C6 carbide and the matrix as the -phase (Figure 11). Electron diffraction also revealed the existence of the orientation relationship between the matrix and carbide: {200} || {200}M23C6. Figure 12a documents a straight grain boundary with discrete secondary particles in material A. Along the full length of the boundary, the particles of triangular or rectangular shapes were observed. As shown in Figure 12b, this kind of particles does not copy the shape of the boundary as in the case of the precipitates at the serrated boundaries. The size of the precipitates was approxi- mately 300 nm. The particles were identified, with elec- tron diffraction (Figure 13), as M23C6 carbide and the matrix as phase . However, in the case of the straight grain boundary, no orientation relationship between the matrix and the precipitate was spotted. The precipitates extracted at the grain boundary in material B as M23C6 carbide were identified on the replicas using electron diffraction, as documented below I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 725 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 12: Detail of straight grain boundary, material A, bright field; b) detail of precipitates at the grain boundary, material A, dark field; reflection (2 -2 4)M23C6 was used for the dark-field display Figure 10: Detail of serrated grain boundary, material A, bright field; b) detail of precipitates at the grain boundary, material A, dark field; reflection (2 0 0) M23C6 was used for the dark-field display Figure 11: Point diffraction spectra – phase  and M23C6 carbide extracted at the grain boundaries confirmed the presence of carbides. The carbides contained minor elements such as Mo, Cr and Ni (Table 2). The M6C ((Co,Ni)3Mo3C) carbide was not confirmed by electron diffraction (confirmed only by the EDX anal- I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT 724 Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 8: a) Extracted small particles of irregular shapes at the grain boundary, material B and b) point-diffraction spectra of MC carbide Figure 6: EDX map, carbon-rich grain boundary, material A Figure 9: Typical EDX spectra of precipitates in material A: a) carbide MC and b) carbide M6C Figure 7: a) Detail of a precipitate at the grain boundary, material A and b) point-diffraction spectra of M23C6 carbide ysis); a slight difference between the lattice parameters of the M23C6 and M6C (aM23C6 = 0.1065 nm aM6C = 0.1085 nm) carbides makes the identification difficult. Table 2: Approximate chemical composition of carbides in Material A M23C6 MC M6C Avr. (w/%)  Avr. (w/%)  Avr. (w/%)  Mo 28.6 ±6.1 64.7 ±6.3 22.9 ±5.0 Ti 4.6 ±3.6 26.0 ±5.0 3.3 ±1.3 Cr 66.8 ±9.6 8.8 ±4.3 27.5 ±3.6 Ni 46.3 ±6.8 Details of the serrated boundaries and precipitates in Material A, observed with TEM on the foils, can be seen in Figures 10a to 10b. As shown in the figures, the asymmetry of the serration was observed on some parts of the serrated boundaries when the irregularity of the serration exhibited a high difference in  at the observed parts of the boundary (on the edges,  = 800–900 nm; in the middle of the boundary,  = 150–200 nm). On the serrated grain boundary documented in Figure 10b, secondary particles of different sizes were noticed. The shape of the particles copied the serration of the boundary. Using electron diffraction, the particles were identified as the M23C6 carbide and the matrix as the -phase (Figure 11). Electron diffraction also revealed the existence of the orientation relationship between the matrix and carbide: {200} || {200}M23C6. Figure 12a documents a straight grain boundary with discrete secondary particles in material A. Along the full length of the boundary, the particles of triangular or rectangular shapes were observed. As shown in Figure 12b, this kind of particles does not copy the shape of the boundary as in the case of the precipitates at the serrated boundaries. The size of the precipitates was approxi- mately 300 nm. The particles were identified, with elec- tron diffraction (Figure 13), as M23C6 carbide and the matrix as phase . However, in the case of the straight grain boundary, no orientation relationship between the matrix and the precipitate was spotted. The precipitates extracted at the grain boundary in material B as M23C6 carbide were identified on the replicas using electron diffraction, as documented below I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 725 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 12: Detail of straight grain boundary, material A, bright field; b) detail of precipitates at the grain boundary, material A, dark field; reflection (2 -2 4)M23C6 was used for the dark-field display Figure 10: Detail of serrated grain boundary, material A, bright field; b) detail of precipitates at the grain boundary, material A, dark field; reflection (2 0 0) M23C6 was used for the dark-field display Figure 11: Point diffraction spectra – phase  and M23C6 carbide in Figure 14. Other precipitates were not identified, which does not exclude their presence in Material B. From the results of the EDX analysis (Figure 15, Table 3) of the observed precipitates in Material B, the presence of MC and M6C carbides is also possible. Figures 16a to 16b document details of straight boundaries and precipitates at the grain boundary in material B on the foils. It can be noticed that the secondary particles extracted at the grain boundaries have a polyhedral or rectangular shape, and none of the observed particles copies the shape of the boundary. The size of these precipitates is in range of 100–200 nm. I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT 726 Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 13: Point diffraction spectra – phase  and M23C6 carbide Figure 16: Detail of straight grain boundary, material B, bright field; b) detail of precipitates at the grain boundary, material B, dark field; reflection (5 -1 -3)M23C6 was used for the dark-field display Figure 14: Extracted small particle of an irregular shape at the grain boundary, material A; b) point diffraction spectra of M23C6 carbide Figure 15: Typical EDX spectra of precipitates in material B: a) M23C6 carbide, b) M6C carbide Table 3: Approximate chemical composition of carbides in material B M23C6 MC M6C Avr. (w/%)  Avr. (w/%)  Avr. (w/%)  Mo 27.3 ±6.4 62.9 ±5.6 24.6 ±3.2 Ti 4.3 ±4.9 25.0 ±4.7 4.1 ±1.8 Cr 61.7 ±9.4 10.5 ±3.5 26.0 ±3.1 Ni 42.6 ±5.1 From the diffraction spectra (Figure 17), secondary particles at the grain boundaries were identified as M23C6 carbide. No relationship between the matrix and preci- pitate was observed in all the diffraction spectra in material B. 4 CONCLUSION To study the serration and precipitates at the grain boundaries and identify secondary particles in Alloy 263, both SEM and TEM were used in the research. To form serrated grain boundaries, material A was heat treated under special conditions, based on the studies from the literature. Material A was compared with material B, which underwent the conventional heat treatment based on the material list for Alloy 263. The experimental results can be summarized as follows: • Serrated and straight grain boundaries were observed in material A. Approximately 39 % of the observed boundaries were serrated. In material B, only straight grain boundaries were observed. • In both cases, precipitates at the grain boundaries were identified as carbide M23C6 using electron diffraction. In material A, carbide MC was also ob- served. • The EDX analysis revealed three types of carbides, based on their chemical compositions, in both ma- terials: MC, M6C and M23C6. • Electron diffraction also revealed the orientation relationship between the matrix and the precipitate at the serrated grain boundaries: {200} {200}M23C6. No orientation relationship was observed at the straight grain boundaries. Acknowledgments The authors wish to thank the financial support of the Slovak Republic Scientific Grant Agency (VEGA) within Grant No. 1/0402/13. 5 REFERENCES 1 P. J. Maziasz, I. G. Wright, J. P. Shingledecker, T. B. Gibbons, R. R. Romanosky, Defining the materials issues and research needs for ultra-supercritical steam turbines, Proc. 4th Inter. Conf. Advances in Materials Technology for Fossil Power Plants, Hilton Head Island, USA, 2005, 602–622 2 R. J. Campbell, Increasing the Efficiency of Existing Coal-Fired Power Plants, https://fas.org/sgp/crs/misc/R43343.pdf, 14.06.2017 3 G. Stein-Brzozowska, D. M. Flórez, J. Maier, G. Scheffknecht, Nickel-base superalloys for ultra-supercritical coal-fired power plants: Fireside corrosion. Laboratory studies and power plant expo- sures, Fuel, 108, (2013), 521–533, doi:10.1016/j.fuel.2012.11.081 4 A. K. Koul, G. H. Gessinger, On the mechanism of serrated grain boundary formation in Ni-based superalloys, Acta Metallurgica, 31, (1983) 7, 1061–1069, doi:10.1016/0001-6160(83)90202-X 5 H. L. Danflou, M. Marty, A. Walder, Formation of serrated grain boundaries and their effect on the mechanical properties in a P/M nickel base superalloy, Proc. 7th Inter. Symp. on Superalloys, Seven Springs Mountain, USA, 1992, 63–72, doi:10.7449/1992/ Super- alloys_1992_63_72 6 R. J. Mitchell, H. Y. Li, Z. W. Huang, On the formation of serrated grain boundaries and fan type structures in an advanced polycry- stalline nickel-base superalloy, Journal of Materials Processing Tech- nology, 209 (2009) 2, 1011–1017, doi:10.1016/j.jmatprotec.2008. 03.008 7 C. L. Qiu, P. Andrews, On the formation of irregular-shaped gamma prime and serrated grain boundaries in a nickel-based superalloy during continuous cooling, Materials Characterization, 76 (2013), 28–34, doi:10.1016/j.matchar.2012.11.012 8 A. C. Yeh, K. W. Lu, C. M. Kuo, H. Y. Bor, C. N. Wei, Effect of Serrated Grain Boundaries on the Creep Property of Inconel 718 Superalloy, Materials Science and Engineering A, 530 (2011), 525–529, doi:10.1016/j.msea.2011.10.014 9 D. H. Ping, Y. F. Gu, C. Y. Cui, H. Harada, Grain boundary segre- gation in a Ni–Fe-based (Alloy 718) superalloy, Materials Science and Engineering A, 456 (2007) 1–2, 99–102, doi:10.1016/j.msea. 2007.01.090 10 L. Jiang, R. Hu, H. Kou, J. Li, G. Bai, H. Fu, The effect of M23C6 carbides on the formation of grain boundary serrations in a wrought Ni-based superalloy, Materials Science and Engineering A, 536 (2012), 37–44, doi:10.1016/j.msea.2011.11.060 11 Y. S. Lim, D. J. Kim, S. S. Hwang, H. P. Kim, S. W. Kim, M23C6 precipitation behavior and grain boundary serration in Ni-based Alloy 690, Materials Characterization, 96 (2014), 28–39, doi:10.1016/j.matchar.2014.07.008 12 H. U. Hong, I. S. Kim, B. G. Choi, M. Y. Kim, C. Y. Jo, The effect of grain boundary serration on creep resistance in a wrought nickel-based superalloy, Materials Science and Engineering A, 517 (2009) 1–2, 125–131, doi:10.1016/j.msea.2009.03.071 13 H. U. Hong, I. S. Kim, B. G. Choi, Y. S. Yoo, C. Y. Jo, On the role of grain boundary serration in simulated weld heat-affected zone liquation of a wrought nickel-based superalloy, Metallurgical and Materials Transactions A, 43 (2012) 1, 173–181, doi:10.1007/ s11661-011-0837-2 14 H. U. Hong, I. S. Kim, B. G. Choi, Y. S. Yoo, C. Y. Jo, On the Mechanism of Serrated Grain Boundary Formation in Ni-Based Superalloys with Low '? Volume Fraction, Proc. 12th Inter. Symp. on Superalloys, Seven Springs Mountain, USA, 2012, 53–61, doi:10.1002/9781118516430.ch6 15 H. U. Hong, F. H. Latief, T. Blanc, I. S. Kim, B. G. Choi, C. Y. Jo, J. H. Lee, Influence of chromium content on microstructure and grain boundary serration formation in a ternary Ni-xCr-0.1C model alloy, Materials Chemistry and Physics, 148 (2014) 3, 1194–1201, doi:10.1016/j.matchemphys.2014.09.047 16 J. C. Zhao, V. Ravikumar, A. M. Beltran, Phase Precipitation and Phase Stability in Nimonic 263, Metallurgical and Materials Transactions A, 32 (2001) 6, 1271–1282, doi:10.1007/s11661- 001-0217-4 I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 727 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 17: Point diffraction spectra – phase  and M23C6 carbide in Figure 14. Other precipitates were not identified, which does not exclude their presence in Material B. From the results of the EDX analysis (Figure 15, Table 3) of the observed precipitates in Material B, the presence of MC and M6C carbides is also possible. Figures 16a to 16b document details of straight boundaries and precipitates at the grain boundary in material B on the foils. It can be noticed that the secondary particles extracted at the grain boundaries have a polyhedral or rectangular shape, and none of the observed particles copies the shape of the boundary. The size of these precipitates is in range of 100–200 nm. I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT 726 Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 13: Point diffraction spectra – phase  and M23C6 carbide Figure 16: Detail of straight grain boundary, material B, bright field; b) detail of precipitates at the grain boundary, material B, dark field; reflection (5 -1 -3)M23C6 was used for the dark-field display Figure 14: Extracted small particle of an irregular shape at the grain boundary, material A; b) point diffraction spectra of M23C6 carbide Figure 15: Typical EDX spectra of precipitates in material B: a) M23C6 carbide, b) M6C carbide Table 3: Approximate chemical composition of carbides in material B M23C6 MC M6C Avr. (w/%)  Avr. (w/%)  Avr. (w/%)  Mo 27.3 ±6.4 62.9 ±5.6 24.6 ±3.2 Ti 4.3 ±4.9 25.0 ±4.7 4.1 ±1.8 Cr 61.7 ±9.4 10.5 ±3.5 26.0 ±3.1 Ni 42.6 ±5.1 From the diffraction spectra (Figure 17), secondary particles at the grain boundaries were identified as M23C6 carbide. No relationship between the matrix and preci- pitate was observed in all the diffraction spectra in material B. 4 CONCLUSION To study the serration and precipitates at the grain boundaries and identify secondary particles in Alloy 263, both SEM and TEM were used in the research. To form serrated grain boundaries, material A was heat treated under special conditions, based on the studies from the literature. Material A was compared with material B, which underwent the conventional heat treatment based on the material list for Alloy 263. The experimental results can be summarized as follows: • Serrated and straight grain boundaries were observed in material A. Approximately 39 % of the observed boundaries were serrated. In material B, only straight grain boundaries were observed. • In both cases, precipitates at the grain boundaries were identified as carbide M23C6 using electron diffraction. In material A, carbide MC was also ob- served. • The EDX analysis revealed three types of carbides, based on their chemical compositions, in both ma- terials: MC, M6C and M23C6. • Electron diffraction also revealed the orientation relationship between the matrix and the precipitate at the serrated grain boundaries: {200} {200}M23C6. No orientation relationship was observed at the straight grain boundaries. Acknowledgments The authors wish to thank the financial support of the Slovak Republic Scientific Grant Agency (VEGA) within Grant No. 1/0402/13. 5 REFERENCES 1 P. J. Maziasz, I. G. Wright, J. P. Shingledecker, T. B. Gibbons, R. R. Romanosky, Defining the materials issues and research needs for ultra-supercritical steam turbines, Proc. 4th Inter. Conf. Advances in Materials Technology for Fossil Power Plants, Hilton Head Island, USA, 2005, 602–622 2 R. J. Campbell, Increasing the Efficiency of Existing Coal-Fired Power Plants, https://fas.org/sgp/crs/misc/R43343.pdf, 14.06.2017 3 G. Stein-Brzozowska, D. M. Flórez, J. Maier, G. Scheffknecht, Nickel-base superalloys for ultra-supercritical coal-fired power plants: Fireside corrosion. Laboratory studies and power plant expo- sures, Fuel, 108, (2013), 521–533, doi:10.1016/j.fuel.2012.11.081 4 A. K. Koul, G. H. Gessinger, On the mechanism of serrated grain boundary formation in Ni-based superalloys, Acta Metallurgica, 31, (1983) 7, 1061–1069, doi:10.1016/0001-6160(83)90202-X 5 H. L. Danflou, M. Marty, A. Walder, Formation of serrated grain boundaries and their effect on the mechanical properties in a P/M nickel base superalloy, Proc. 7th Inter. Symp. on Superalloys, Seven Springs Mountain, USA, 1992, 63–72, doi:10.7449/1992/ Super- alloys_1992_63_72 6 R. J. Mitchell, H. Y. Li, Z. W. Huang, On the formation of serrated grain boundaries and fan type structures in an advanced polycry- stalline nickel-base superalloy, Journal of Materials Processing Tech- nology, 209 (2009) 2, 1011–1017, doi:10.1016/j.jmatprotec.2008. 03.008 7 C. L. Qiu, P. Andrews, On the formation of irregular-shaped gamma prime and serrated grain boundaries in a nickel-based superalloy during continuous cooling, Materials Characterization, 76 (2013), 28–34, doi:10.1016/j.matchar.2012.11.012 8 A. C. Yeh, K. W. Lu, C. M. Kuo, H. Y. Bor, C. N. Wei, Effect of Serrated Grain Boundaries on the Creep Property of Inconel 718 Superalloy, Materials Science and Engineering A, 530 (2011), 525–529, doi:10.1016/j.msea.2011.10.014 9 D. H. Ping, Y. F. Gu, C. Y. Cui, H. Harada, Grain boundary segre- gation in a Ni–Fe-based (Alloy 718) superalloy, Materials Science and Engineering A, 456 (2007) 1–2, 99–102, doi:10.1016/j.msea. 2007.01.090 10 L. Jiang, R. Hu, H. Kou, J. Li, G. Bai, H. Fu, The effect of M23C6 carbides on the formation of grain boundary serrations in a wrought Ni-based superalloy, Materials Science and Engineering A, 536 (2012), 37–44, doi:10.1016/j.msea.2011.11.060 11 Y. S. Lim, D. J. Kim, S. S. Hwang, H. P. Kim, S. W. Kim, M23C6 precipitation behavior and grain boundary serration in Ni-based Alloy 690, Materials Characterization, 96 (2014), 28–39, doi:10.1016/j.matchar.2014.07.008 12 H. U. Hong, I. S. Kim, B. G. Choi, M. Y. Kim, C. Y. Jo, The effect of grain boundary serration on creep resistance in a wrought nickel-based superalloy, Materials Science and Engineering A, 517 (2009) 1–2, 125–131, doi:10.1016/j.msea.2009.03.071 13 H. U. Hong, I. S. Kim, B. G. Choi, Y. S. Yoo, C. Y. Jo, On the role of grain boundary serration in simulated weld heat-affected zone liquation of a wrought nickel-based superalloy, Metallurgical and Materials Transactions A, 43 (2012) 1, 173–181, doi:10.1007/ s11661-011-0837-2 14 H. U. Hong, I. S. Kim, B. G. Choi, Y. S. Yoo, C. Y. Jo, On the Mechanism of Serrated Grain Boundary Formation in Ni-Based Superalloys with Low '? Volume Fraction, Proc. 12th Inter. Symp. on Superalloys, Seven Springs Mountain, USA, 2012, 53–61, doi:10.1002/9781118516430.ch6 15 H. U. Hong, F. H. Latief, T. Blanc, I. S. Kim, B. G. Choi, C. Y. Jo, J. H. Lee, Influence of chromium content on microstructure and grain boundary serration formation in a ternary Ni-xCr-0.1C model alloy, Materials Chemistry and Physics, 148 (2014) 3, 1194–1201, doi:10.1016/j.matchemphys.2014.09.047 16 J. C. Zhao, V. Ravikumar, A. M. Beltran, Phase Precipitation and Phase Stability in Nimonic 263, Metallurgical and Materials Transactions A, 32 (2001) 6, 1271–1282, doi:10.1007/s11661- 001-0217-4 I. SLATKOVSKÝ et al.: INVESTIGATION OF GRAIN BOUNDARIES IN ALLOY 263 AFTER SPECIAL HEAT TREATMENT Materiali in tehnologije / Materials and technology 51 (2017) 5, 721–727 727 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 17: Point diffraction spectra – phase  and M23C6 carbide J. PTA^INOVÁ et al.: FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT ... 729–733 FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT FOR 17 H AND DOUBLE TEMPERING LOMNA @ILAVOST LEDEBURITNEGA JEKLA VANADIS 6 PO TOPLOTNI OBDELAVI S 17-URNIM PODHLAJEVANJEM IN DVOJNIM POPU[^ANJEM Jana Pta~inová1, Peter Jur~i1, Ivo Dlouhý2 1Institute of Materials Science, Faculty of Materials Science and Technology Trnava, Jána Bottu 25, 917 24 Trnava, Slovakia 2Institute of Physics of Materials, Academy of Sciences of the Czech Republic, Zizkova 22, 61662 Brno, Czech Republic jana.ptacinova@stuba.sk Prejem rokopisa – received: 2016-06-20; sprejem za objavo – accepted for publication: 2017-03-01 doi:10.17222/mit.2016.118 The influence of sub-zero treatment on the fracture toughness of Cr-V ledeburitic steel Vanadis 6 was examined, in comparison with the same material processed without the sub-zero period. The microstructure of the material consists of the matrix and several types of carbides – eutectic carbides (ECs), secondary carbides (SCs) and small globular carbides (SGCs). Small amounts of retained austenite were also present in the microstructure, but only in the untempered or low-temperature tempered steel. Sub-zero treatment increases the amount of small globular carbides. On the other hand, tempering results in a decrease in the population density of these particles. The fracture toughness of conventionally heat-treated steel firstly increases with the tempering, but then it decreases rapidly when the steel is tempered at the temperature of secondary hardening. In the case of the sub-zero treated material, the fracture toughness is correspondingly lower when the material is tempered at low temperatures, but it becomes slightly higher in the temperature range normally used for secondary hardening. Generally, one can say that the fracture toughness follows well the values of the hardness of the material, except in the narrow temperature range in the case of the sub-zero treated steel, where a "window" for a simultaneous enhancement of hardness and toughness exists. Keywords: ledeburitic steel, sub-zero treatment, fracture toughness, carbides, fracture surface Avtorji so preu~evali vpliv standardne toplotne obdelave v kombinaciji s podhlajevanjem na lomno `ilavost Cr-V ledeburitnega jekla Vanadis 6 v primerjavi z enakim materialom brez podhlajevanja. Mikrostruktura materiala sestoji iz matrice in ve~ vrst karbidov – evtekti~nih karbidov (angl. ECs), sekundarnih karbidov (angl. SCs), in manj{ih globularnih karbidov (angl. SGCs). Majhne koli~ine zaostalega avstenita so bile prisotne v mikrostrukturi, vendar le v nepopu{~enem ali v nizkotemperaturno- popu{~enenem jeklu. Podhlajevanje pove~uje koli~ino manj{ih globularnih karbidov. Po drugi strani pa se s popu{~anjem zmanj{uje populacijska gostota teh delcev. Lomna `ilavost konvencionalno toplotno obdelanega jekla se najprej pove~uje s povi{evanjem temperature popu{~anja vendar se pri~ne hitro zmanj{evati (zni`evati), ko je prekora~ena temperatura sekun- darnega utrjevanja. Pri toplotni obdelavi materiala s podhlajevanjem, se lomna `ilavost posledi~no zni`uje, ko je le-to popu{~eno na ni`jih temperaturah, vendar se rahlo pove~uje v obmo~ju temperature, ki se navadno uporablja za sekundarno utrjevanje. V splo{nem je mo`no re~i, da lomna `ilavost sledi vrednostim trdote materiala, razen v primeru toplotno obdelanega jekla v kombinaciji s podhlajevanjem, kjer v zelo ozkem temperaturnem intervalu obstaja "temperaturno okno". Tam dose`emo isto~asno pove~anje `ilavosti in trdote. Klju~ne besede: ledeburitno jeklo, podhlajevanje, lomna `ilavost, karbidi, lomna povr{ina 1 INTRODUCTION In modern tooling, it is necessary that the tool mate- rials have a good wear performance and high hardness, accompanied with at least acceptable toughness. Other- wise the tools might fail very early, before any wear damage can occur. The toughness of steels quantifies their resistance to the initiation of brittle fracture under given conditions. For real tools made of Cr- and Cr-V ledeburitic steels, the toughness determines their capability to resist either the chipping or a total failure. The toughness, being expressed by the three-point bending strength, b, of brittle ledeburitic steels is the highest when the material is soft-annealed and it decreases as a consequence of the application of heat treatment. However, the higher the austenitizing temperature (TA), the lower is the toughness and, at a constant TA, the tempering first results in an in- crease in the toughness, which is followed by its decrease at the maximum secondary hardening.1 Hence, the application of proper heat treatment is a compromise between the requirements for the maximum hardness and at least acceptable toughness. There is a shortage of literature pertaining to the effect of the sub-zero treatment (SZT) on the toughness. D. N. Collins and J. Dormer reported that the toughness of the AISI D2 steel decreased with an application of the SZT up to the temperature of –70 °C, which was followed by a moderate increase in the toughness when a lower processing temperature was used for the SZT.2,3 It should be noticed here that the samples were low-tem- perature tempered at 200 °C. Materiali in tehnologije / Materials and technology 51 (2017) 5, 729–733 729 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 620.178.2:669.15-194.58:691.714:62-712 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)729(2017) J. PTA^INOVÁ et al.: FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT ... 729–733 FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT FOR 17 H AND DOUBLE TEMPERING LOMNA @ILAVOST LEDEBURITNEGA JEKLA VANADIS 6 PO TOPLOTNI OBDELAVI S 17-URNIM PODHLAJEVANJEM IN DVOJNIM POPU[^ANJEM Jana Pta~inová1, Peter Jur~i1, Ivo Dlouhý2 1Institute of Materials Science, Faculty of Materials Science and Technology Trnava, Jána Bottu 25, 917 24 Trnava, Slovakia 2Institute of Physics of Materials, Academy of Sciences of the Czech Republic, Zizkova 22, 61662 Brno, Czech Republic jana.ptacinova@stuba.sk Prejem rokopisa – received: 2016-06-20; sprejem za objavo – accepted for publication: 2017-03-01 doi:10.17222/mit.2016.118 The influence of sub-zero treatment on the fracture toughness of Cr-V ledeburitic steel Vanadis 6 was examined, in comparison with the same material processed without the sub-zero period. The microstructure of the material consists of the matrix and several types of carbides – eutectic carbides (ECs), secondary carbides (SCs) and small globular carbides (SGCs). Small amounts of retained austenite were also present in the microstructure, but only in the untempered or low-temperature tempered steel. Sub-zero treatment increases the amount of small globular carbides. On the other hand, tempering results in a decrease in the population density of these particles. The fracture toughness of conventionally heat-treated steel firstly increases with the tempering, but then it decreases rapidly when the steel is tempered at the temperature of secondary hardening. In the case of the sub-zero treated material, the fracture toughness is correspondingly lower when the material is tempered at low temperatures, but it becomes slightly higher in the temperature range normally used for secondary hardening. Generally, one can say that the fracture toughness follows well the values of the hardness of the material, except in the narrow temperature range in the case of the sub-zero treated steel, where a "window" for a simultaneous enhancement of hardness and toughness exists. Keywords: ledeburitic steel, sub-zero treatment, fracture toughness, carbides, fracture surface Avtorji so preu~evali vpliv standardne toplotne obdelave v kombinaciji s podhlajevanjem na lomno `ilavost Cr-V ledeburitnega jekla Vanadis 6 v primerjavi z enakim materialom brez podhlajevanja. Mikrostruktura materiala sestoji iz matrice in ve~ vrst karbidov – evtekti~nih karbidov (angl. ECs), sekundarnih karbidov (angl. SCs), in manj{ih globularnih karbidov (angl. SGCs). Majhne koli~ine zaostalega avstenita so bile prisotne v mikrostrukturi, vendar le v nepopu{~enem ali v nizkotemperaturno- popu{~enenem jeklu. Podhlajevanje pove~uje koli~ino manj{ih globularnih karbidov. Po drugi strani pa se s popu{~anjem zmanj{uje populacijska gostota teh delcev. Lomna `ilavost konvencionalno toplotno obdelanega jekla se najprej pove~uje s povi{evanjem temperature popu{~anja vendar se pri~ne hitro zmanj{evati (zni`evati), ko je prekora~ena temperatura sekun- darnega utrjevanja. Pri toplotni obdelavi materiala s podhlajevanjem, se lomna `ilavost posledi~no zni`uje, ko je le-to popu{~eno na ni`jih temperaturah, vendar se rahlo pove~uje v obmo~ju temperature, ki se navadno uporablja za sekundarno utrjevanje. V splo{nem je mo`no re~i, da lomna `ilavost sledi vrednostim trdote materiala, razen v primeru toplotno obdelanega jekla v kombinaciji s podhlajevanjem, kjer v zelo ozkem temperaturnem intervalu obstaja "temperaturno okno". Tam dose`emo isto~asno pove~anje `ilavosti in trdote. Klju~ne besede: ledeburitno jeklo, podhlajevanje, lomna `ilavost, karbidi, lomna povr{ina 1 INTRODUCTION In modern tooling, it is necessary that the tool mate- rials have a good wear performance and high hardness, accompanied with at least acceptable toughness. Other- wise the tools might fail very early, before any wear damage can occur. The toughness of steels quantifies their resistance to the initiation of brittle fracture under given conditions. For real tools made of Cr- and Cr-V ledeburitic steels, the toughness determines their capability to resist either the chipping or a total failure. The toughness, being expressed by the three-point bending strength, b, of brittle ledeburitic steels is the highest when the material is soft-annealed and it decreases as a consequence of the application of heat treatment. However, the higher the austenitizing temperature (TA), the lower is the toughness and, at a constant TA, the tempering first results in an in- crease in the toughness, which is followed by its decrease at the maximum secondary hardening.1 Hence, the application of proper heat treatment is a compromise between the requirements for the maximum hardness and at least acceptable toughness. There is a shortage of literature pertaining to the effect of the sub-zero treatment (SZT) on the toughness. D. N. Collins and J. Dormer reported that the toughness of the AISI D2 steel decreased with an application of the SZT up to the temperature of –70 °C, which was followed by a moderate increase in the toughness when a lower processing temperature was used for the SZT.2,3 It should be noticed here that the samples were low-tem- perature tempered at 200 °C. Materiali in tehnologije / Materials and technology 51 (2017) 5, 729–733 729 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 620.178.2:669.15-194.58:691.714:62-712 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)729(2017) The fracture toughness (KIC) (the resistance against the crack propagation) of the ledeburitic steels can be quantified using either pre-cracked (bend) specimens or chevron-notch technique.4–6 The fracture toughness of ledeburitic steels is very low, whereas the KIC values usually follow the three-point bending strength. The effect of the SZT on the fracture toughness was reported by several groups of investigators. D. Das et al., for instance, pointed out that the SZT carried out at tem- peratures of (–75, –125 and –196) °C led to a decrease in the fracture toughness compared to conventionally heat-treated samples when low-temperature tempered at 210 °C for 2 h.6 The variations in the fracture toughness were attributed to the reduction of the retained-austenite amount and to the increase in the population density of SGCs in the microstructure after the SZT. In our recent paper, it was demonstrated that the effect of the SZT on the toughness is marginal, but the effect of this process- ing on KIC is rather positive when the material is tempered at the temperature normally used for secondary hardening.7 The main goal of the current investigations is to characterize the variations in the toughness and fracture toughness of powder-metallurgy (PM) ledeburitic steel Vanadis 6, as a result of different heat-treatment sche- dules (austenitizing, sub-zero treatment, tempering) and to relate them to the microstructural alterations, due to these treatments, like the reduction of the retained-auste- nite amount, increase in the carbide count and others.8–10 2 MATERIAL AND EXPERIMENTAL METHODS 2.1 Material and processing The experimental material was tool steel Vanadis 6 with nominally 2.1 % C, 1.0 % Si, 0.4 % Mn, 6.8 % Cr, 1.5 % Mo, 5.4 % V and Fe as balance, made with PM. The conventional heat treatment (CHT) consisted of the following steps: heating up the material to the desired TA (1050 °C) in a vacuum furnace, holding it at the tem- perature for 30 min and nitrogen gas quenching (5 bar). One set of samples was processed without the inclusion of the SZT between quenching and tempering, while the other samples were subjected to the SZT carried out at the temperature of -196 °C for 17 h. Double tempering (2 h + 2 h) was performed at temperatures in the range of 100–600 °C. 2.2 Experimental methods Metallographic samples were prepared in the stan- dard way and etched with a picric-acid reagent for the SEM observation. The microstructure was recorded using a JEOL JSM 7600 F device equipped with an EDS detector (Oxford Instruments), at an acceleration voltage of 15 kV. Details of the categorization of the carbides have been published recently.10,11 Macro-hardness measurements were completed using the Vickers (HV10) method. Five measurements were made on the metallo- graphic specimens processed with any combination of heat-treatment parameters, and both the mean values and the standard deviations were then calculated. For the fracture-toughness determination, pre-cracked specimens predetermined for bending, with dimensions of 10 mm × 10 mm × 55 mm were used. For the pre-cracking of the samples, four bending fixtures and a resonance freq- uency machine (Cracktronic 8024) were used, which allowed the fatigue-crack initiation and further propa- gation through the actual frequency of the cycling to be controlled. In addition, the crack development was checked on both sides of a sample using digital long-distance micro- scopic methods. Both the pre-crack preparation and bend tests were carried out at room temperature according to the ISO12137 standard.12 For the test, an Instron 8862 machine was used and a loading rate of 0.1 mm/min was applied. Specimen deflection was measured by means of an inductive transducer integrated directly into the loading axis. In total, five samples were tested for each investigated condition. Fracture-surface morphology was investigated with a scanning electron microscope JEOL JSM 7600F with an EDS detector (Oxford Instruments). The topography of fracture surfaces was studied with a confocal laser scanning microscope Zeiss LSM 700. Three-dimensional topographical resolution was achieved using the ZEN 2009 software. 3 RESULTS AND DISCUSSION 3.1 Microstructure SEM micrographs, Figure 1, show an example of the microstructure evolution of the material after the sub-zero treatment in liquid nitrogen (for 17 h) and tempering at different temperatures. The microstructure after the sub-zero treatment consists of a matrix made of martensite, a small amount of retained austenite and carbides (Figure 1a). As expected, the character of the matrix microstructure changes with increasing tempering temperature (Figures 1b to 1f). Due to the tempering, the martensite becomes more sensitive to the etching agent (the so-called tempered martensite). This is a com- monly known fact and it is related to the precipitation of carbides during tempering. Retained austenite is trans- formed into martensite. The portion of carbides is also changed due to the tempering. The volume fractions of both eutectic carbides and secondary carbides are invariant over the range of the heat-treatment parameters used in the experiments. The population density of the small globular carbides increased with the application of the sub-zero treatment, but rapidly decreased with the in- creasing temperature of the tempering (Figure 2). 3.2 Hardness characteristics The bulk hardness of the non-SZT and SZT Vanadis 6 steel, as a function of the tempering temperature, is J. PTA^INOVÁ et al.: FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT ... 730 Materiali in tehnologije / Materials and technology 51 (2017) 5, 729–733 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS shown in Table 1. The as-quenched hardness of the conventionally heat-treated steel that was not tempered was 838±4.33 HV10. The hardness of the SZT steel soaked in liquid nitrogen for 17 h and not tempered was correspondingly higher, e.g., 917±7.58 HV10. These results show that the as-quenched bulk hardness of the Vanadis 6 steel was improved due to the sub-zero treat- ment. The hardness of all the samples then decreases with the increasing tempering temperature. Here it is interesting that the hardness of the SZT steel remains higher up to the tempering temperature of 450 °C and then it drops more intensely than that of the convention- ally quenched samples. As a result, the hardness of the specimens tempered in the range of the temperatures commonly used for secondary hardening is lower for the SZT steel than for the steel achieved after conventional quenching and tempering. 3.3 Evaluation of the SZT effect on the fracture tough- ness The fracture toughness vs. heat treatment schedules are presented in Figure 3a (CHT) and Figure 3b (SZT). In the case of the CHT, the fracture toughness of the untempered material was measured to be 16.4 MPa m1/2 and it firstly increased, due to the tempering, to 19.6 MPa m1/2 with a subsequent slight decrease; however, the decrease in the fracture toughness markedly accelerated, at the temperature normally used for secondary harden- ing, to a value of 14.8 MPa m1/2. For the material subjected to the SZT in liquid nitrogen for 17 h, the fracture toughness before tempering was 13.3 MPa m1/2 (i.e., lower than that of the CHT steel). Then, the fracture toughness increased with the tempering temperature to more than 15 MPa m1/2 and just slightly decreased at a tempering temperature of 450 °C. It is interesting that the KIC values were higher for the SZT steel than for the CHT steel at the temperature of secondary hardening (530 °C). Table 1: Hardness of the Vanadis 6 ledeburitic steel Tempering of samples CHT SZT HV10 ± HV10 ± untempered 838 4.3 917 7.6 tempered at 170 °C 784 3.7 870 5.4 tempered at 330 °C 736 7 851 15.7 tempered at 450 °C 724 3 816 7.8 tempered at 530 °C 758 5.9 701 3 The behaviour of fracture toughness vs. heat treat- ment schedules can be classified as logical because one can expect higher KIC at lower hardness and lower KIC at elevated hardness. In other words, the application of the SZT decreases KIC when the steel is low-temperature tempered. This fact should be considered by toolmakers and users of tools in all the cases when they expect an increase in the wear resistance due to a high hardness resulting from the sub-zero treatment. A very interesting fact was found for the material tempered at the temperature of secondary hardening – KIC was higher for the SZT material. This is in good agreement with the recently published results where a similar tendency was found for the same SZT steel treated for 4 h and 10 h.7 One can say that this kind of KIC behaviour can be expected because of the lower hardness of the SZT steel when tempered in this temperature range; however, in the mentioned paper, it was also found that the increase of KIC is accompanied with a better wear performance of the material, due to the presence of a higher amount of carbides. The enhanced number of carbides, compared to the material after the conventional quenching, was also identified in the current work. Hence, one can expect that a sub-zero treated material would have a better wear performance, too. These results are also interesting from J. PTA^INOVÁ et al.: FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 729–733 731 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Amount of carbides in the Vanadis 6 steel after conventional quenching, SZT for 17 h and SZT and subsequent tempering at different temperatures Figure 1: SEM micrographs showing the microstructure of Vanadis 6 ledeburitic steel: a) after quenching, subsequent SZT and b) tempering at 170 °C, c) 330 °C, d) 450 °C, e) 530 °C, f) 600 °C The fracture toughness (KIC) (the resistance against the crack propagation) of the ledeburitic steels can be quantified using either pre-cracked (bend) specimens or chevron-notch technique.4–6 The fracture toughness of ledeburitic steels is very low, whereas the KIC values usually follow the three-point bending strength. The effect of the SZT on the fracture toughness was reported by several groups of investigators. D. Das et al., for instance, pointed out that the SZT carried out at tem- peratures of (–75, –125 and –196) °C led to a decrease in the fracture toughness compared to conventionally heat-treated samples when low-temperature tempered at 210 °C for 2 h.6 The variations in the fracture toughness were attributed to the reduction of the retained-austenite amount and to the increase in the population density of SGCs in the microstructure after the SZT. In our recent paper, it was demonstrated that the effect of the SZT on the toughness is marginal, but the effect of this process- ing on KIC is rather positive when the material is tempered at the temperature normally used for secondary hardening.7 The main goal of the current investigations is to characterize the variations in the toughness and fracture toughness of powder-metallurgy (PM) ledeburitic steel Vanadis 6, as a result of different heat-treatment sche- dules (austenitizing, sub-zero treatment, tempering) and to relate them to the microstructural alterations, due to these treatments, like the reduction of the retained-auste- nite amount, increase in the carbide count and others.8–10 2 MATERIAL AND EXPERIMENTAL METHODS 2.1 Material and processing The experimental material was tool steel Vanadis 6 with nominally 2.1 % C, 1.0 % Si, 0.4 % Mn, 6.8 % Cr, 1.5 % Mo, 5.4 % V and Fe as balance, made with PM. The conventional heat treatment (CHT) consisted of the following steps: heating up the material to the desired TA (1050 °C) in a vacuum furnace, holding it at the tem- perature for 30 min and nitrogen gas quenching (5 bar). One set of samples was processed without the inclusion of the SZT between quenching and tempering, while the other samples were subjected to the SZT carried out at the temperature of -196 °C for 17 h. Double tempering (2 h + 2 h) was performed at temperatures in the range of 100–600 °C. 2.2 Experimental methods Metallographic samples were prepared in the stan- dard way and etched with a picric-acid reagent for the SEM observation. The microstructure was recorded using a JEOL JSM 7600 F device equipped with an EDS detector (Oxford Instruments), at an acceleration voltage of 15 kV. Details of the categorization of the carbides have been published recently.10,11 Macro-hardness measurements were completed using the Vickers (HV10) method. Five measurements were made on the metallo- graphic specimens processed with any combination of heat-treatment parameters, and both the mean values and the standard deviations were then calculated. For the fracture-toughness determination, pre-cracked specimens predetermined for bending, with dimensions of 10 mm × 10 mm × 55 mm were used. For the pre-cracking of the samples, four bending fixtures and a resonance freq- uency machine (Cracktronic 8024) were used, which allowed the fatigue-crack initiation and further propa- gation through the actual frequency of the cycling to be controlled. In addition, the crack development was checked on both sides of a sample using digital long-distance micro- scopic methods. Both the pre-crack preparation and bend tests were carried out at room temperature according to the ISO12137 standard.12 For the test, an Instron 8862 machine was used and a loading rate of 0.1 mm/min was applied. Specimen deflection was measured by means of an inductive transducer integrated directly into the loading axis. In total, five samples were tested for each investigated condition. Fracture-surface morphology was investigated with a scanning electron microscope JEOL JSM 7600F with an EDS detector (Oxford Instruments). The topography of fracture surfaces was studied with a confocal laser scanning microscope Zeiss LSM 700. Three-dimensional topographical resolution was achieved using the ZEN 2009 software. 3 RESULTS AND DISCUSSION 3.1 Microstructure SEM micrographs, Figure 1, show an example of the microstructure evolution of the material after the sub-zero treatment in liquid nitrogen (for 17 h) and tempering at different temperatures. The microstructure after the sub-zero treatment consists of a matrix made of martensite, a small amount of retained austenite and carbides (Figure 1a). As expected, the character of the matrix microstructure changes with increasing tempering temperature (Figures 1b to 1f). Due to the tempering, the martensite becomes more sensitive to the etching agent (the so-called tempered martensite). This is a com- monly known fact and it is related to the precipitation of carbides during tempering. Retained austenite is trans- formed into martensite. The portion of carbides is also changed due to the tempering. The volume fractions of both eutectic carbides and secondary carbides are invariant over the range of the heat-treatment parameters used in the experiments. The population density of the small globular carbides increased with the application of the sub-zero treatment, but rapidly decreased with the in- creasing temperature of the tempering (Figure 2). 3.2 Hardness characteristics The bulk hardness of the non-SZT and SZT Vanadis 6 steel, as a function of the tempering temperature, is J. PTA^INOVÁ et al.: FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT ... 730 Materiali in tehnologije / Materials and technology 51 (2017) 5, 729–733 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS shown in Table 1. The as-quenched hardness of the conventionally heat-treated steel that was not tempered was 838±4.33 HV10. The hardness of the SZT steel soaked in liquid nitrogen for 17 h and not tempered was correspondingly higher, e.g., 917±7.58 HV10. These results show that the as-quenched bulk hardness of the Vanadis 6 steel was improved due to the sub-zero treat- ment. The hardness of all the samples then decreases with the increasing tempering temperature. Here it is interesting that the hardness of the SZT steel remains higher up to the tempering temperature of 450 °C and then it drops more intensely than that of the convention- ally quenched samples. As a result, the hardness of the specimens tempered in the range of the temperatures commonly used for secondary hardening is lower for the SZT steel than for the steel achieved after conventional quenching and tempering. 3.3 Evaluation of the SZT effect on the fracture tough- ness The fracture toughness vs. heat treatment schedules are presented in Figure 3a (CHT) and Figure 3b (SZT). In the case of the CHT, the fracture toughness of the untempered material was measured to be 16.4 MPa m1/2 and it firstly increased, due to the tempering, to 19.6 MPa m1/2 with a subsequent slight decrease; however, the decrease in the fracture toughness markedly accelerated, at the temperature normally used for secondary harden- ing, to a value of 14.8 MPa m1/2. For the material subjected to the SZT in liquid nitrogen for 17 h, the fracture toughness before tempering was 13.3 MPa m1/2 (i.e., lower than that of the CHT steel). Then, the fracture toughness increased with the tempering temperature to more than 15 MPa m1/2 and just slightly decreased at a tempering temperature of 450 °C. It is interesting that the KIC values were higher for the SZT steel than for the CHT steel at the temperature of secondary hardening (530 °C). Table 1: Hardness of the Vanadis 6 ledeburitic steel Tempering of samples CHT SZT HV10 ± HV10 ± untempered 838 4.3 917 7.6 tempered at 170 °C 784 3.7 870 5.4 tempered at 330 °C 736 7 851 15.7 tempered at 450 °C 724 3 816 7.8 tempered at 530 °C 758 5.9 701 3 The behaviour of fracture toughness vs. heat treat- ment schedules can be classified as logical because one can expect higher KIC at lower hardness and lower KIC at elevated hardness. In other words, the application of the SZT decreases KIC when the steel is low-temperature tempered. This fact should be considered by toolmakers and users of tools in all the cases when they expect an increase in the wear resistance due to a high hardness resulting from the sub-zero treatment. A very interesting fact was found for the material tempered at the temperature of secondary hardening – KIC was higher for the SZT material. This is in good agreement with the recently published results where a similar tendency was found for the same SZT steel treated for 4 h and 10 h.7 One can say that this kind of KIC behaviour can be expected because of the lower hardness of the SZT steel when tempered in this temperature range; however, in the mentioned paper, it was also found that the increase of KIC is accompanied with a better wear performance of the material, due to the presence of a higher amount of carbides. The enhanced number of carbides, compared to the material after the conventional quenching, was also identified in the current work. Hence, one can expect that a sub-zero treated material would have a better wear performance, too. These results are also interesting from J. PTA^INOVÁ et al.: FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 729–733 731 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Amount of carbides in the Vanadis 6 steel after conventional quenching, SZT for 17 h and SZT and subsequent tempering at different temperatures Figure 1: SEM micrographs showing the microstructure of Vanadis 6 ledeburitic steel: a) after quenching, subsequent SZT and b) tempering at 170 °C, c) 330 °C, d) 450 °C, e) 530 °C, f) 600 °C the point of view of industrial practice. They indicate that it is possible to increase the wear performance, along the toughness, of the material in a certain tem- pering-temperature range. 3.4 Fracture-surface morphology The micro-mechanics of fracture propagation is demonstrated through representative SEM micrographs showing how the fractures appear in the cases of the steel that was sub-zero treated at –196 °C for 17 h and not tempered (the lowest KIC) and the steel that was sub-zero treated at –196 °C for 17 h and tempered at 170 °C (higher KIC). The fracture surface of the sample that was not tempered and the one that was tempered at 170 °C are shown in Figure 4. At first glance, there is not any difference between the mentioned fracture surfaces; they appear flat and shiny, Figures 4a and 4c. However, a more detailed observation, at higher magnifications, makes it clear that the fracture surface of the sample after low-temperature tempering contains an enhanced number of sites with micro-plastic deformation (PLD), located mainly at the carbide/matrix interfaces, Figure 4b. On the other hand, there are very few such sites on the fracture surface of the untempered specimen, Figure 4d, and the fracture surface contains a number of cleavage facettes/regions (CL). The topographies of the fracture surfaces of the CHT Vanadis 6 steel tempered at 170 °C and sub-zero treated, and the untempered one, expressed with surface rough- ness Ra are shown in Table 2 and Figure 5. The average roughness of the fracture surface of the conventionally quenched steel tempered at 170 °C was 4.2703±0.248 mm. J. PTA^INOVÁ et al.: FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT ... 732 Materiali in tehnologije / Materials and technology 51 (2017) 5, 729–733 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 4: Representative SEM micrographs of the fracture surfaces of the specimens processed with: a), b) SZT at –196 °C for 17 h and tem- pered at 170 °C and c), d) SZT at –196 °C for 17 h and not tempered Figure 3: Hardness HV10 and fracture toughness KIC of Vanadis 6 steel in dependence on the heat treatment applied: a) conventional quenching, b) sub-zero treatment Figure 5: Confocal-microscope micrographs showing the fracture surfaces of KIC specimens: a) CHT and tempered at 170 °C, b) SZT and not tempered The roughness of the sub-zero treated and untempered Vanadis 6 steel was lower, i.e., 2.918±0.323 μm. At this point, it should be noted that in this stage of experimental efforts, only the materials with the highest KIC and the one with the lowest KIC were checked with confocal microscopy. However, the measurements gave unambiguous differences in the roughness, where the higher roughness corresponded to the higher KIC, i.e., to the conventionally quenched and low-temperature tem- pered sample. Table 2: Roughness of the fracture surface of Vanadis 6 ledeburitic steel No. of measurement Tempered at 170 °C SZT anduntempered 1. 4.136 2.591 2. 4.750 3.507 3. 3.925 2.656 Ra (μm) 4.2703±0.248 2.918±0.323 4 CONCLUSIONS The main conclusions based on the presented results are as follows: • Sub-zero treatment reduces the amount of retained austenite and increases the population density of small globular carbides in the microstructure. • The amount of small globular carbides decreases with the application of tempering treatment. The higher the tempering temperature, the more signifi- cant is the reduction of the carbide count. • The hardness of the sub-zero treated material is higher than that of the conventionally quenched one. Also, this tendency is preserved when steel is low- temperature tempered. • On the other hand, the hardness of the conventionally quenched steel becomes higher than that of the SZT one when tempered at the temperature of secondary hardening. • The fracture toughness is generally higher for the conventionally quenched steel Vanadis 6 except in the case when the material is tempered to the se- condary hardness. One can conclude that KIC follows the reciprocal value of the hardness, i.e., the harder the material the lower is the fracture toughness. • The differences in KIC are reflected in the topo- graphies of the fracture surfaces, represented by the surface roughness – the higher the fracture toughness, the higher is the roughness of the fracture surface. Acknowledgements This paper is the result of implementing project CE for development and application of advanced diagnostic methods in processing of metallic and non-metallic materials, ITMS: 26220120048, supported by the Re- search & Development Operational Programme funded by the ERDF. This research was supported by the grant project VEGA 1/0735/14. 5 REFERENCES 1 P. Jur~i, Cr-V Ledeburitic Cold-Work Tool Steels, Mater. Tehnol., 45 (2011) 5, 383–394 2 D. N. Collins, Cryogenic treatment of tool steels, Advanced Ma- terials and Processes, 12 (1998), 24–29 3 D. N. Collins, J. Dormer, Deep Cryogenic Treatment of a D2 Cold- Work Tool Steel, Heat Treatment of Metals, 24 (1997) 3, 71–74 4 H. Berns, C. Bröckmann, Fracture of Hot Formed Ledeburitic Chro- mium Steels, Engineering Fracture Mechanics, 58 (1997), 311–325, doi:10.1016/S0013-7944(97)00118-5 5 I. Dlouhý, M. Holzmann, J. Valka, Using chevron notched specimens for determining fracture toughness of bearing steels, Kovove mate- riály/Metallic materials, 32 (1994), 3–13 6 D. Das, R. Sarkar, A. K. Dutta, K. K. Ray, Influence of sub-zero treatments on fracture toughness of AISI D2 steel, Materials Science and Engineering, A 528 (2010), 589–603, doi:10.1016/j.msea.2010. 09.057 7 J. Sobotová, P. Jur~i, I. Dlouhý, The effect of subzero treatment on microstructure, fracture toughness and wear resistance of Vanadis 6 tool steel, Materials Science and Engineering, A 652 (2016), 192–204, doi:10.1016/j.msea.2015.11.078 8 K. Amini, A. Akhbarizadeh, S. Javadpour, Investigating the effect of holding duration on the microstructure of 1.2080 tool steel during the deep cryogenic treatment, Vacuum, 86 (2012), 1534–1540, doi:10.1016/j.vacuum.2012.02.013 9 D. Das, A. K. Dutta, K. K. Ray, Sub-zero treatments of AISI D2 steel: Part I, Microstructure and hardness, Materials Science and Engineering, A 527 (2010), 2182–2193, doi:10.1016/j.msea.2009. 10.070 10 P. Jur~i, M. Dománková, ¼. ^aplovi~, J. Pta~inová, J. Sobotová, P. Salabová, O. Prikner, B. [u{tar{i~, D. Jenko, Microstructure and hardness of sub-zero treated and untempered P/M Vanadis 6 ledebu- ritic tool steel, Vacuum, 111 (2015), 92–101, doi:10.1016/j.vacuum. 2014.10.004 11 P. Bílek, J. Sobotová, P. Jur~i, Evaluation of the Microstructural Changes in Cr-V Ledeburitic Tool Steel Depending on the Austeni- tization Temperature, Mater. Tehnol., 45 (2011), 489–493 12 ^SN EN ISO 12137: 2010 – Metallic materials – determination of plane strain fracture toughness, European Committee for Standardi- zation, Brussels J. PTA^INOVÁ et al.: FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 729–733 733 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS the point of view of industrial practice. They indicate that it is possible to increase the wear performance, along the toughness, of the material in a certain tem- pering-temperature range. 3.4 Fracture-surface morphology The micro-mechanics of fracture propagation is demonstrated through representative SEM micrographs showing how the fractures appear in the cases of the steel that was sub-zero treated at –196 °C for 17 h and not tempered (the lowest KIC) and the steel that was sub-zero treated at –196 °C for 17 h and tempered at 170 °C (higher KIC). The fracture surface of the sample that was not tempered and the one that was tempered at 170 °C are shown in Figure 4. At first glance, there is not any difference between the mentioned fracture surfaces; they appear flat and shiny, Figures 4a and 4c. However, a more detailed observation, at higher magnifications, makes it clear that the fracture surface of the sample after low-temperature tempering contains an enhanced number of sites with micro-plastic deformation (PLD), located mainly at the carbide/matrix interfaces, Figure 4b. On the other hand, there are very few such sites on the fracture surface of the untempered specimen, Figure 4d, and the fracture surface contains a number of cleavage facettes/regions (CL). The topographies of the fracture surfaces of the CHT Vanadis 6 steel tempered at 170 °C and sub-zero treated, and the untempered one, expressed with surface rough- ness Ra are shown in Table 2 and Figure 5. The average roughness of the fracture surface of the conventionally quenched steel tempered at 170 °C was 4.2703±0.248 mm. J. PTA^INOVÁ et al.: FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT ... 732 Materiali in tehnologije / Materials and technology 51 (2017) 5, 729–733 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 4: Representative SEM micrographs of the fracture surfaces of the specimens processed with: a), b) SZT at –196 °C for 17 h and tem- pered at 170 °C and c), d) SZT at –196 °C for 17 h and not tempered Figure 3: Hardness HV10 and fracture toughness KIC of Vanadis 6 steel in dependence on the heat treatment applied: a) conventional quenching, b) sub-zero treatment Figure 5: Confocal-microscope micrographs showing the fracture surfaces of KIC specimens: a) CHT and tempered at 170 °C, b) SZT and not tempered The roughness of the sub-zero treated and untempered Vanadis 6 steel was lower, i.e., 2.918±0.323 μm. At this point, it should be noted that in this stage of experimental efforts, only the materials with the highest KIC and the one with the lowest KIC were checked with confocal microscopy. However, the measurements gave unambiguous differences in the roughness, where the higher roughness corresponded to the higher KIC, i.e., to the conventionally quenched and low-temperature tem- pered sample. Table 2: Roughness of the fracture surface of Vanadis 6 ledeburitic steel No. of measurement Tempered at 170 °C SZT anduntempered 1. 4.136 2.591 2. 4.750 3.507 3. 3.925 2.656 Ra (μm) 4.2703±0.248 2.918±0.323 4 CONCLUSIONS The main conclusions based on the presented results are as follows: • Sub-zero treatment reduces the amount of retained austenite and increases the population density of small globular carbides in the microstructure. • The amount of small globular carbides decreases with the application of tempering treatment. The higher the tempering temperature, the more signifi- cant is the reduction of the carbide count. • The hardness of the sub-zero treated material is higher than that of the conventionally quenched one. Also, this tendency is preserved when steel is low- temperature tempered. • On the other hand, the hardness of the conventionally quenched steel becomes higher than that of the SZT one when tempered at the temperature of secondary hardening. • The fracture toughness is generally higher for the conventionally quenched steel Vanadis 6 except in the case when the material is tempered to the se- condary hardness. One can conclude that KIC follows the reciprocal value of the hardness, i.e., the harder the material the lower is the fracture toughness. • The differences in KIC are reflected in the topo- graphies of the fracture surfaces, represented by the surface roughness – the higher the fracture toughness, the higher is the roughness of the fracture surface. Acknowledgements This paper is the result of implementing project CE for development and application of advanced diagnostic methods in processing of metallic and non-metallic materials, ITMS: 26220120048, supported by the Re- search & Development Operational Programme funded by the ERDF. This research was supported by the grant project VEGA 1/0735/14. 5 REFERENCES 1 P. Jur~i, Cr-V Ledeburitic Cold-Work Tool Steels, Mater. Tehnol., 45 (2011) 5, 383–394 2 D. N. Collins, Cryogenic treatment of tool steels, Advanced Ma- terials and Processes, 12 (1998), 24–29 3 D. N. Collins, J. Dormer, Deep Cryogenic Treatment of a D2 Cold- Work Tool Steel, Heat Treatment of Metals, 24 (1997) 3, 71–74 4 H. Berns, C. Bröckmann, Fracture of Hot Formed Ledeburitic Chro- mium Steels, Engineering Fracture Mechanics, 58 (1997), 311–325, doi:10.1016/S0013-7944(97)00118-5 5 I. Dlouhý, M. Holzmann, J. Valka, Using chevron notched specimens for determining fracture toughness of bearing steels, Kovove mate- riály/Metallic materials, 32 (1994), 3–13 6 D. Das, R. Sarkar, A. K. Dutta, K. K. Ray, Influence of sub-zero treatments on fracture toughness of AISI D2 steel, Materials Science and Engineering, A 528 (2010), 589–603, doi:10.1016/j.msea.2010. 09.057 7 J. Sobotová, P. Jur~i, I. Dlouhý, The effect of subzero treatment on microstructure, fracture toughness and wear resistance of Vanadis 6 tool steel, Materials Science and Engineering, A 652 (2016), 192–204, doi:10.1016/j.msea.2015.11.078 8 K. Amini, A. Akhbarizadeh, S. Javadpour, Investigating the effect of holding duration on the microstructure of 1.2080 tool steel during the deep cryogenic treatment, Vacuum, 86 (2012), 1534–1540, doi:10.1016/j.vacuum.2012.02.013 9 D. Das, A. K. Dutta, K. K. Ray, Sub-zero treatments of AISI D2 steel: Part I, Microstructure and hardness, Materials Science and Engineering, A 527 (2010), 2182–2193, doi:10.1016/j.msea.2009. 10.070 10 P. Jur~i, M. Dománková, ¼. ^aplovi~, J. Pta~inová, J. Sobotová, P. Salabová, O. Prikner, B. [u{tar{i~, D. Jenko, Microstructure and hardness of sub-zero treated and untempered P/M Vanadis 6 ledebu- ritic tool steel, Vacuum, 111 (2015), 92–101, doi:10.1016/j.vacuum. 2014.10.004 11 P. Bílek, J. Sobotová, P. Jur~i, Evaluation of the Microstructural Changes in Cr-V Ledeburitic Tool Steel Depending on the Austeni- tization Temperature, Mater. Tehnol., 45 (2011), 489–493 12 ^SN EN ISO 12137: 2010 – Metallic materials – determination of plane strain fracture toughness, European Committee for Standardi- zation, Brussels J. PTA^INOVÁ et al.: FRACTURE TOUGHNESS OF LEDEBURITIC VANADIS 6 STEEL AFTER SUB-ZERO TREATMENT ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 729–733 733 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS C. XIANG et al.: ELECTRONIC AND OPTICAL PROPERTIES OF THE SPINEL OXIDES ... 735–743 ELECTRONIC AND OPTICAL PROPERTIES OF THE SPINEL OXIDES MgxZn1-xAl2O4 BY FIRST-PRINCIPLES CALCULATIONS ELEKTRONSKE IN OPTI^NE LASTNOSTI SPINELNIH OKSIDOV MgxZn1-xAl2O4, IZPELJANE IZ TEORETI^NIH OSNOV Chao Xiang1, Jianxiong Zhang1, Yun Lu2, Dong Tian3, Cheng Peng1 1Yangtze Normai University, School of Mechanical and Engineering, Fuling 408000, China 2Chiba University, Institute of Material Science and Engineering, Chiba 2790000, Japan 3Kunming University of Science and Technology, Faculty of Science, Kunming 650093, China 1254618608@qq.com Prejem rokopisa – received: 2016-10-07; sprejem za objavo – accepted for publication: 2017-02-10 doi:10.17222/mit.2016.296 The structural, electronic and optical properties of perfect MgxZn1-xAl2O4 oxides have been studied by first-principles calculations within the generalized gradient approximation of the density functional theory. It is interesting to note that a linear increase of cell volume (V) with increasing doping amount (x) occurs. The band gap increases in the series from 3.851 eV to 5.079 eV, which is in agreement with theoretical and experimental values. In addition, a blue shift of the absorption shoulder is observed in the UV region with the increase of x, as predicted by the imaginary part 2() of the dielectric function at zero frequency as well as bandgap. This can be explained by the threshold of the electronic transition from O-2p to the empty Mg-3p electron states due to the substitution of Zn with Mg. The real part 1() of the dielectric function located at zero frequency has a square fit relationship with refractive index n(0), which is 1.71–1.77 from x=0 to x=1. The energy-loss function shows that the replacement of Zn by Mg is responsible for a decrease in the intensity of the sharp peaks. The reflectivity shows that a higher coefficient of reflectivity (R(0)) at zero frequency corresponds to a smaller bandgap. Keywords: electronic transitions, dielectric function, refractive index, adsorption shoulder Preiskovali smo strukturne, elektronske in opti~ne lastnosti idealnih MgxZn1-xAl2O4 oksidov, izpeljane iz teoreti~nih osnov znotraj posplo{ene gradientne aproksimacije funkcionalne teorije gostote. Opazili smo linearno pove~anje celi~nega volumna (V) z nara{~ajo~o koncentracijo Mg (x). [irina prepovedanega pasu nara{~a od 3.851 eV to 5.079 eV v skladu s teoreti~nimi in eksperimentalnimi vrednostmi. Poleg tega ob nara{~anju dele`a Mg opazimo modri premik dodatnega absorpcijskega vrha v UV obmo~ju, kakor napovedujeta vrednosti imaginarnega dela 2() dielektri~ne funkcije pri frekvenci 0 in prepovedanega pasu. To je mogo~e pojasniti s pragom elektronskega prehoda elektronov iz O-2p v prazen Mg-3p zaradi nadomestitve Zn z Mg. Kvadrat realnega dela 1() dielektri~ne funkcije pri frekvenci 0 se ujema z lomnim koli~nikom n(0), ki je 1.71–1.77 za x=0 do x=1. Funkcija izgube energije, ka`e, da zamenjava Zn z Mg povzro~a zmanj{anje intenzitete ostrih vrhov. Reflektivnost ka`e, da vi{ji koeficient refleksije (R(0)) pri frekvenci 0 odgovarja manj{i {irini prepovedanega pasu. Klju~ne besede: prehodi elektronov, dielektri~na funkcija, lomni koli~nik, dodatni absorpcijski vrh 1 INTRODUCTION Spinal oxides with the general chemical formula AB2O4 have a close-packed, face-centered-cubic struc- ture (space group Fd3m) characterized by two symmetrically distinct polyhedra: a tetrahedron and an octahedron. They are widely used in various fields such as catalysis, gas sensor, semiconductor, biomedical, catalyst carrier, as well as electroluminescent displays owing to their catalytic, physical, structural, electronic and optical properties.1,2 Among them, MgAl2O4 and ZnAl2O4 have high-temperature resistance,3,4 and they are highly reflective for wavelengths in the ultraviolet (UV) region, which make them candidate materials for reflective optical coating in aerospace applications.5 In particular, MgAl2O4 is one of the potential candidates for the full wave band transparent window materials with high transmittance in IR- and visible-wavelength even extending to microwave ranges,6,7 and it also can be used as lamps and lasers,8 transparent ceramic material for high-temperature,9 transparent armor and glass.10 Similarly, ZnAl2O4 can be used as a ceramic material similar to MgAl2O4. It can be used as a transparent conductor, optical material and dielectric material,11,12 and it is suitable for UV photoelectronic applications.13 Simultaneously, much work has been done on the structural, electronic and optical properties of MgAl2O4 and ZnAl2O4 over the past few years.14–27 The effect of point vacancies on the spectral properties of MgAl2O4 has been studied by S. L. Jiang et al.14 They revealed that the absorption peak at 5.3 eV is attributed to the neutral oxygen vacancy Vo 0 , while two peaks at 3.2 eV and 4.75 eV are attributed to the 1+ charged oxygen vacancy Vo 1 + . A related mechanism of transparency in MgAl2O4 nano- ceramics prepared by sintering under high pressure and low temperature has been studied by J. Zhang et al.15, who suggested that the decrease in the transparency with increasing temperature (>700 °C) is therefore a result of the light scattering at large pores. The low-temperature, high-pressure preparation of transparent nanocrystalline MgAl2O4 ceramics has been investigated by T. C. Lu et Materiali in tehnologije / Materials and technology 51 (2017) 5, 735–743 735 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 67.017:621.3.011.5:535.327 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)735(2017) C. XIANG et al.: ELECTRONIC AND OPTICAL PROPERTIES OF THE SPINEL OXIDES ... 735–743 ELECTRONIC AND OPTICAL PROPERTIES OF THE SPINEL OXIDES MgxZn1-xAl2O4 BY FIRST-PRINCIPLES CALCULATIONS ELEKTRONSKE IN OPTI^NE LASTNOSTI SPINELNIH OKSIDOV MgxZn1-xAl2O4, IZPELJANE IZ TEORETI^NIH OSNOV Chao Xiang1, Jianxiong Zhang1, Yun Lu2, Dong Tian3, Cheng Peng1 1Yangtze Normai University, School of Mechanical and Engineering, Fuling 408000, China 2Chiba University, Institute of Material Science and Engineering, Chiba 2790000, Japan 3Kunming University of Science and Technology, Faculty of Science, Kunming 650093, China 1254618608@qq.com Prejem rokopisa – received: 2016-10-07; sprejem za objavo – accepted for publication: 2017-02-10 doi:10.17222/mit.2016.296 The structural, electronic and optical properties of perfect MgxZn1-xAl2O4 oxides have been studied by first-principles calculations within the generalized gradient approximation of the density functional theory. It is interesting to note that a linear increase of cell volume (V) with increasing doping amount (x) occurs. The band gap increases in the series from 3.851 eV to 5.079 eV, which is in agreement with theoretical and experimental values. In addition, a blue shift of the absorption shoulder is observed in the UV region with the increase of x, as predicted by the imaginary part 2() of the dielectric function at zero frequency as well as bandgap. This can be explained by the threshold of the electronic transition from O-2p to the empty Mg-3p electron states due to the substitution of Zn with Mg. The real part 1() of the dielectric function located at zero frequency has a square fit relationship with refractive index n(0), which is 1.71–1.77 from x=0 to x=1. The energy-loss function shows that the replacement of Zn by Mg is responsible for a decrease in the intensity of the sharp peaks. The reflectivity shows that a higher coefficient of reflectivity (R(0)) at zero frequency corresponds to a smaller bandgap. Keywords: electronic transitions, dielectric function, refractive index, adsorption shoulder Preiskovali smo strukturne, elektronske in opti~ne lastnosti idealnih MgxZn1-xAl2O4 oksidov, izpeljane iz teoreti~nih osnov znotraj posplo{ene gradientne aproksimacije funkcionalne teorije gostote. Opazili smo linearno pove~anje celi~nega volumna (V) z nara{~ajo~o koncentracijo Mg (x). [irina prepovedanega pasu nara{~a od 3.851 eV to 5.079 eV v skladu s teoreti~nimi in eksperimentalnimi vrednostmi. Poleg tega ob nara{~anju dele`a Mg opazimo modri premik dodatnega absorpcijskega vrha v UV obmo~ju, kakor napovedujeta vrednosti imaginarnega dela 2() dielektri~ne funkcije pri frekvenci 0 in prepovedanega pasu. To je mogo~e pojasniti s pragom elektronskega prehoda elektronov iz O-2p v prazen Mg-3p zaradi nadomestitve Zn z Mg. Kvadrat realnega dela 1() dielektri~ne funkcije pri frekvenci 0 se ujema z lomnim koli~nikom n(0), ki je 1.71–1.77 za x=0 do x=1. Funkcija izgube energije, ka`e, da zamenjava Zn z Mg povzro~a zmanj{anje intenzitete ostrih vrhov. Reflektivnost ka`e, da vi{ji koeficient refleksije (R(0)) pri frekvenci 0 odgovarja manj{i {irini prepovedanega pasu. Klju~ne besede: prehodi elektronov, dielektri~na funkcija, lomni koli~nik, dodatni absorpcijski vrh 1 INTRODUCTION Spinal oxides with the general chemical formula AB2O4 have a close-packed, face-centered-cubic struc- ture (space group Fd3m) characterized by two symmetrically distinct polyhedra: a tetrahedron and an octahedron. They are widely used in various fields such as catalysis, gas sensor, semiconductor, biomedical, catalyst carrier, as well as electroluminescent displays owing to their catalytic, physical, structural, electronic and optical properties.1,2 Among them, MgAl2O4 and ZnAl2O4 have high-temperature resistance,3,4 and they are highly reflective for wavelengths in the ultraviolet (UV) region, which make them candidate materials for reflective optical coating in aerospace applications.5 In particular, MgAl2O4 is one of the potential candidates for the full wave band transparent window materials with high transmittance in IR- and visible-wavelength even extending to microwave ranges,6,7 and it also can be used as lamps and lasers,8 transparent ceramic material for high-temperature,9 transparent armor and glass.10 Similarly, ZnAl2O4 can be used as a ceramic material similar to MgAl2O4. It can be used as a transparent conductor, optical material and dielectric material,11,12 and it is suitable for UV photoelectronic applications.13 Simultaneously, much work has been done on the structural, electronic and optical properties of MgAl2O4 and ZnAl2O4 over the past few years.14–27 The effect of point vacancies on the spectral properties of MgAl2O4 has been studied by S. L. Jiang et al.14 They revealed that the absorption peak at 5.3 eV is attributed to the neutral oxygen vacancy Vo 0 , while two peaks at 3.2 eV and 4.75 eV are attributed to the 1+ charged oxygen vacancy Vo 1 + . A related mechanism of transparency in MgAl2O4 nano- ceramics prepared by sintering under high pressure and low temperature has been studied by J. Zhang et al.15, who suggested that the decrease in the transparency with increasing temperature (>700 °C) is therefore a result of the light scattering at large pores. The low-temperature, high-pressure preparation of transparent nanocrystalline MgAl2O4 ceramics has been investigated by T. C. Lu et Materiali in tehnologije / Materials and technology 51 (2017) 5, 735–743 735 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 67.017:621.3.011.5:535.327 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)735(2017) al.,16 indicating that the nanoceramics are highly transparent even though their relative densities are all less than 99 %, owing to the low or negligible light scattering from the nanosized grains and pores. The optical properties of ZnAl2O4 nanomaterials obtained by the hydrothermal method have been investigated by Miron and co-workers,17 demonstrating that the band gap is determined from the absorbance spectra, and it de- pends strongly on the temperature used for further heating the samples. A first-principles study on struc- tural, electronic and optical properties of spinel oxides ZnAl2O4, ZnGa2O4 and ZnIn2O4 has been carried out by F. Zerarga and co-workers,18 implying that the peaks and structures in the optical spectra are assigned to interband transitions. The fabrication of transparent polycrystalline ZnAl2O4 – a new optical bulk ceramic – has been inve- stigated by Goldstein and co-workers19, who suggested that specimens have a high transparency (ILT78\%;  = 800 nm; t = 2 mm). Plus, the differences in struc- tural, electronic and optical performance between alumi- num spinel MgAl2O4 and ZnAl2O4 have been presented.28–30 To the best of our knowledge, the struc- ture, electronic and optical properties of MgxZn1-xAl2O4 (0 for the phases appearing in the material at various milling times were determined. The apparatus factors were eliminated using LaB6 (SRM 660a). The morphology of the initial powders and the microstructure of the sintered material were tested using an optic microscope and scanning microscope JEOL JSM 6480 with an accelerating voltage of 20 kV. Additionally, a chemical analysis was performed using an EDS detector manufactured by IXRF using the standard calibration method. An analysis of the microstructure was performed on microscopic images with the use of ImageJ, which is the software used for image processing and analysis. The size and shape of the grains were determined using the planimetric method. The study also included determining the inhomogeneity of the size of the grains. A quantita- tive analysis of the morphology of the material was performed using the stereological parameters presented below: a) the size of the area of a section of grain a (μm2), b) dimensionless lengthening factor (Equation (1): f h wg1 = (1) where: h – height and w – width of the smallest rectangle described on the object, c) dimensionless shape factor – the circularity: f F Lg 2 2 4 = π (2) where: F – area of the analyzed object; L – perimeter of the analyzed object, d) grain-size change-ability factor:   = x a (3) where: x – grain-size standard deviation; a – grain mean value, the number of analyzed elements per area unit of the image: NA [ ]1 2/mm . The thermal behavior of the martensitic transforma- tion was studied using a differential scanning calorimeter (DSC) Mettler Toledo DSC-1. Transformation tempera- tures were determined from the thermograms measured at a heating rate of 10 deg/min and a thermal range of –120 °C to 600 °C. The surface morphology, hardness and elastic modu- lus were tested using a Hysitron Triboindenter Ti950 with AFM QScope 250. The measurements were calculated from the load-displacement data obtained with the nanoindentation, using a three-sided, pyramidal, diamond (Berkovich) indenter tip. An area of 40 μm × 40 μm was analyzed at a scanning frequency of 0.25 Hz. The microhardness measurement was carried out parallel to the nanoindentation tests. The Vickers microhardness measurement was taken at a load of 500 N and a loading time of 10 s on a 401MVD microhardness tester. G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... 796 Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 3 RESULTS AND DISCUSSION SEM micrographs of the initial metal powders are presented in Figure 1. The titanium-powder morphology is irregular, having sharp corners and rough surfaces. The dispersion of the Ti particles is wide and ranges from a few μm to even 50 μm. The niobium powder has a much wider dispersion of the particles in comparison to the Ti powder. Some of the particles have the size below 10 μm, and the shape of these particles is irregular. Simultaneously, there are also very large particles, the sizes of which are greater even than 50 μm and these particles have an irregular, polyhedral shape. As a result of the work-hard- ening effect and high density of defects such as dislo- cations and vacancies created during long ball-milling times, the fracture and welding of powders reach an equilibrium state. This state leads to the formation of rather equiaxed and small particles.16,17 However, due to the higher bonding strength of the finer particles, their ability to undergo further plastic deformation is de- creased and a higher force is required to fracture the small particles during the ball-milling process.18 It is worth mentioning that the reduction in the particle size at a given milling time can also be influenced by the milling technique. The morphologies of the Ti-26Nb (a/%) powders at different milling stages are shown in Figure 2. The powder after the shortest milling time shows globular particles. In the picture with the higher magnification (Figure 2b), some cracks were observed, which is the beginning of the cyclic process of material milling during high-energy ball milling. In contrast, in the samples after 70 h of milling, the particles became larger and the shape of the particles became less globular. Moreover, the particles of this sample are cold welded and the powder particles, in particular, are composed of the welded layers of the material. This behavior of the material during the high-energy milling is consistent with the conduct and following steps of mechanical alloying.19 As shown by the other studies, the mechani- cal-alloying method leads to a decreased particle size, the -Ti phase being stabilized by Nb, and an increased surface area between Ti and Nb. This promotes good homogeneity of the solid solution.20 Figure 3 shows a summary of diffraction patterns of the material after the mechanical synthesis, depending on the milling time of the initial powders. The X-ray qualitative analysis showed that only  and  phases are present in the material after milling for 50 h and 70 h. Change of the profile lines of individual diffraction patterns are clearly visible. The broadening of the particle-size distribution at longer ball-milling times is a typical behavior of high- energy ball-milling process.17,19,21,22 For the  and  phases, it was observed that a decrease in the crystallite size took place simultaneously with an increase in the milling time (Table 1). In addition, in the case of the  phase there was a tendency towards a decrease in the lattice strain <a/a> (Table 1). The estimated size of the crystallites for the powder after milling for 70 h is 9(1) nm and 16(2) nm for the  and  phases, respectively. The observed <a/a> lattice strain values indicate growth in the  phase (3.43E-03 %) and in the  phase G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 797 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: X-ray diffraction patterns of the powder after 50 h and 70 h of the milling process Figure 1: SEM micrographs of the initial Ti and Nb powders at different magnifications Figure 2: SEM micrographs of particles after different milling times: a), b) 50 h of milling and c), d) 70 h these alloys are obtained with a conventional method such as arc melting. One of the alternatives to the widely used arc melting is powder metallurgy (PM). In the scientific literature, there are evidences of successful preparations of titanium-based materials with this pro- duction method.8–10 This production method has a number of significant advantages. It allows creating alloys of materials with large differences between their melting temperatures or densities. The most important advantages of PM for the production of biomaterials are its low cost and the possibility to produce highly porous materials.4,11 The aim of the present paper is to determine the possibility to produce a Ti–26Nb (a/%) alloy using the powder-metallurgy method. Simultaneously, this re- search is to develop a technology of producing Ti-based alloys and estimate the influence of changing parameters during the mechanical alloying. Such research is the starting point for further development of nickel-free alloys with the shape-memory effect based on titanium and niobium. In this article, the influence of the milling time on the microstructure and mechanical properties of the produced alloy is investigated. 2 EXPERIMENTAL PART Commercial elemental powders of Ti (Atlantic Equipment Engineers (AEE), 99.7 %, <20 μm) and Nb (Atlantic Equipment Engineers (AEE), 99.8 %, < 5 μm) were used as the initial materials for the synthesis of the alloy. Elemental metal powders with the nominal com- position of Ti-26Nb (a/%) were milled for two different milling times, 50 h and 70 h, in a planetary ball mill Fritch PULVERISETTE 7 premium line. To prevent the powder from oxidation as much as possible, the process was carried out in an argon-protective atmosphere. During the shorter milling time, the speed of 200 min–1 was applied, and for the longer milling time (with additional 20 h), the speed of 400 min–1 was applied. The powders were cold isostatically pressed under a 750 MPa pressure without any substances that improve the porosity (e.g., space holders). Then the material was sintered at 1000 °C for 24 h and cooled in the furnace to room temperature. The crystalline structure and phase content of the sintered materials were tested with X-ray diffraction. The refinement of the X-ray diffraction pattern was carried out using Rietveld’s whole X-ray profile fitting technique with the DBWS 9807a program.12 The profile function used to adjust the calculated diffractograms to the observed ones was the pseudo-Voigt one.13,14 The weight fraction of each component was determined based on the optimized scale factors with the use of the relation proposed by Hill and Howard.15,16 Based on the Williamson–Hall method, the crystallite sizes (D) and lattice strain <a/a> for the phases appearing in the material at various milling times were determined. The apparatus factors were eliminated using LaB6 (SRM 660a). The morphology of the initial powders and the microstructure of the sintered material were tested using an optic microscope and scanning microscope JEOL JSM 6480 with an accelerating voltage of 20 kV. Additionally, a chemical analysis was performed using an EDS detector manufactured by IXRF using the standard calibration method. An analysis of the microstructure was performed on microscopic images with the use of ImageJ, which is the software used for image processing and analysis. The size and shape of the grains were determined using the planimetric method. The study also included determining the inhomogeneity of the size of the grains. A quantita- tive analysis of the morphology of the material was performed using the stereological parameters presented below: a) the size of the area of a section of grain a (μm2), b) dimensionless lengthening factor (Equation (1): f h wg1 = (1) where: h – height and w – width of the smallest rectangle described on the object, c) dimensionless shape factor – the circularity: f F Lg 2 2 4 = π (2) where: F – area of the analyzed object; L – perimeter of the analyzed object, d) grain-size change-ability factor:   = x a (3) where: x – grain-size standard deviation; a – grain mean value, the number of analyzed elements per area unit of the image: NA [ ]1 2/mm . The thermal behavior of the martensitic transforma- tion was studied using a differential scanning calorimeter (DSC) Mettler Toledo DSC-1. Transformation tempera- tures were determined from the thermograms measured at a heating rate of 10 deg/min and a thermal range of –120 °C to 600 °C. The surface morphology, hardness and elastic modu- lus were tested using a Hysitron Triboindenter Ti950 with AFM QScope 250. The measurements were calculated from the load-displacement data obtained with the nanoindentation, using a three-sided, pyramidal, diamond (Berkovich) indenter tip. An area of 40 μm × 40 μm was analyzed at a scanning frequency of 0.25 Hz. The microhardness measurement was carried out parallel to the nanoindentation tests. The Vickers microhardness measurement was taken at a load of 500 N and a loading time of 10 s on a 401MVD microhardness tester. G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... 796 Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 3 RESULTS AND DISCUSSION SEM micrographs of the initial metal powders are presented in Figure 1. The titanium-powder morphology is irregular, having sharp corners and rough surfaces. The dispersion of the Ti particles is wide and ranges from a few μm to even 50 μm. The niobium powder has a much wider dispersion of the particles in comparison to the Ti powder. Some of the particles have the size below 10 μm, and the shape of these particles is irregular. Simultaneously, there are also very large particles, the sizes of which are greater even than 50 μm and these particles have an irregular, polyhedral shape. As a result of the work-hard- ening effect and high density of defects such as dislo- cations and vacancies created during long ball-milling times, the fracture and welding of powders reach an equilibrium state. This state leads to the formation of rather equiaxed and small particles.16,17 However, due to the higher bonding strength of the finer particles, their ability to undergo further plastic deformation is de- creased and a higher force is required to fracture the small particles during the ball-milling process.18 It is worth mentioning that the reduction in the particle size at a given milling time can also be influenced by the milling technique. The morphologies of the Ti-26Nb (a/%) powders at different milling stages are shown in Figure 2. The powder after the shortest milling time shows globular particles. In the picture with the higher magnification (Figure 2b), some cracks were observed, which is the beginning of the cyclic process of material milling during high-energy ball milling. In contrast, in the samples after 70 h of milling, the particles became larger and the shape of the particles became less globular. Moreover, the particles of this sample are cold welded and the powder particles, in particular, are composed of the welded layers of the material. This behavior of the material during the high-energy milling is consistent with the conduct and following steps of mechanical alloying.19 As shown by the other studies, the mechani- cal-alloying method leads to a decreased particle size, the -Ti phase being stabilized by Nb, and an increased surface area between Ti and Nb. This promotes good homogeneity of the solid solution.20 Figure 3 shows a summary of diffraction patterns of the material after the mechanical synthesis, depending on the milling time of the initial powders. The X-ray qualitative analysis showed that only  and  phases are present in the material after milling for 50 h and 70 h. Change of the profile lines of individual diffraction patterns are clearly visible. The broadening of the particle-size distribution at longer ball-milling times is a typical behavior of high- energy ball-milling process.17,19,21,22 For the  and  phases, it was observed that a decrease in the crystallite size took place simultaneously with an increase in the milling time (Table 1). In addition, in the case of the  phase there was a tendency towards a decrease in the lattice strain <a/a> (Table 1). The estimated size of the crystallites for the powder after milling for 70 h is 9(1) nm and 16(2) nm for the  and  phases, respectively. The observed <a/a> lattice strain values indicate growth in the  phase (3.43E-03 %) and in the  phase G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 797 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: X-ray diffraction patterns of the powder after 50 h and 70 h of the milling process Figure 1: SEM micrographs of the initial Ti and Nb powders at different magnifications Figure 2: SEM micrographs of particles after different milling times: a), b) 50 h of milling and c), d) 70 h (2.31E-03 %). The analysis of the diffraction patters obtained using the Rietveld method showed that that high-energy milling has an influence on the lattice parameters. Table 1: Changes of the average crystallite sizes (D) and lattice distortions (<a/a>) of the  and  phases of the milled powders Phase Parameters Milling time (h) 50 70  D (nm) 22(2) 16(1) <a/a> (%) 2.76E-03 3.43E-03  D (nm) 15(2) 9(1) <a/a> (%) 2.04E-03 2.31E-03 Table 2 presents lattice parameters determined for individual phases and the corresponding ICDD data sheets. In the case of the  phase, a slight deviation of the a0 lattice parameters from the catalogue data was observed. A similar type of cell contraction was discovered for the a0 parameter of the  phase and for the c0 parameter, but these changes are smaller, resulting from the shift systems for the phases of the hexagonal system. It should be stressed that for the sample after 70 h of milling, the deviation of both lattice parameters can be noted. It is probably the effect of the alternate cold welding and deposition of the particles during the mechanical synthesis. Table 2: Lattice parameters and contents of  and  phases of the milled powders Phase Lattice parameters (nm) ICDD*  a0 0.2970 c0 0.4720  a0 0.3307 * International Centre for Diffraction Data® – a scientific organization dedicated to collecting, editing, publishing and distributing powder- diffraction data for the identification of materials The quantitative phase analysis of the material after each stage show a successive progressive synthesis of the as-prepared powders in relation to the  phase (77.5(11)) % mass fraction and 82.6(12) % mass fraction, for 50 h and 70 h milling times, respectively). The X-ray qualitative analysis showed that only  and  phases are present in the samples annealed at 1000 °C for 24 h from the powder previously milled for 50 h and 70 h, as shown in Figure 4. The XRD shows that there were no obvious diffraction peaks of elemental Nb remaining in the sintered Ti–26Nb (a/%). The diffraction peaks attributable to the  and  phases from the pattern confirm a duplex microstructure. M. Tahara et al.23 confirmed only the  phase for Ti-26Nb (a/%) produced by arc melting. The predominant phase in all the samples was the body-centered-cubic (bcc)  phase. This indicates that Nb was diffused into Ti, thereby lead- ing to the formation of  +  phases after the sintering. This is significant because two-phase Ti alloys, with the major  phase and the minor  phase, possess good com- prehensive properties including high yield strength, ex- cellent corrosion resistance and good fracture toughness. The presence of unalloyed Ti particles in materials may arise from different facts. According to S. N. Patankar and F. H. Froes24, it could be correlated with the brittleness of the Nb particles, as compared to the Ti particles, since these brittle particles adhere to the sur- faces of ductile particles. Another reason could be the limited solubility of Nb in -Ti. It is known that iso- morphous -stabilizer elements, such as Nb, exhibit slower diffusivities in Ti than the eutectoid elements of Fe, Co and Ni.25 The structural analysis of the diffraction patters obtained using the Rietveld method showed that the lattice parameters for all the tested samples become reduced. In comparison to the ICDD data, it was found that high-energy milling and annealing have the smallest, but significant effect on the reduction of the size of a unit cell of the  and  phases. The quantitative phase anal- ysis (Table 3) of the material after annealing showed a significant synthesis of the as-prepared powders in rela- tion to the  phase 96.6(12) w/% and 98.8(12) w/% for samples previously milled for 50 h and 70 h, respecti- vely). Table 3: Lattice parameters and contents of  and  phases after annealing (A) the samples at 1000 °C for 24 h from the powder previously milled for 50 h and 70 h Phase Lattice parameters (nm) Phase Contents (w/%) ICDD* Rietveld 50 h + A 70 h + A 50 h + A 70 h + A  a0 0.2970 0.2898(2) 0.2891 (2)  3.4 (6) 1.2 (4) c0 0.4720 0.4715(4) 0.4710 (4)  96.6 (12) 98.8 (12)  a0 0.3307 0.3287(2) 0.3286 (2) * International Centre for Diffraction Data®– a scientific organization dedicated to collecting, editing, publishing and distributing powder-diffraction data for the identification of materials Optical and SEM micrographs of the studied samples are presented in Figures 5 and 6. They clearly show the G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... 798 Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 4: X-ray diffraction patterns for the samples annealed (+A) at 1000 °C for 24 h from the powder previously milled for 50 h and 70 h changes in the morphology of the material after sintering with respect to various milling times. The common feature of the samples is a hierarchical structure. After the milling, large-sized particles caused the formation of a porous material. The pores were large and intercon- nected (Figures 5a and 5c) in spite of the applied pressure of 750 MPa during the isostatic pressing. In the case of the sample milled for 50 h (Figure 5b), an anomalous grain growth was observed. SEM micro- graphs (Figure 6) revealed further differences between the obtained microstructures. The powders ball-milled for a long time cannot move and deform effectively during the compaction. This may be primarily due to the effects of the decrease in the particle size, solid-solution strengthening and work hardening of the particles during the ball-milling process.26,27 Amongst these factors, the work-hardening effect appears to be the main deter- minant in governing the plastic behavior of the powder particles during the compaction. In the ball-milling process, work hardening is caused by the interaction of dislocations with each other through the continuous mechanical impact of the balls on the powder particles.16 A quantitative analysis of the microstructure was conducted including the determination of the cross-sec- tional area of the grains and the subgrains, as well as the average circularity of the subgrains of the microsections based on microscopic images, the results of which are presented in Tables 4 to 6. Significant changes between the samples in the average a of the grain cross-sectional area can also be observed. For the samples milled for 50 h and 70 h, the average section area of grains was 9.09 μm2 and 3.55 μm2, respectively (Table 4). A change in the time and rotations per minute during milling causes a difference in the parameter. Table 4: Results for the cross-sectional area of the grains for the samples milled for different times Parameter Sample Minimumvalue Maximum value Average value Std. dev. a (μm2) 50 h 0.51 267.57 9.09 12.68 70 h 0.51 32.35 3.55 3.45 Table 5: Results for the circularity of the grains for the samples milled for different times Parameter Sample Minimumvalue Maximum value Average value Std. dev. fg2 50 h 0.10 1.00 0.60 0.15 70 h 0.11 0.96 0.59 0.15 Table 6: Results for the dimensionless lengthening factor of the grains for the samples milled for different times Parameter Sample Minimumvalue Maximum value Average value Std. dev. fg1 50 h 0.32 3.31 1.06 0.37 70 h 0.28 3.30 1.06 0.37 Histograms of the size of the area of the section of the grains in the samples milled for 50 h and 70 h are presented in Figures 7 and 8, respectively. For the sam- ple milled for 70 h, a larger fraction of the grains with the section area in the range of 0.5–35 μm can be ob- served. Milling for 70 h caused a significant fragmen- tation of the grains with the section area below 30 μm2. G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 799 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: a), b) SEM micrographs of Ti-26Nb (a/%) after different milling times – 50 h and c), d) 70 h and annealing Figure 5: a), b) Optical micrographs of Ti-26Nb (a/%) at different magnifications, produced in different milling times – 50 h and c), d) 70 h and annealing Figure 7: Histogram of the grain average area for the sample milled for 50 h and annealed (2.31E-03 %). The analysis of the diffraction patters obtained using the Rietveld method showed that that high-energy milling has an influence on the lattice parameters. Table 1: Changes of the average crystallite sizes (D) and lattice distortions (<a/a>) of the  and  phases of the milled powders Phase Parameters Milling time (h) 50 70  D (nm) 22(2) 16(1) <a/a> (%) 2.76E-03 3.43E-03  D (nm) 15(2) 9(1) <a/a> (%) 2.04E-03 2.31E-03 Table 2 presents lattice parameters determined for individual phases and the corresponding ICDD data sheets. In the case of the  phase, a slight deviation of the a0 lattice parameters from the catalogue data was observed. A similar type of cell contraction was discovered for the a0 parameter of the  phase and for the c0 parameter, but these changes are smaller, resulting from the shift systems for the phases of the hexagonal system. It should be stressed that for the sample after 70 h of milling, the deviation of both lattice parameters can be noted. It is probably the effect of the alternate cold welding and deposition of the particles during the mechanical synthesis. Table 2: Lattice parameters and contents of  and  phases of the milled powders Phase Lattice parameters (nm) ICDD*  a0 0.2970 c0 0.4720  a0 0.3307 * International Centre for Diffraction Data® – a scientific organization dedicated to collecting, editing, publishing and distributing powder- diffraction data for the identification of materials The quantitative phase analysis of the material after each stage show a successive progressive synthesis of the as-prepared powders in relation to the  phase (77.5(11)) % mass fraction and 82.6(12) % mass fraction, for 50 h and 70 h milling times, respectively). The X-ray qualitative analysis showed that only  and  phases are present in the samples annealed at 1000 °C for 24 h from the powder previously milled for 50 h and 70 h, as shown in Figure 4. The XRD shows that there were no obvious diffraction peaks of elemental Nb remaining in the sintered Ti–26Nb (a/%). The diffraction peaks attributable to the  and  phases from the pattern confirm a duplex microstructure. M. Tahara et al.23 confirmed only the  phase for Ti-26Nb (a/%) produced by arc melting. The predominant phase in all the samples was the body-centered-cubic (bcc)  phase. This indicates that Nb was diffused into Ti, thereby lead- ing to the formation of  +  phases after the sintering. This is significant because two-phase Ti alloys, with the major  phase and the minor  phase, possess good com- prehensive properties including high yield strength, ex- cellent corrosion resistance and good fracture toughness. The presence of unalloyed Ti particles in materials may arise from different facts. According to S. N. Patankar and F. H. Froes24, it could be correlated with the brittleness of the Nb particles, as compared to the Ti particles, since these brittle particles adhere to the sur- faces of ductile particles. Another reason could be the limited solubility of Nb in -Ti. It is known that iso- morphous -stabilizer elements, such as Nb, exhibit slower diffusivities in Ti than the eutectoid elements of Fe, Co and Ni.25 The structural analysis of the diffraction patters obtained using the Rietveld method showed that the lattice parameters for all the tested samples become reduced. In comparison to the ICDD data, it was found that high-energy milling and annealing have the smallest, but significant effect on the reduction of the size of a unit cell of the  and  phases. The quantitative phase anal- ysis (Table 3) of the material after annealing showed a significant synthesis of the as-prepared powders in rela- tion to the  phase 96.6(12) w/% and 98.8(12) w/% for samples previously milled for 50 h and 70 h, respecti- vely). Table 3: Lattice parameters and contents of  and  phases after annealing (A) the samples at 1000 °C for 24 h from the powder previously milled for 50 h and 70 h Phase Lattice parameters (nm) Phase Contents (w/%) ICDD* Rietveld 50 h + A 70 h + A 50 h + A 70 h + A  a0 0.2970 0.2898(2) 0.2891 (2)  3.4 (6) 1.2 (4) c0 0.4720 0.4715(4) 0.4710 (4)  96.6 (12) 98.8 (12)  a0 0.3307 0.3287(2) 0.3286 (2) * International Centre for Diffraction Data®– a scientific organization dedicated to collecting, editing, publishing and distributing powder-diffraction data for the identification of materials Optical and SEM micrographs of the studied samples are presented in Figures 5 and 6. They clearly show the G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... 798 Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 4: X-ray diffraction patterns for the samples annealed (+A) at 1000 °C for 24 h from the powder previously milled for 50 h and 70 h changes in the morphology of the material after sintering with respect to various milling times. The common feature of the samples is a hierarchical structure. After the milling, large-sized particles caused the formation of a porous material. The pores were large and intercon- nected (Figures 5a and 5c) in spite of the applied pressure of 750 MPa during the isostatic pressing. In the case of the sample milled for 50 h (Figure 5b), an anomalous grain growth was observed. SEM micro- graphs (Figure 6) revealed further differences between the obtained microstructures. The powders ball-milled for a long time cannot move and deform effectively during the compaction. This may be primarily due to the effects of the decrease in the particle size, solid-solution strengthening and work hardening of the particles during the ball-milling process.26,27 Amongst these factors, the work-hardening effect appears to be the main deter- minant in governing the plastic behavior of the powder particles during the compaction. In the ball-milling process, work hardening is caused by the interaction of dislocations with each other through the continuous mechanical impact of the balls on the powder particles.16 A quantitative analysis of the microstructure was conducted including the determination of the cross-sec- tional area of the grains and the subgrains, as well as the average circularity of the subgrains of the microsections based on microscopic images, the results of which are presented in Tables 4 to 6. Significant changes between the samples in the average a of the grain cross-sectional area can also be observed. For the samples milled for 50 h and 70 h, the average section area of grains was 9.09 μm2 and 3.55 μm2, respectively (Table 4). A change in the time and rotations per minute during milling causes a difference in the parameter. Table 4: Results for the cross-sectional area of the grains for the samples milled for different times Parameter Sample Minimumvalue Maximum value Average value Std. dev. a (μm2) 50 h 0.51 267.57 9.09 12.68 70 h 0.51 32.35 3.55 3.45 Table 5: Results for the circularity of the grains for the samples milled for different times Parameter Sample Minimumvalue Maximum value Average value Std. dev. fg2 50 h 0.10 1.00 0.60 0.15 70 h 0.11 0.96 0.59 0.15 Table 6: Results for the dimensionless lengthening factor of the grains for the samples milled for different times Parameter Sample Minimumvalue Maximum value Average value Std. dev. fg1 50 h 0.32 3.31 1.06 0.37 70 h 0.28 3.30 1.06 0.37 Histograms of the size of the area of the section of the grains in the samples milled for 50 h and 70 h are presented in Figures 7 and 8, respectively. For the sam- ple milled for 70 h, a larger fraction of the grains with the section area in the range of 0.5–35 μm can be ob- served. Milling for 70 h caused a significant fragmen- tation of the grains with the section area below 30 μm2. G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 799 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: a), b) SEM micrographs of Ti-26Nb (a/%) after different milling times – 50 h and c), d) 70 h and annealing Figure 5: a), b) Optical micrographs of Ti-26Nb (a/%) at different magnifications, produced in different milling times – 50 h and c), d) 70 h and annealing Figure 7: Histogram of the grain average area for the sample milled for 50 h and annealed The circularity of grains was 0.60 for 50 h of milling and 0.59 for 70 h of milling, which corresponds to the dimensionless lengthening factor for both samples (Tables 5 to 6). The coefficient of the grain variability was also determined. For all the samples milled for 50 h, it was 1.4, and for the sample milled for 70 h, it was 0.9. It means that for the sample milled for a shorter time, the size of the cross-sectional area is much more diversified in comparison to the sample milled for 70 h. The longer milling time (70 h) caused a greater fragmentation of the material. The difference in the average section area of the grains was significant, especially when compared to the size of the powder after milling. No significant diffe- rences in the size of the particles depending on the milling time were noticed. The number of the elements that were tested per area unit of the image was also determined; for the samples annealed and previously milled for 50 h and 70 h, the number of elements per area was estimated to be 82379 [ ]1 2/mm and 199951 [ ]1 2/mm respectively. Distribution maps of the elements for the sintered samples revealed single regions rich in titanium in the case of the material milled for 50 h (Figure 9) Also, a slightly higher concentration of Ti was observed inside the grains. In contrast, milling for 70 h allowed us to obtain a material with an even distribution of the ele- ments on its surface (Figure 10). In the case of this alloy, the increase in the milling time allows a more homogeneous structure. Figures 11 and 12 show the effect of an analysis with the Hysitron Tribointender Ti950 with AFM QScope 250, which allowed us to compare the surface topogra- phies of the produced samples. For the material milled for 50 h and sintered, a higher diversification of the sample surface was observed. On the sample, initially synthetized for 70 h of milling, we can observe smaller and more regular grains in comparison to the sample ex- posed to the shorter milling time, which was confirmed with a stereological analysis. G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... 800 Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 8: Histograms of the grain average area for the sample milled for 70 h and annealed. The upper histogram presents the distribution of the average area of the grains smaller than 36 μm2 for this sample Figure 10: Distribution maps of the elements for the sample milled for 70 h and annealed Figure 9: Distribution maps of the elements for the sample milled for 50 h and annealed Another important aspect of the produced material is its properties. DSC results showed (Figure 13) the presence of an endothermic peak on the heating curve for both samples. In the case of the sample milled for 70 h, the peak was slightly shifted toward higher temperatures. The peaks occurred at about 490 °C and 520 °C for the samples milled for 50 h and 70 h, respectively. The pre- sence of the peaks can indicate the presence of a partial phase transformation   . Moreover, the milling time has an impact on the temperature of transformation. However, no significant changes were noticed on the cooling curves. The nanoindentation method allowed the determina- tion of the reduced Young’s modulus and hardness of the material, and the results of this analysis of the material after different milling times and sintering are presented in Table 7. Table 7: Results of the nanoindentation analysis – hardness and modulus of both samples Samples Contact depth(nm) Hardness (GPa) Modulus (GPa) 50 h 89(14) 0.99(0.31) 48(19) 70 h 42(12) 3.86(1.71) 95(32) The measurement is based on the curves of the load and the penetration depth of the indenter (Figures 14 and 15). After 50 h of milling, the material obtained much lower values of hardness and modulus in comparison to the material milled for 70 h. For the samples milled for 50 h and 70 h, the hardness was 0.99±0.31 GPa and 3.86±1.71GPa, respectively. A similar situation applied to the values of the modulus. For the material milled for 50 h, it was 48±19 GPa, and for the material milled for 70 h, it was 95±32 GPa. Such low modulus values, especially in the case of the sample with the shorter milling time, are very promising as they are close to the value of the modulus of a human bone. G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 801 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 13: DSC analysis of the samples milled for 50 h and 70 h and also annealed Figure 12: Surface morphology of the area of 40 μm × 40 μm for the sample milled for 70 h and annealed Figure 11: Surface morphology of the area of 40 μm × 40 μm for the sample milled for 50 h and annealed The circularity of grains was 0.60 for 50 h of milling and 0.59 for 70 h of milling, which corresponds to the dimensionless lengthening factor for both samples (Tables 5 to 6). The coefficient of the grain variability was also determined. For all the samples milled for 50 h, it was 1.4, and for the sample milled for 70 h, it was 0.9. It means that for the sample milled for a shorter time, the size of the cross-sectional area is much more diversified in comparison to the sample milled for 70 h. The longer milling time (70 h) caused a greater fragmentation of the material. The difference in the average section area of the grains was significant, especially when compared to the size of the powder after milling. No significant diffe- rences in the size of the particles depending on the milling time were noticed. The number of the elements that were tested per area unit of the image was also determined; for the samples annealed and previously milled for 50 h and 70 h, the number of elements per area was estimated to be 82379 [ ]1 2/mm and 199951 [ ]1 2/mm respectively. Distribution maps of the elements for the sintered samples revealed single regions rich in titanium in the case of the material milled for 50 h (Figure 9) Also, a slightly higher concentration of Ti was observed inside the grains. In contrast, milling for 70 h allowed us to obtain a material with an even distribution of the ele- ments on its surface (Figure 10). In the case of this alloy, the increase in the milling time allows a more homogeneous structure. Figures 11 and 12 show the effect of an analysis with the Hysitron Tribointender Ti950 with AFM QScope 250, which allowed us to compare the surface topogra- phies of the produced samples. For the material milled for 50 h and sintered, a higher diversification of the sample surface was observed. On the sample, initially synthetized for 70 h of milling, we can observe smaller and more regular grains in comparison to the sample ex- posed to the shorter milling time, which was confirmed with a stereological analysis. G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... 800 Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 8: Histograms of the grain average area for the sample milled for 70 h and annealed. The upper histogram presents the distribution of the average area of the grains smaller than 36 μm2 for this sample Figure 10: Distribution maps of the elements for the sample milled for 70 h and annealed Figure 9: Distribution maps of the elements for the sample milled for 50 h and annealed Another important aspect of the produced material is its properties. DSC results showed (Figure 13) the presence of an endothermic peak on the heating curve for both samples. In the case of the sample milled for 70 h, the peak was slightly shifted toward higher temperatures. The peaks occurred at about 490 °C and 520 °C for the samples milled for 50 h and 70 h, respectively. The pre- sence of the peaks can indicate the presence of a partial phase transformation   . Moreover, the milling time has an impact on the temperature of transformation. However, no significant changes were noticed on the cooling curves. The nanoindentation method allowed the determina- tion of the reduced Young’s modulus and hardness of the material, and the results of this analysis of the material after different milling times and sintering are presented in Table 7. Table 7: Results of the nanoindentation analysis – hardness and modulus of both samples Samples Contact depth(nm) Hardness (GPa) Modulus (GPa) 50 h 89(14) 0.99(0.31) 48(19) 70 h 42(12) 3.86(1.71) 95(32) The measurement is based on the curves of the load and the penetration depth of the indenter (Figures 14 and 15). After 50 h of milling, the material obtained much lower values of hardness and modulus in comparison to the material milled for 70 h. For the samples milled for 50 h and 70 h, the hardness was 0.99±0.31 GPa and 3.86±1.71GPa, respectively. A similar situation applied to the values of the modulus. For the material milled for 50 h, it was 48±19 GPa, and for the material milled for 70 h, it was 95±32 GPa. Such low modulus values, especially in the case of the sample with the shorter milling time, are very promising as they are close to the value of the modulus of a human bone. G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 801 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 13: DSC analysis of the samples milled for 50 h and 70 h and also annealed Figure 12: Surface morphology of the area of 40 μm × 40 μm for the sample milled for 70 h and annealed Figure 11: Surface morphology of the area of 40 μm × 40 μm for the sample milled for 50 h and annealed The nanosized indentation measurements of the mechanical properties are from the selected specific areas, which neutralize the effects of the pores or defects on the results. A large measurement uncertainty is even more surprising. Assuming that the material is mechani- cally widely heterogeneous, it seems necessary to apply heat treatment to the material. The microhardness measurement of the sample milled for 50 h revealed a higher value of the microhardness and it is 288.6(±31.0) HV0.5, which is around 2.82 GPa. For the sample milled for 70 h, the revealed value is 256.2(±14.6) HV0.5, which is approximately 2.51 GPa. However, we observed a relatively high uncertainty of the measurement, which made it difficult to clearly evaluate the differences in the microhardness measurement. 4 CONCLUSIONS Based on the investigation into the microstructure and mechanical properties of the Ti-26Nb (a/%) alloy, the main conclusions can be drawn: • It was proved that the microstructure, the degree of porosity and mechanical properties can be adjusted by means of mechanical alloying. • A longer milling time (70 h) resulted in an increased degree of structure fragmentation. Young’s modulus increased twice as well. • The porous samples sintered from the powders milled for 50 hours have a very low elastic modulus of 48(19) GPa, which is similar to that of natural bones. • Ball-milling for 50 h led to the formation of a rela- tively equiaxed shape and powder particles of a smaller size and with a more asymmetrical particle- size distribution than that of the powders ball milled for 70 h. • Based on XRD, it was found that the process of high-energy milling for 50 h and 70 h makes it pos- sible to obtain a solid nanocrystalline solution  + . Acknowledgments This work was supported by the Polish National Science Centre (Polish: Narodowe Centrum Nauki, abbr. NCN) under the research project no. UMO-2011/03/ D/ST8/04884 5 REFERENCES 1 M. Long, H. J. Rack, Titanium alloys in total joint replacement – a materials science perspective, Biomaterials, 19 (1998) 18, 1621–1639, doi:10.1016/S0142-9612(97)00146-4 2 Y. Li, C. Yang, H. Zhao, S. Qu, X. Li, Y. Li, New Developments of Ti-Based Alloys for Biomedical Applications, Mater., 7 (2014) 3, 1709–1800, doi:10.3390/ma7031709 3 M. Niinomi, Mechanical biocompatibilities of titanium alloys for biomedical application, J. Mech. Behav. Biomed. Mater., 1 (2008) 1, 30–42, doi:10.1016/j.jmbbm.2007.07.001 4 A. Biesiekierski, J. Wang, M. Gepreel, C. Wena, A new look at biomedical Ti-based shape memory alloys, Acta Biomater. 8 (2012) 5, 1661–1669, doi:10.1016/j.actbio.2012.01.018 5 S. Miyazaki, H.Y. Kim, Basic characteristics of titanium–nickel (Ti–Ni)-based and titanium–niobium (Ti–Nb)-based alloys, Shape Memory and Superelastic Alloys, Appl. Technol., (2011), 15–42, doi:10.1016/B978-1-84569-707-5.50002-X 6 M. Tahara, H. Y. Kim, H. Hosoda, S. Miyazaki, Cyclic deformation behavior of Ti-26 at.% Nb alloy, Acta Mater., 25 (2009), 2461–2469, doi:10.1016/j.actamat.2009.01.037 7 H. Y. Kim, Y. Ikehara, J. I. Kim, H. Hosoda, S. Miyazaki, Marten- sitic transformation, shape memory effect and superelasticity of Ti–Nb binary alloys, Acta Mater. 54 (2006), 2419–2429, doi:10.1016/j.actamat.2006.01.019 8 G. Dercz, I. Matu³a, M. Zubko, A. Liberska, Structure characteri- zation of biomedical Ti-Mo-Sn prepared by mechanical alloying method, Acta Phys. Pol. A, 130 (2016), 1029–1032, doi:10.12693/ APhysPolA.130.1029 9 I. Matu³a, G. Dercz, M. Zubko, L. Paj¹k, Influence of high energy milling time on the Ti-50Ta biomedical alloy structure, Acta Phys. Pol. A, 130 (2016), 1033–1036, doi:10.12693/APhysPolA.130.1033 10 X. Rao, C. L. Chu, Y. Y. Zheng, Phase composition, microstructure, and mechanical properties of porous Ti-Nb-Zr alloys prepared by a two-step foaming powder metallurgy method, J. Mech. Behav. Biomed., 34 (2014), 27–36, doi:10.1016/j.jmbbm.2014.02.001 11 G. Ryan, A. Pandit, D. P. Apatsidis, Fabrication methods of porous metals for use in orthopaedic applications, Biomaterials, 27 (2006) 13, 2651–2670, doi:10.1016/j.biomaterials.2005.12.002 G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... 802 Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 15: Diagram of reduced modulus to contact depth for both samples, after 50 h and 70 h of milling Figure 14: Diagram of hardness to contact depth for both samples, after 50 h and 70 h of milling 12 H. M. Rietveld, A Profile Refinement Method for Nuclear and Magnetic Structures, J. Appl. Cryst., 3 (1969), 65–69, doi:10.1107/ S0021889869006558 13 R. J. Hill, C. J. Howard, Quantitative phase analysis from neutron powder diffraction data using the Rietveld method. J. Appl. Cryst., 20 (1987), 467–474, doi:10.1107/S0021889887086199 14 G. Dercz, D. Oleszak, K. Prusik, L. Paj¹k, Rietveld-based quanti- tative analysis of multiphase powders with nanocrystalline NiAl and FeAl phases, Rev. Adv. Mater. Sci., 8 (2008), 764–768 15 Y. Li, Y. Cui, F. Zhanga, H. Xua, Shape memory behavior in Ti–Zr alloys, Scripta Mater., 64 (2011) 6, 584–587, doi:/10.1016/ j.scriptamat.2010.11.048 16 A. Nouri, P. D. Hodgson, C. Wen, Effect of ball-milling time on the structural characteristics of biomedical porous Ti–Sn–Nb alloy, Mater. Sci. Eng. C, 31 (2011), 921–928, doi:10.1016/j.msec.2011. 02.011 17 G. Dercz, I. Matu³a, M. Zubko, J. Dercz, Phase composition and microstructure of new Ti–Ta–Nb–Zr biomedical alloys prepared by mechanical alloying method, Powder Diffr., (2017), 1–7, doi:10.1017/S0885715617000045 18 L. Lü, M. O. Lai, Mechanical Alloying, Kluwer Academic Pub- lishers, Boston, 1998 19 C. Suryanarayana, Mechanical alloying and milling, Prog. Mater. Sci., 46 (2001) 1–2, 1–184, doi:10.1016/S0079-6425(99)00010-9 20 A. Omran, K. Woo, H. B. Lee, Mechanical properties of –Ti- 35Nb-2.5Sn alloy synthesized by mechanical alloying and pulsed current activated sintering, Metall. Mater. Trans. A, 43 (2012) 12, 4866–4874, doi:10.1007/s11661-012-1298-y 21 G. Dercz, B. Formanek, K. Prusik, L. Paj¹k, Microstructure of Ni(Cr)-TiC-Cr3C2-Cr7C3 composite powder, J. Mater. Process. Tech., 162 (2005), 15–19, doi:10.1016/jmatprotec.2005.02.004 22 G. Dercz, L. Paj¹k, B. Formanek, Dispersion analysis of NiAl- TiC-Al2O3 composite powder ground in a high-energy attritorial mill, J. Mater. Process. Tech., 175 (2006), 334–337, doi:10.1016/ j.jmatprotec.2005.04.060 23 M. Tahara, H. Y. Kim, H. Hosoda, S. Miyazaki, Cyclic deformation behavior of a Ti–26 at.% Nb alloy, Acta Mater., 57 (2009), 2461–2469, doi:10.1016/j.actamat.2009.01.037 24 E. K. Molchanova, Phase Diagrams of Titanium Alloys (Translation of Atlas Diagram Sostoyaniya Titanovyk Splavov), Israel Program for Scientific Translations, Jerusalem, 1965, 116 25 S. N. Patankar, F. H. Froes, Transformation of mechanically alloyed Nb-Sn powder to Nb3Sn, Metall. Mater. Trans. A, 35 (2004), 3009–3012, doi:10.1007/s11661-004-0248-8 26 P. J. James, Particle Deformation During Cold Isostatic Pressing of Metal Powders, Powder Metall., 20 (1977), 199–204, doi:10.1179/ pom.1977.20.4.199 27 L. Lu, M. O. Lai, G. Li, Influence of sintering process on the mecha- nical property and microstructure of ball milled composite compacts, Mater. Res. Bull., 31 (1996), 453–464, doi:10.1016/S0025-5408(96) 00024-4 G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 803 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS The nanosized indentation measurements of the mechanical properties are from the selected specific areas, which neutralize the effects of the pores or defects on the results. A large measurement uncertainty is even more surprising. Assuming that the material is mechani- cally widely heterogeneous, it seems necessary to apply heat treatment to the material. The microhardness measurement of the sample milled for 50 h revealed a higher value of the microhardness and it is 288.6(±31.0) HV0.5, which is around 2.82 GPa. For the sample milled for 70 h, the revealed value is 256.2(±14.6) HV0.5, which is approximately 2.51 GPa. However, we observed a relatively high uncertainty of the measurement, which made it difficult to clearly evaluate the differences in the microhardness measurement. 4 CONCLUSIONS Based on the investigation into the microstructure and mechanical properties of the Ti-26Nb (a/%) alloy, the main conclusions can be drawn: • It was proved that the microstructure, the degree of porosity and mechanical properties can be adjusted by means of mechanical alloying. • A longer milling time (70 h) resulted in an increased degree of structure fragmentation. Young’s modulus increased twice as well. • The porous samples sintered from the powders milled for 50 hours have a very low elastic modulus of 48(19) GPa, which is similar to that of natural bones. • Ball-milling for 50 h led to the formation of a rela- tively equiaxed shape and powder particles of a smaller size and with a more asymmetrical particle- size distribution than that of the powders ball milled for 70 h. • Based on XRD, it was found that the process of high-energy milling for 50 h and 70 h makes it pos- sible to obtain a solid nanocrystalline solution  + . Acknowledgments This work was supported by the Polish National Science Centre (Polish: Narodowe Centrum Nauki, abbr. NCN) under the research project no. UMO-2011/03/ D/ST8/04884 5 REFERENCES 1 M. Long, H. J. Rack, Titanium alloys in total joint replacement – a materials science perspective, Biomaterials, 19 (1998) 18, 1621–1639, doi:10.1016/S0142-9612(97)00146-4 2 Y. Li, C. Yang, H. Zhao, S. Qu, X. Li, Y. Li, New Developments of Ti-Based Alloys for Biomedical Applications, Mater., 7 (2014) 3, 1709–1800, doi:10.3390/ma7031709 3 M. Niinomi, Mechanical biocompatibilities of titanium alloys for biomedical application, J. Mech. Behav. Biomed. Mater., 1 (2008) 1, 30–42, doi:10.1016/j.jmbbm.2007.07.001 4 A. Biesiekierski, J. Wang, M. Gepreel, C. Wena, A new look at biomedical Ti-based shape memory alloys, Acta Biomater. 8 (2012) 5, 1661–1669, doi:10.1016/j.actbio.2012.01.018 5 S. Miyazaki, H.Y. Kim, Basic characteristics of titanium–nickel (Ti–Ni)-based and titanium–niobium (Ti–Nb)-based alloys, Shape Memory and Superelastic Alloys, Appl. Technol., (2011), 15–42, doi:10.1016/B978-1-84569-707-5.50002-X 6 M. Tahara, H. Y. Kim, H. Hosoda, S. Miyazaki, Cyclic deformation behavior of Ti-26 at.% Nb alloy, Acta Mater., 25 (2009), 2461–2469, doi:10.1016/j.actamat.2009.01.037 7 H. Y. Kim, Y. Ikehara, J. I. Kim, H. Hosoda, S. Miyazaki, Marten- sitic transformation, shape memory effect and superelasticity of Ti–Nb binary alloys, Acta Mater. 54 (2006), 2419–2429, doi:10.1016/j.actamat.2006.01.019 8 G. Dercz, I. Matu³a, M. Zubko, A. Liberska, Structure characteri- zation of biomedical Ti-Mo-Sn prepared by mechanical alloying method, Acta Phys. Pol. A, 130 (2016), 1029–1032, doi:10.12693/ APhysPolA.130.1029 9 I. Matu³a, G. Dercz, M. Zubko, L. Paj¹k, Influence of high energy milling time on the Ti-50Ta biomedical alloy structure, Acta Phys. Pol. A, 130 (2016), 1033–1036, doi:10.12693/APhysPolA.130.1033 10 X. Rao, C. L. Chu, Y. Y. Zheng, Phase composition, microstructure, and mechanical properties of porous Ti-Nb-Zr alloys prepared by a two-step foaming powder metallurgy method, J. Mech. Behav. Biomed., 34 (2014), 27–36, doi:10.1016/j.jmbbm.2014.02.001 11 G. Ryan, A. Pandit, D. P. Apatsidis, Fabrication methods of porous metals for use in orthopaedic applications, Biomaterials, 27 (2006) 13, 2651–2670, doi:10.1016/j.biomaterials.2005.12.002 G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... 802 Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 15: Diagram of reduced modulus to contact depth for both samples, after 50 h and 70 h of milling Figure 14: Diagram of hardness to contact depth for both samples, after 50 h and 70 h of milling 12 H. M. Rietveld, A Profile Refinement Method for Nuclear and Magnetic Structures, J. Appl. Cryst., 3 (1969), 65–69, doi:10.1107/ S0021889869006558 13 R. J. Hill, C. J. Howard, Quantitative phase analysis from neutron powder diffraction data using the Rietveld method. J. Appl. Cryst., 20 (1987), 467–474, doi:10.1107/S0021889887086199 14 G. Dercz, D. Oleszak, K. Prusik, L. Paj¹k, Rietveld-based quanti- tative analysis of multiphase powders with nanocrystalline NiAl and FeAl phases, Rev. Adv. Mater. Sci., 8 (2008), 764–768 15 Y. Li, Y. Cui, F. Zhanga, H. Xua, Shape memory behavior in Ti–Zr alloys, Scripta Mater., 64 (2011) 6, 584–587, doi:/10.1016/ j.scriptamat.2010.11.048 16 A. Nouri, P. D. Hodgson, C. Wen, Effect of ball-milling time on the structural characteristics of biomedical porous Ti–Sn–Nb alloy, Mater. Sci. Eng. C, 31 (2011), 921–928, doi:10.1016/j.msec.2011. 02.011 17 G. Dercz, I. Matu³a, M. Zubko, J. Dercz, Phase composition and microstructure of new Ti–Ta–Nb–Zr biomedical alloys prepared by mechanical alloying method, Powder Diffr., (2017), 1–7, doi:10.1017/S0885715617000045 18 L. Lü, M. O. Lai, Mechanical Alloying, Kluwer Academic Pub- lishers, Boston, 1998 19 C. Suryanarayana, Mechanical alloying and milling, Prog. Mater. Sci., 46 (2001) 1–2, 1–184, doi:10.1016/S0079-6425(99)00010-9 20 A. Omran, K. Woo, H. B. Lee, Mechanical properties of –Ti- 35Nb-2.5Sn alloy synthesized by mechanical alloying and pulsed current activated sintering, Metall. Mater. Trans. A, 43 (2012) 12, 4866–4874, doi:10.1007/s11661-012-1298-y 21 G. Dercz, B. Formanek, K. Prusik, L. Paj¹k, Microstructure of Ni(Cr)-TiC-Cr3C2-Cr7C3 composite powder, J. Mater. Process. Tech., 162 (2005), 15–19, doi:10.1016/jmatprotec.2005.02.004 22 G. Dercz, L. Paj¹k, B. Formanek, Dispersion analysis of NiAl- TiC-Al2O3 composite powder ground in a high-energy attritorial mill, J. Mater. Process. Tech., 175 (2006), 334–337, doi:10.1016/ j.jmatprotec.2005.04.060 23 M. Tahara, H. Y. Kim, H. Hosoda, S. Miyazaki, Cyclic deformation behavior of a Ti–26 at.% Nb alloy, Acta Mater., 57 (2009), 2461–2469, doi:10.1016/j.actamat.2009.01.037 24 E. K. Molchanova, Phase Diagrams of Titanium Alloys (Translation of Atlas Diagram Sostoyaniya Titanovyk Splavov), Israel Program for Scientific Translations, Jerusalem, 1965, 116 25 S. N. Patankar, F. H. Froes, Transformation of mechanically alloyed Nb-Sn powder to Nb3Sn, Metall. Mater. Trans. A, 35 (2004), 3009–3012, doi:10.1007/s11661-004-0248-8 26 P. J. James, Particle Deformation During Cold Isostatic Pressing of Metal Powders, Powder Metall., 20 (1977), 199–204, doi:10.1179/ pom.1977.20.4.199 27 L. Lu, M. O. Lai, G. Li, Influence of sintering process on the mecha- nical property and microstructure of ball milled composite compacts, Mater. Res. Bull., 31 (1996), 453–464, doi:10.1016/S0025-5408(96) 00024-4 G. DERCZ, I. MATU£A: EFFECT OF BALL MILLING ON THE PROPERTIES OF THE POROUS Ti–26Nb ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 795–803 803 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... 805–811 EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL PROPERTIES AND EROSIVE-WEAR BEHAVIOR OF A WOOD-DUST-PARTICLE-REINFORCED PHENOL FORMALDEHYDE COMPOSITE VPLIVI DODATKA KOKOSOVIH VLAKEN FENOL-FORMALDEHIDNEMU KOMPOZITU, OJA^ANEM Z LESNIM PRAHOM, NA NJEGOVE MEHANSKE LASTNOSTI IN EROZIJSKO OBRABO Arul Sujin Jose1, Ayyanar Athijayamani2, Kalimuthu Ramanathan3, Susaiyappan Sidhardhan4 1Lourdes Mount College of Engineering and Technology, Department of Mechanical Engineering, Kanyakumari, Tamilnadu, India 2Government College of Engineering, Department of Mechanical Engineering, Bodinayakkanur, Tamilnadu, India 3Alagappa Chettiar College of Engineering and Technology, Department of Mechanical Engineering, Karaikudi, Tamilnadu, India 4Government College of Engineering, Department of Civil Engineering, Tirunelveli, Tamilnadu, India athimania@gmail.com Prejem rokopisa – received: 2016-09-23; sprejem za objavo – accepted for publication: 2017-01-22 doi:10.17222/mit.2016.284 Several attempts were made to investigate the effects of various process parameters on the mechanical properties and wear behavior of synthetic and natural cellulosic fibers and also particle-reinforced polymer composites. However, very few studies were carried out on the effects of various process parameters on the mechanical and wear behavior of phenol formaldehyde (PF) composites reinforced with natural cellulosic fibers and particles. Therefore, in the present study, an attempt was made to observe the effects of various process parameters on the mechanical and wear behavior of wood-dust (WD) and coir-pith (CP) particle-reinforced resole-type PF composites. First, the mechanical properties of a WD/PF composite were studied based on the content of CP particles. Then, the erosive-wear behavior of the WD/PF composite was studied with respect to five different parameters such particle content, erodent size, impact velocity, impingement angle, and standoff distance. The erosive experiments were carried out for five different parameters based on the Taguchi experimental design (L27). The results show that the mechanical properties of the WD/PF composite increase with an addition of CP particles. The increment in the composite modulus was higher than that of the composite strength. The erosive test results indicate that the erosion-wear rate is affected by the particle content, impingement angle, erodent size and impact velocity. Brittle-erosion behavior was identified on the surface of the composite with a heavy erosive wear occurring at a 60° impingement angle. Keywords: biowaste particles, phenol formaldehyde, composites, mechanical properties, erosive-wear resistance, Taguchi method Izvedenih je bilo `e kar nekaj poizkusov v zvezi z u~inki razli~nih procesnih parametrov na mehanske lastnosti in obrabo polimernih kompozitov oja~anih s sinteti~nimi in naravnimi celuloznimi vlakni in/ali delci. Toda zelo malo raziskav je bilo izvedenih glede vpliva razli~nih procesnih parametrov na mehanske lastnosti in obrabo fenolformaldehidnih (angl. PF) kompozitov, oja~anih z naravnimi celuloznimi vlakni in delci. Tako je v pri~ujo~em delu predstavljen vpliv razli~nih procesnih parametrov na mehanske lastnosti in obrabo PF kompozitov, ki so bili oja~ani z delci lesnega prahu (angl. WD) in delci kokosa (angl. CP). Te vrste kompozitov se uporabljajo za izdelavo podplatov ~evljev. Najprej so bile dolo~ene mehanske lastnosti WD/PF kompozitov glede na vsebnost CP delcev. Sledili so preizkusi in analize erozijske obrabe WD/PF kompozitov glede na vsebnost (koli~ino) delcev v kompozitu, velikost, hitrost in razdaljo u~inkovanja erozijskega sredsta ter njegov vpadni kot. Preizkusi so temeljili na analizi s Taguchijevo metodo (L27) s petimi razli~nimi parametri. Rezultati so pokazali, da se mehanske lastnosti WD/PF kompozitov izbolj{ujejo z dodajanjem CP delcev. Povi{anje modula kompozitov je bilo ve~je od pove~anja trdnosti kompozita. Erozijski testi ka`ejo, da je hitrost erozijske obrabe posledica vseh procesnih parametrov, to je: vsebnosti delcev, udarnega kota, hitrosti in velikosti delcev izbranega erozijskega sredstva. Najve~ja obraba zaradi erozije je bila dose`ena (ugotovljena) pri 60 stopinjskem vpadnem kotu abrazijskega sredstva z nastalimi po{kodbami krhkega zna~aja. Klju~ne besede: delci bioodpadkov, fenolformaldehid, kompoziti, mehanske lastnosti, odpornost proti erozijski obrabi, Taguchi metoda 1 INTRODUCTION Recently, polymer composites reinforced with syn- thetic materials have been replaced with polymer com- posites reinforced with bio-based natural materials. The bio-based natural materials have many advantages over the synthetic materials like renewability, biodegrada- bility, abundant availability, low costs, etc.1–3 In India, particularly in the South Indian region, the biowaste materials like wood dust, coir pith, groundnut shell, coconut shell, cashew nut shell, etc. are abundantly available because in that region, coconut, groundnut and cashew nut are cultivated in large amounts. A number of timber and oil mills are also available in the Southern region of Tamilnadu, India. Therefore, bioparticles are thrown away after producing useful materials and Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 805 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 67.017:620.1:620.163 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)805(2017) A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... 805–811 EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL PROPERTIES AND EROSIVE-WEAR BEHAVIOR OF A WOOD-DUST-PARTICLE-REINFORCED PHENOL FORMALDEHYDE COMPOSITE VPLIVI DODATKA KOKOSOVIH VLAKEN FENOL-FORMALDEHIDNEMU KOMPOZITU, OJA^ANEM Z LESNIM PRAHOM, NA NJEGOVE MEHANSKE LASTNOSTI IN EROZIJSKO OBRABO Arul Sujin Jose1, Ayyanar Athijayamani2, Kalimuthu Ramanathan3, Susaiyappan Sidhardhan4 1Lourdes Mount College of Engineering and Technology, Department of Mechanical Engineering, Kanyakumari, Tamilnadu, India 2Government College of Engineering, Department of Mechanical Engineering, Bodinayakkanur, Tamilnadu, India 3Alagappa Chettiar College of Engineering and Technology, Department of Mechanical Engineering, Karaikudi, Tamilnadu, India 4Government College of Engineering, Department of Civil Engineering, Tirunelveli, Tamilnadu, India athimania@gmail.com Prejem rokopisa – received: 2016-09-23; sprejem za objavo – accepted for publication: 2017-01-22 doi:10.17222/mit.2016.284 Several attempts were made to investigate the effects of various process parameters on the mechanical properties and wear behavior of synthetic and natural cellulosic fibers and also particle-reinforced polymer composites. However, very few studies were carried out on the effects of various process parameters on the mechanical and wear behavior of phenol formaldehyde (PF) composites reinforced with natural cellulosic fibers and particles. Therefore, in the present study, an attempt was made to observe the effects of various process parameters on the mechanical and wear behavior of wood-dust (WD) and coir-pith (CP) particle-reinforced resole-type PF composites. First, the mechanical properties of a WD/PF composite were studied based on the content of CP particles. Then, the erosive-wear behavior of the WD/PF composite was studied with respect to five different parameters such particle content, erodent size, impact velocity, impingement angle, and standoff distance. The erosive experiments were carried out for five different parameters based on the Taguchi experimental design (L27). The results show that the mechanical properties of the WD/PF composite increase with an addition of CP particles. The increment in the composite modulus was higher than that of the composite strength. The erosive test results indicate that the erosion-wear rate is affected by the particle content, impingement angle, erodent size and impact velocity. Brittle-erosion behavior was identified on the surface of the composite with a heavy erosive wear occurring at a 60° impingement angle. Keywords: biowaste particles, phenol formaldehyde, composites, mechanical properties, erosive-wear resistance, Taguchi method Izvedenih je bilo `e kar nekaj poizkusov v zvezi z u~inki razli~nih procesnih parametrov na mehanske lastnosti in obrabo polimernih kompozitov oja~anih s sinteti~nimi in naravnimi celuloznimi vlakni in/ali delci. Toda zelo malo raziskav je bilo izvedenih glede vpliva razli~nih procesnih parametrov na mehanske lastnosti in obrabo fenolformaldehidnih (angl. PF) kompozitov, oja~anih z naravnimi celuloznimi vlakni in delci. Tako je v pri~ujo~em delu predstavljen vpliv razli~nih procesnih parametrov na mehanske lastnosti in obrabo PF kompozitov, ki so bili oja~ani z delci lesnega prahu (angl. WD) in delci kokosa (angl. CP). Te vrste kompozitov se uporabljajo za izdelavo podplatov ~evljev. Najprej so bile dolo~ene mehanske lastnosti WD/PF kompozitov glede na vsebnost CP delcev. Sledili so preizkusi in analize erozijske obrabe WD/PF kompozitov glede na vsebnost (koli~ino) delcev v kompozitu, velikost, hitrost in razdaljo u~inkovanja erozijskega sredsta ter njegov vpadni kot. Preizkusi so temeljili na analizi s Taguchijevo metodo (L27) s petimi razli~nimi parametri. Rezultati so pokazali, da se mehanske lastnosti WD/PF kompozitov izbolj{ujejo z dodajanjem CP delcev. Povi{anje modula kompozitov je bilo ve~je od pove~anja trdnosti kompozita. Erozijski testi ka`ejo, da je hitrost erozijske obrabe posledica vseh procesnih parametrov, to je: vsebnosti delcev, udarnega kota, hitrosti in velikosti delcev izbranega erozijskega sredstva. Najve~ja obraba zaradi erozije je bila dose`ena (ugotovljena) pri 60 stopinjskem vpadnem kotu abrazijskega sredstva z nastalimi po{kodbami krhkega zna~aja. Klju~ne besede: delci bioodpadkov, fenolformaldehid, kompoziti, mehanske lastnosti, odpornost proti erozijski obrabi, Taguchi metoda 1 INTRODUCTION Recently, polymer composites reinforced with syn- thetic materials have been replaced with polymer com- posites reinforced with bio-based natural materials. The bio-based natural materials have many advantages over the synthetic materials like renewability, biodegrada- bility, abundant availability, low costs, etc.1–3 In India, particularly in the South Indian region, the biowaste materials like wood dust, coir pith, groundnut shell, coconut shell, cashew nut shell, etc. are abundantly available because in that region, coconut, groundnut and cashew nut are cultivated in large amounts. A number of timber and oil mills are also available in the Southern region of Tamilnadu, India. Therefore, bioparticles are thrown away after producing useful materials and Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 805 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 67.017:620.1:620.163 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)805(2017) dumped on the land of the village nearest to these in- dustries. Many studies have already reported on the properties of different bio-based natural-fiber-reinforced polymer composites in different conditions.3–7 But, only few reports are available on the properties of polymer com- posites filled with biowaste particles.8–11 In this investi- gation, an attempt was made to study the mechanical and wear behavior of PF composites reinforced with biowaste particles (WD and CP). Mechanical properties of wood-dust-particle-reinforced PF composites were evaluated based on the content of coir pith. The erosive- wear behavior of the composites was studied using five different parameters such as particle content, erodent size, impingement angle, impact velocity and standoff distance. The erosive experiments were conducted as per the Taguchi experimental design. The parameters used for erosive-wear tests were also analyzed using an analysis of variance with the wear rate. 2 MATERIALS AND METHODOLOGY 2.1 Materials Wood-dust particles were collected from the Kumar Timber and Sawmill, Karaikudi, Tamilnadu, India. Coir-pith particles were collected from the Coir Industry, Sozhavanthan, Tamilnadu, India. From the collected wood-dust and CP particles, microparticles with the average size of 800 microns were separated using a sieving machine available in our composite laboratory. The resole-type PF liquid resin was procured, together with a cross-linking agent (divinylbenzene) and acidic catalyst (hydrochloric acid), from POOJA Chemicals, Madurai, Tamilnadu, India. 2.2 Preparation of the composites A hardboard mold box with dimensions of 150 mm × 150 mm × 3 mm was used to prepare the wood-dust and coir-pith-particle composite plates using the hand lay-up technique. Wood dust/coir pith/phenol formaldehyde composites were fabricated at three different concentra- tions of wood-dust and coir-pith particles, i.e., (20, 30 and 40) % mass fractions. The amount of WD particles was maintained at a fixed level of 20 % mass fraction. Three different amounts of CP particles (0–20 % mass frac- tions) were hybridized with the constant amount of WD particles, i.e., 20WD/0CP, 20WD/10CP, 20WD/20CP. The weight percentage of WD and CP particles and designation of the composites are given in Table 1. Prior to the process, the particles were dried in sunlight for 12 h. The PF resin with the particles was mixed with a mechanical stirrer at room temperature for 30 min. Then, the cross-linking agent and acidic catalyst were also mixed into the mixture of phenol formaldehyde/particles and once again stirred with the mechanical stirrer for 15 min. After that, the mixture was poured into the mold box and allowed to cure at room temperature for 48 h. 2.3 Testing composite specimens Composite specimens were characterized using me- chanical tests such as tensile, flexural and impact tests. The tensile tests were conducted on an FIE universal testing machine (UTE 40 HGFL) in accordance with ASTM D638-10.12 The flexural tests were performed on the same testing machine in accordance with ASTM D790-10.13 The impact tests were carried out on an Izod impact machine according to ISO 180.14 All the tests were conducted at room temperature and atmospheric pressure. 2.4 Taguchi experimental design The erosive behavior of the WD/CP/PF composite was studied based on the Taguchi method and analysis of variance techniques. Experiments were performed as per Taguchi experimental design (an orthogonal array) because it is a systematic and efficient approach to get the optimum range of process parameters with a good performance. The number of experiments can be reduced due to the constructed orthogonal array, which provides a set of well-balanced experiments.15 The results obtained with this experimental design are transformed into signal-to-noise (S/N) ratios, which serve as objective functions for the optimization of parameters and help with the result analysis. There are three S/N ratios available for the optimization of several static problems: the smaller-the-better (used to minimize the response), the nominal-the-better (used whenever an ideal quality is equated with a particular nominal value.) and the larger- the-better ratio (used to maximize the response). Among these three characteristics, the minimum erosion rate comes under the smaller-the-better characteristic, which can be expressed as Equation (1): S/N = –10 Log10 (1) (the mean of the sum of squares of the measured data) A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... 806 Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Table 1: The weight percentage of WD and CP particles and designation of the composites Total weight percentage of particles in the composites Weight percentage of resin Weight percentage of WD particles Weight percentage of CP particles Designation of composites 20 80 20 0 20WD/0CP 30 70 20 10 20WD/10CP 40 60 20 20 20WD/20CP The five different process parameters at three levels are used in this study to observe the erosive behavior of the WD/CP/PF composite. Therefore, the actual number of experiments, based on the traditional experimental design, should be 243 (35). But, this number is reduced to 27 experiments using the Taguchi technique. The pro- cess parameters and their setting levels for the erosion test of the WD/CP/PF composite are presented in Table 2. In these experiments, the following parameters are fixed throughout the process: the type of erodent is silica, the erodent feed rate is 10.0±1.0 g/min, the nozzle length is 80 mm, and the nozzle diameter is 3 mm. Table 2: The erosive process parameters with their designation and setting levels Process Parameters and their designation Level I Level II Level III Particle content: (A) wt% 20 30 40 Impact velocity: (B) m/sec 41 52 63 Impingement angle: (C) degree 30 60 90 Erodent size: (D) ìm 300 500 700 Stand-off distance: (E) mm 80 120 160 2.5 Erosion test The erosive tests of the WD/CP/PF composite speci- mens were conducted as shown in the schematic diagram of the erosion process (Figure 1). The main components of the erosion-test apparatus are the erodent feeder box, erodent feeder nozzle, mixing chamber, nozzle of the mixing chamber, air-flow vent, sample holder and ero- dent collector. Dry silica sand with three different sizes (300, 500 and 700) μm was used as the erodent in the erosion tests. After the test, the composite samples were taken from the apparatus and cleaned with acetone. Then, the cleaned composite specimens were dried and weighed using a precision digital balance at an accuracy of ±0.1 mg. The composite samples were weighed before and after the erosion tests and their difference is termed as the weight loss. Then, the weight loss was recorded and used for the erosion-rate calculation. Generally, the erosion rate can be obtained as the ratio of the weight loss of samples to the weight of the eroding particle. The process was repeated until the steady-state erosion was reached. 3 RESULTS AND DISCUSSION 3.1 Mechanical properties of the composites Mechanical tests were carried out on the WD/CP/PF composites and their results are presented in Figure 2a. The neat-resin sample had a tensile strength of 29.8 MPa, tensile modulus of 1168.4 MPa, flexural strength of 34.7 MPa, flexural modulus of 1257.4 MPa, and impact strength of 1.24 KJ/m2. It can be seen that the tensile strength and modulus of the PF composite in- crease with an increase in the particle content. The tensile strength of the 20WD/PF composite is almost the same as that of the neat-resin sample. It shows that the addition of WD particles enhances the strength of the PF composite. The WD/PF composite without the addition of CP particles has a tensile strength of 30.4 MPa and this value increases to 41.7 MPa with the incorporation of 10 % mass fraction of CP particles; after that, it de- creases to 36.8 MPa with the addition of 20 % mass fraction of CP particles. This may be due to a poor interfacial bonding between the particles and the matrix, i.e., a weak transfer of stress. Moreover, the stress con- centration in the PF matrix may be created due to the corner edges of the irregularly shaped WD and CP parti- cles. Due to the addition of 10 % mass fraction and 20 % mass fraction of CP particles, the tensile strength of the WD/PF composite increases by about 37.17 % and 21.1 %, respectively. Figure 2a also shows the tensile- modulus values of the WD/CP/PF composites with res- pect to the particle content. The composite also reached the tensile-modulus value of the neat-resin sample with the particle addition of 20 % mass fraction. The tensile- modulus value of the WD/PF composite increased with the further addition of CP particles. The maximum mo- dulus value was observed at 40 % mass fraction of the particles. The results of the flexural tests of the WD/CP/PF composites with respect to the particle content are given in Figure 2b. It is interesting to note that the flexural strength and modulus of the WD/PF composite increase with the addition of CP particles. The flexural strength of the WD/PF composite is slightly lower than the value of the neat-resin sample. The maximum values of the flexu- ral strength and modulus were identified at the 40 % addition. The flexural strength of the WD/PF composite A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 807 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 1: Schematic diagram of the erosive process of the WD/CP/PF composite dumped on the land of the village nearest to these in- dustries. Many studies have already reported on the properties of different bio-based natural-fiber-reinforced polymer composites in different conditions.3–7 But, only few reports are available on the properties of polymer com- posites filled with biowaste particles.8–11 In this investi- gation, an attempt was made to study the mechanical and wear behavior of PF composites reinforced with biowaste particles (WD and CP). Mechanical properties of wood-dust-particle-reinforced PF composites were evaluated based on the content of coir pith. The erosive- wear behavior of the composites was studied using five different parameters such as particle content, erodent size, impingement angle, impact velocity and standoff distance. The erosive experiments were conducted as per the Taguchi experimental design. The parameters used for erosive-wear tests were also analyzed using an analysis of variance with the wear rate. 2 MATERIALS AND METHODOLOGY 2.1 Materials Wood-dust particles were collected from the Kumar Timber and Sawmill, Karaikudi, Tamilnadu, India. Coir-pith particles were collected from the Coir Industry, Sozhavanthan, Tamilnadu, India. From the collected wood-dust and CP particles, microparticles with the average size of 800 microns were separated using a sieving machine available in our composite laboratory. The resole-type PF liquid resin was procured, together with a cross-linking agent (divinylbenzene) and acidic catalyst (hydrochloric acid), from POOJA Chemicals, Madurai, Tamilnadu, India. 2.2 Preparation of the composites A hardboard mold box with dimensions of 150 mm × 150 mm × 3 mm was used to prepare the wood-dust and coir-pith-particle composite plates using the hand lay-up technique. Wood dust/coir pith/phenol formaldehyde composites were fabricated at three different concentra- tions of wood-dust and coir-pith particles, i.e., (20, 30 and 40) % mass fractions. The amount of WD particles was maintained at a fixed level of 20 % mass fraction. Three different amounts of CP particles (0–20 % mass frac- tions) were hybridized with the constant amount of WD particles, i.e., 20WD/0CP, 20WD/10CP, 20WD/20CP. The weight percentage of WD and CP particles and designation of the composites are given in Table 1. Prior to the process, the particles were dried in sunlight for 12 h. The PF resin with the particles was mixed with a mechanical stirrer at room temperature for 30 min. Then, the cross-linking agent and acidic catalyst were also mixed into the mixture of phenol formaldehyde/particles and once again stirred with the mechanical stirrer for 15 min. After that, the mixture was poured into the mold box and allowed to cure at room temperature for 48 h. 2.3 Testing composite specimens Composite specimens were characterized using me- chanical tests such as tensile, flexural and impact tests. The tensile tests were conducted on an FIE universal testing machine (UTE 40 HGFL) in accordance with ASTM D638-10.12 The flexural tests were performed on the same testing machine in accordance with ASTM D790-10.13 The impact tests were carried out on an Izod impact machine according to ISO 180.14 All the tests were conducted at room temperature and atmospheric pressure. 2.4 Taguchi experimental design The erosive behavior of the WD/CP/PF composite was studied based on the Taguchi method and analysis of variance techniques. Experiments were performed as per Taguchi experimental design (an orthogonal array) because it is a systematic and efficient approach to get the optimum range of process parameters with a good performance. The number of experiments can be reduced due to the constructed orthogonal array, which provides a set of well-balanced experiments.15 The results obtained with this experimental design are transformed into signal-to-noise (S/N) ratios, which serve as objective functions for the optimization of parameters and help with the result analysis. There are three S/N ratios available for the optimization of several static problems: the smaller-the-better (used to minimize the response), the nominal-the-better (used whenever an ideal quality is equated with a particular nominal value.) and the larger- the-better ratio (used to maximize the response). Among these three characteristics, the minimum erosion rate comes under the smaller-the-better characteristic, which can be expressed as Equation (1): S/N = –10 Log10 (1) (the mean of the sum of squares of the measured data) A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... 806 Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Table 1: The weight percentage of WD and CP particles and designation of the composites Total weight percentage of particles in the composites Weight percentage of resin Weight percentage of WD particles Weight percentage of CP particles Designation of composites 20 80 20 0 20WD/0CP 30 70 20 10 20WD/10CP 40 60 20 20 20WD/20CP The five different process parameters at three levels are used in this study to observe the erosive behavior of the WD/CP/PF composite. Therefore, the actual number of experiments, based on the traditional experimental design, should be 243 (35). But, this number is reduced to 27 experiments using the Taguchi technique. The pro- cess parameters and their setting levels for the erosion test of the WD/CP/PF composite are presented in Table 2. In these experiments, the following parameters are fixed throughout the process: the type of erodent is silica, the erodent feed rate is 10.0±1.0 g/min, the nozzle length is 80 mm, and the nozzle diameter is 3 mm. Table 2: The erosive process parameters with their designation and setting levels Process Parameters and their designation Level I Level II Level III Particle content: (A) wt% 20 30 40 Impact velocity: (B) m/sec 41 52 63 Impingement angle: (C) degree 30 60 90 Erodent size: (D) ìm 300 500 700 Stand-off distance: (E) mm 80 120 160 2.5 Erosion test The erosive tests of the WD/CP/PF composite speci- mens were conducted as shown in the schematic diagram of the erosion process (Figure 1). The main components of the erosion-test apparatus are the erodent feeder box, erodent feeder nozzle, mixing chamber, nozzle of the mixing chamber, air-flow vent, sample holder and ero- dent collector. Dry silica sand with three different sizes (300, 500 and 700) μm was used as the erodent in the erosion tests. After the test, the composite samples were taken from the apparatus and cleaned with acetone. Then, the cleaned composite specimens were dried and weighed using a precision digital balance at an accuracy of ±0.1 mg. The composite samples were weighed before and after the erosion tests and their difference is termed as the weight loss. Then, the weight loss was recorded and used for the erosion-rate calculation. Generally, the erosion rate can be obtained as the ratio of the weight loss of samples to the weight of the eroding particle. The process was repeated until the steady-state erosion was reached. 3 RESULTS AND DISCUSSION 3.1 Mechanical properties of the composites Mechanical tests were carried out on the WD/CP/PF composites and their results are presented in Figure 2a. The neat-resin sample had a tensile strength of 29.8 MPa, tensile modulus of 1168.4 MPa, flexural strength of 34.7 MPa, flexural modulus of 1257.4 MPa, and impact strength of 1.24 KJ/m2. It can be seen that the tensile strength and modulus of the PF composite in- crease with an increase in the particle content. The tensile strength of the 20WD/PF composite is almost the same as that of the neat-resin sample. It shows that the addition of WD particles enhances the strength of the PF composite. The WD/PF composite without the addition of CP particles has a tensile strength of 30.4 MPa and this value increases to 41.7 MPa with the incorporation of 10 % mass fraction of CP particles; after that, it de- creases to 36.8 MPa with the addition of 20 % mass fraction of CP particles. This may be due to a poor interfacial bonding between the particles and the matrix, i.e., a weak transfer of stress. Moreover, the stress con- centration in the PF matrix may be created due to the corner edges of the irregularly shaped WD and CP parti- cles. Due to the addition of 10 % mass fraction and 20 % mass fraction of CP particles, the tensile strength of the WD/PF composite increases by about 37.17 % and 21.1 %, respectively. Figure 2a also shows the tensile- modulus values of the WD/CP/PF composites with res- pect to the particle content. The composite also reached the tensile-modulus value of the neat-resin sample with the particle addition of 20 % mass fraction. The tensile- modulus value of the WD/PF composite increased with the further addition of CP particles. The maximum mo- dulus value was observed at 40 % mass fraction of the particles. The results of the flexural tests of the WD/CP/PF composites with respect to the particle content are given in Figure 2b. It is interesting to note that the flexural strength and modulus of the WD/PF composite increase with the addition of CP particles. The flexural strength of the WD/PF composite is slightly lower than the value of the neat-resin sample. The maximum values of the flexu- ral strength and modulus were identified at the 40 % addition. The flexural strength of the WD/PF composite A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 807 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 1: Schematic diagram of the erosive process of the WD/CP/PF composite was increased by about 68.99 % due to the incorporation of 20 % mass fraction of CP particles. The impact-strength values of the WD/CP/PF com- posites after the impact tests are presented in Figure 2c. It can be seen that the impact strength of the WD/PF composite is slightly lower than the value of the neat- resin sample. It is also shown that the impact strength of the WD/PF composite increases with the addition of 10 % mass fraction of CP particles, but it decreases with the incorporation of 20 % mass fraction of CP particles. This may be due to a poor adhesion between the particles and the matrix. It may also be due to the stress concentration of the resin matrix. It is also observed that the incor- poration of 10 % mass fraction and 20 % mass fraction of CP particles shows 15.57 % and 8.19 % higher impact values compared to the WD/PF composite. 3.2 Steady-state erosion: effects of the impingement angle Generally, the erosive-wear behavior of polymer composite materials can be categorized as brittle and ductile. A ductile erosive situation is created in thermo- plastic polymer composites, whereas the a brittle erosive situation may be created in thermosetting polymer composites. For the steady-state-erosion analysis of the WD/CP/PF composites, an erosion test was carried out based on eight different impingement angles (20, 30, 40, 50, 60, 70, 80 and 90)°, keeping all the other process parameters constant (the initial level values). The effects of the impingement angles on the erosion rate of the WD/CP/PF composites are presented in Fig- ure 3. From this figure, it can be observed that the erosion rate is high at the impingement angle of 60° for all the composite specimens, irrespective of the particle content. However, a more brittle erosive behavior was identified for the 40 % mass-fraction (20WD/20CP) composite specimen. However, in the 20 % and 30 % mass-fraction composite specimens, a semi-brittle erosive behavior was identified. This may be due to the addition of WD and CP particles to the PF composites. When the higher amounts of particles are added to the polymer material, it behaves as a typical brittle material. Therefore, the brittleness of the composites with 20 % mass fraction and 30 % mass fraction of particles is lower than the composite with 40 % mass fraction of particles. Due to this, a semi-brittle erosive situation exists during the erosive process. It is also clear from Figure 3 that the erosion rate increased with the increase in the particle content. This may be due to the increased hardness of the PF composite material caused by the addition of WD and CP particles. 3.3 Analysis of the erosion rate The erosion rates for 27 combinations of the erosive experiments conducted on the WD/CP/PF composites are given in Table 3. The erosion analysis was made with popular software, namely, MINITAB 17. From Table 3, it can be concluded that the parameter combina- tion of particle loading-A (level II = 30 % mass fraction), impact velocity-B (level I = 41 m/s), impingement angle-C (level I = 30°), erodent size-D (level II = 500 um) and standoff distance-E (160 mm) gives the mini- mum erosion rate (189. 8 mg/kg). Moreover, another parameter combination (experiment number 4) allows the next level of the minimum erosion rate (194.1 mg/kg). The difference between these two erosion rates is small, as seen from Table 3. Anyway, the first para- meter combination mentioned above is recognized as the better combination of parameters to obtain the minimum erosion rate. The overall mean of the signal-to-noise ratio for the erosion rate is found to be 49.75 dB. A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... 808 Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: Steady-state erosive behavior of WD/CP/PF composites for eight different impingement angles Figure 2: Variation of: a) tensile property, b) flexural property, and c) impact strength based on the particle content The effects of five erosive-process parameters on the erosion rate are graphically presented in Figure 4a. From this figure, it can be clearly concluded that para- meter A (the particle content), parameter B (the impact velocity) and parameter C (the impingement angle) are the most significant parameters. Parameter E (the stan- doff distance) shows a moderately significant influence, while parameter D (the erodent size) has a relatively less significant influence. Figure 4b shows the interaction between the erosive parameters. From this figure, it is observed that a moderate interaction exists between para- meters A and B, and between A and C. The interaction between parameters B and C is below the moderate level. Figures 5a to 5c show 3D surface plots of the ero- sion rate with significant process parameters. The observation is similar to the one made of the interaction plots of the erosion rate. From the erosion test analysis of the WD/CP/PF composites, it can be concluded that the erodent size is most insignificant for the erosion rate. The standoff distance shows relatively less significance when compared to the other three process parameters (particle content, impact velocity and impingement A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 809 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: 3D surface plots of erosion rate vs process parameters: a) A × B, b) A × C, and c) B × C Figure 4: a) Effects of erosive-process parameters and b) effects of interactions of erosive-process parameters on the erosion rate Table 3: The erosion rates and their S/N ratio of WD/CP/PF compo- sites for 27 combinations Experi ment No. A B C D E Erosion rate mg/kg S/N ratio dB 1 20 41 30 300 80 209.5 -46.42 2 20 41 60 500 120 273.8 -48.75 3 20 41 90 700 160 220.7 -46.87 4 20 52 30 500 120 194.1 -45.76 5 20 52 60 700 160 267.3 -48.54 6 20 52 90 300 80 246.2 -47.82 7 20 63 30 700 160 211.9 -46.52 8 20 63 60 300 80 301.3 -49.58 9 20 63 90 500 120 277.1 -48.85 10 30 41 30 500 160 189.8 -45.65 11 30 41 60 700 80 343.5 -50.72 12 30 41 90 300 120 312.7 -49.90 13 30 52 30 700 80 298.9 -49.51 14 30 52 60 300 120 367.2 -51.29 15 30 52 90 500 160 351.3 -50.91 16 30 63 30 300 120 300.8 -49.56 17 30 63 60 500 160 378.5 -51.56 18 30 63 90 700 80 361.9 -51.17 19 40 41 30 700 120 332.8 -50.44 20 40 41 60 300 160 387.5 -51.76 21 40 41 90 500 80 370.6 -51.37 22 40 52 30 300 160 359.1 -51.10 23 40 52 60 500 80 398.3 -52.00 24 40 52 90 700 120 381.7 -51.63 25 40 63 30 500 80 379.2 -51.58 26 40 63 60 700 120 427.6 -52.62 27 40 63 90 300 160 369.8 -51.36 was increased by about 68.99 % due to the incorporation of 20 % mass fraction of CP particles. The impact-strength values of the WD/CP/PF com- posites after the impact tests are presented in Figure 2c. It can be seen that the impact strength of the WD/PF composite is slightly lower than the value of the neat- resin sample. It is also shown that the impact strength of the WD/PF composite increases with the addition of 10 % mass fraction of CP particles, but it decreases with the incorporation of 20 % mass fraction of CP particles. This may be due to a poor adhesion between the particles and the matrix. It may also be due to the stress concentration of the resin matrix. It is also observed that the incor- poration of 10 % mass fraction and 20 % mass fraction of CP particles shows 15.57 % and 8.19 % higher impact values compared to the WD/PF composite. 3.2 Steady-state erosion: effects of the impingement angle Generally, the erosive-wear behavior of polymer composite materials can be categorized as brittle and ductile. A ductile erosive situation is created in thermo- plastic polymer composites, whereas the a brittle erosive situation may be created in thermosetting polymer composites. For the steady-state-erosion analysis of the WD/CP/PF composites, an erosion test was carried out based on eight different impingement angles (20, 30, 40, 50, 60, 70, 80 and 90)°, keeping all the other process parameters constant (the initial level values). The effects of the impingement angles on the erosion rate of the WD/CP/PF composites are presented in Fig- ure 3. From this figure, it can be observed that the erosion rate is high at the impingement angle of 60° for all the composite specimens, irrespective of the particle content. However, a more brittle erosive behavior was identified for the 40 % mass-fraction (20WD/20CP) composite specimen. However, in the 20 % and 30 % mass-fraction composite specimens, a semi-brittle erosive behavior was identified. This may be due to the addition of WD and CP particles to the PF composites. When the higher amounts of particles are added to the polymer material, it behaves as a typical brittle material. Therefore, the brittleness of the composites with 20 % mass fraction and 30 % mass fraction of particles is lower than the composite with 40 % mass fraction of particles. Due to this, a semi-brittle erosive situation exists during the erosive process. It is also clear from Figure 3 that the erosion rate increased with the increase in the particle content. This may be due to the increased hardness of the PF composite material caused by the addition of WD and CP particles. 3.3 Analysis of the erosion rate The erosion rates for 27 combinations of the erosive experiments conducted on the WD/CP/PF composites are given in Table 3. The erosion analysis was made with popular software, namely, MINITAB 17. From Table 3, it can be concluded that the parameter combina- tion of particle loading-A (level II = 30 % mass fraction), impact velocity-B (level I = 41 m/s), impingement angle-C (level I = 30°), erodent size-D (level II = 500 um) and standoff distance-E (160 mm) gives the mini- mum erosion rate (189. 8 mg/kg). Moreover, another parameter combination (experiment number 4) allows the next level of the minimum erosion rate (194.1 mg/kg). The difference between these two erosion rates is small, as seen from Table 3. Anyway, the first para- meter combination mentioned above is recognized as the better combination of parameters to obtain the minimum erosion rate. The overall mean of the signal-to-noise ratio for the erosion rate is found to be 49.75 dB. A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... 808 Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: Steady-state erosive behavior of WD/CP/PF composites for eight different impingement angles Figure 2: Variation of: a) tensile property, b) flexural property, and c) impact strength based on the particle content The effects of five erosive-process parameters on the erosion rate are graphically presented in Figure 4a. From this figure, it can be clearly concluded that para- meter A (the particle content), parameter B (the impact velocity) and parameter C (the impingement angle) are the most significant parameters. Parameter E (the stan- doff distance) shows a moderately significant influence, while parameter D (the erodent size) has a relatively less significant influence. Figure 4b shows the interaction between the erosive parameters. From this figure, it is observed that a moderate interaction exists between para- meters A and B, and between A and C. The interaction between parameters B and C is below the moderate level. Figures 5a to 5c show 3D surface plots of the ero- sion rate with significant process parameters. The observation is similar to the one made of the interaction plots of the erosion rate. From the erosion test analysis of the WD/CP/PF composites, it can be concluded that the erodent size is most insignificant for the erosion rate. The standoff distance shows relatively less significance when compared to the other three process parameters (particle content, impact velocity and impingement A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 809 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: 3D surface plots of erosion rate vs process parameters: a) A × B, b) A × C, and c) B × C Figure 4: a) Effects of erosive-process parameters and b) effects of interactions of erosive-process parameters on the erosion rate Table 3: The erosion rates and their S/N ratio of WD/CP/PF compo- sites for 27 combinations Experi ment No. A B C D E Erosion rate mg/kg S/N ratio dB 1 20 41 30 300 80 209.5 -46.42 2 20 41 60 500 120 273.8 -48.75 3 20 41 90 700 160 220.7 -46.87 4 20 52 30 500 120 194.1 -45.76 5 20 52 60 700 160 267.3 -48.54 6 20 52 90 300 80 246.2 -47.82 7 20 63 30 700 160 211.9 -46.52 8 20 63 60 300 80 301.3 -49.58 9 20 63 90 500 120 277.1 -48.85 10 30 41 30 500 160 189.8 -45.65 11 30 41 60 700 80 343.5 -50.72 12 30 41 90 300 120 312.7 -49.90 13 30 52 30 700 80 298.9 -49.51 14 30 52 60 300 120 367.2 -51.29 15 30 52 90 500 160 351.3 -50.91 16 30 63 30 300 120 300.8 -49.56 17 30 63 60 500 160 378.5 -51.56 18 30 63 90 700 80 361.9 -51.17 19 40 41 30 700 120 332.8 -50.44 20 40 41 60 300 160 387.5 -51.76 21 40 41 90 500 80 370.6 -51.37 22 40 52 30 300 160 359.1 -51.10 23 40 52 60 500 80 398.3 -52.00 24 40 52 90 700 120 381.7 -51.63 25 40 63 30 500 80 379.2 -51.58 26 40 63 60 700 120 427.6 -52.62 27 40 63 90 300 160 369.8 -51.36 angle). It can be concluded that the combination of the parameters – the particle content at level II, the impact velocity at level I and the impingement angle at level I – gives the minimum erosion rate. Therefore, this com- bination is recognized as the best combination of the erosive-process parameters to get the minimum erosion rate within the selected parameter range. 3.4 Analysis of the variance for the erosion rate Taguchi’s analysis of variance can be used to find the set of significant parameters as well as their interactions in any system. In this study, ANOVA is used to under- stand the contribution of the parameters to the erosion rate and the effects of their interactions on the erosion rate of the WD/CP/PF composite. The ANOVA results for the erosion rate of the WD/CP/PF composite are given in Table 4. The last column (p-value) in this table indicates the highly significant parameters and their main effects depend upon the value of p. The p values for the particle content, impact velocity, impingement angle and standoff distance are 0.000, 0.023, 0.003 and 0.186, indicating their great influence on the erosion rate of the composite. The p values of the interactions of the process parameters show a moderate significance and a significance below the moderate level. Table 5 shows the responses for the S/N ratio (the small-the-better charac- teristic). The order of the erosive-process parameters based on their contributions to obtain the minimum ero- sion rate is the particle content, impingement angle, impact velocity, standoff distance and erodent size. Table 4: Analysis of variance for Means of erosion rate Source DF Seq SS Adj SS Adj MS F P A 2 81404 81403.9 40701.9 120.99 0.000 B 2 7520 7519.9 3759.9 11.18 0.023 C 2 25188 25188.1 12594.0 37.44 0.003 D 2 97 96.7 48.4 0.14 0.870 E 2 1778 1778.4 889.2 2.64 0.186 A*B 4 2698 2698.2 674.6 2.01 0.258 A*C 4 3305 3305.3 826.3 2.46 0.203 B*C 4 791 791.1 197.8 0.59 0.690 Residual Error 4 1346 1345.6 336.4 Total 26 124127 Table 5: Response table for signal to noise ratios (smaller is better) Level A B C D E 1 -47.68 -49.10 -48.51 -49.87 -50.02 2 -50.03 -49.84 -50.76 -49.61 -49.87 3 -51.54 -50.31 -49.99 -49.78 -49.37 Delta 3.86 1.21 2.25 0.26 0.65 Rank 1 3 2 5 4 3.5 Confirmation experiment At the end of this study, we carried out a confirma- tory experiment as the final test to validate the estimated results obtained during the erosive analysis of the WD/CP/PF composite. Therefore, the experimental results were verified with the estimated results using the confirmation test. This test was conducted to predict the erosion rate caused by a new set of erosive-process- parameter levels (A2B1C1E3). To predict the S/N ratio, the following equation can be used: S/Np = (A2-T) + (B1-T) + (C1-T) + (E3-T) + T (2) where T is the overall experimental average of the S/N ratio and S/Np is the value of the predicted S/N ratio. The comparison results for the predicted and experimen- tal S/N ratio of the optimum process parameters are given in Table 6. The difference between the predicted and experimental S/N ratio is 0.51, i.e., an error of 1.1 %. It proves that the model can predict the erosion rate with a reasonable accuracy. Table 6: Comparison results of predicted and experimental signal- to-noise ratio of optimal process parameters Optimal erosive process parameters Predicted Experimental Difference Parameter level A2B1C1E3 A2B1C1E3 Predicted- Experimental Erosion rate mg/kg 191.5 189.8 1.7 Signal-to-noise ratio dB -46.17 -45.66 4 CONCLUSIONS Mechanical properties of a WD/CP/PF composite were analyzed based on the particle content. The results show that the tensile strength of the WD/PF composite increased with the addition of 10 % mass fraction of CP particles, but decreased with the addition of 20 % mass fraction of CP particles. The flexural properties of the WD/PF composite increased with the increase in CP particles. The impact strength of the WD/PF composite also increased with the addition of 10 % mass fraction of CP particles and decreased with further addition of CP particles. The steady-state erosion analysis was carried out for eight different impingement angles on the WD/CP/PF composite. The composite with the lower particle content shows a semi-brittle erosive behavior with a higher erosion wear at the 60° impingement angle. On the other hand, the composite with the higher particle content showed a fully brittle nature of the erosive beha- vior with a higher erosion wear at the 60° impingement angle. From the erosive analysis of the WD/CP/PF com- posite, the process parameters like the particle content, impingement angle and impact velocity are found to be the most significant parameters influencing the erosion rate. The standoff distance shows a moderate influence on the erosion rate, while the erodent size shows a less significant influence on the erosion rate. A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... 810 Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 5 REFERENCES 1 G. K. Mani, J. B. B. Rayappan, D. K. Bisoyi, Synthesis and cha- racterization of kapok fibers and its composites, J. of Appl. Sci., 12 (2012), 1661–1665, doi:10.3923/jas.2012.1661.1665 2 J. A. Khan, M. A. Khan, R. Islam, Mechanical, thermal and degra- dation properties of jute fabric – reinforced polypropylene composites: Effect of potassium permanganate as oxidizing agent, Polym. Compos., 34 (2013), 671–680, doi:10.1002/pc.22470 3 S. Singh, D. Deepak, L. Aggarwal, V. K. Gupta, Tensile and flexural behavior of hemp fiber reinforced virgin recycled HDPE matrix com- posites, Procedia Mater. Sci., 6 (2014), 1696–1702, doi:10.1016/ j.mspro.2014.07.155 4 I. V. Surendra, K. V. Rao, K. V. P. P. Chandu, Fabrication and investi- gation of mechanical properties of sisal, jute & okra natural fiber reinforced hybrid polymer composites, Int. J. Eng. Trends Technol., 19 (2015), 116–120, doi: 10.14445/22315381/IJETT-V19P220 5 D. Kurniawan, B. S. Kim, H. Y. Lee, J. Y. 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Rajkumar, Mechanical pro- perties of luffa fiber and ground nut reinforced epoxy polymer hybrid composites, Procedia Eng., 97 (2014), 2042–2051, doi:10.1016/ j.proeng.2014.12.447 12 ASTM D 638-10, Standard test method for tensile properties of plastics, Annual Book of ASTM Standards, 08.01 (2010), 1–16, ASTM International, West Conshohocken 13 ASTM D 790–10, Standard test methods for flexural properties of un-reinforced and reinforced plastics and electrical insulating materials, Annual Book of ASTM Standards, 08.01 (2010), 1–11, ASTM International, West Conshohocken 14 ISO 180:2000, Plastics – determination of Izod impact strength, third edition, ISO Central Secretariat, Switzerland, 2000 15 U. S. Rao, L. L. R. Rodrigues, Influence of machining parameters on tool wear in drilling of GFRP composites –Taguchi analysis and ANOVA methodology, Proc. of the Inter. Conf. on Advances in Mechanical and Robotics Engineering – MRE 2014, 25–29, doi:10.15224/978-1-63248-002-6-84 A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 811 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS angle). It can be concluded that the combination of the parameters – the particle content at level II, the impact velocity at level I and the impingement angle at level I – gives the minimum erosion rate. Therefore, this com- bination is recognized as the best combination of the erosive-process parameters to get the minimum erosion rate within the selected parameter range. 3.4 Analysis of the variance for the erosion rate Taguchi’s analysis of variance can be used to find the set of significant parameters as well as their interactions in any system. In this study, ANOVA is used to under- stand the contribution of the parameters to the erosion rate and the effects of their interactions on the erosion rate of the WD/CP/PF composite. The ANOVA results for the erosion rate of the WD/CP/PF composite are given in Table 4. The last column (p-value) in this table indicates the highly significant parameters and their main effects depend upon the value of p. The p values for the particle content, impact velocity, impingement angle and standoff distance are 0.000, 0.023, 0.003 and 0.186, indicating their great influence on the erosion rate of the composite. The p values of the interactions of the process parameters show a moderate significance and a significance below the moderate level. Table 5 shows the responses for the S/N ratio (the small-the-better charac- teristic). The order of the erosive-process parameters based on their contributions to obtain the minimum ero- sion rate is the particle content, impingement angle, impact velocity, standoff distance and erodent size. Table 4: Analysis of variance for Means of erosion rate Source DF Seq SS Adj SS Adj MS F P A 2 81404 81403.9 40701.9 120.99 0.000 B 2 7520 7519.9 3759.9 11.18 0.023 C 2 25188 25188.1 12594.0 37.44 0.003 D 2 97 96.7 48.4 0.14 0.870 E 2 1778 1778.4 889.2 2.64 0.186 A*B 4 2698 2698.2 674.6 2.01 0.258 A*C 4 3305 3305.3 826.3 2.46 0.203 B*C 4 791 791.1 197.8 0.59 0.690 Residual Error 4 1346 1345.6 336.4 Total 26 124127 Table 5: Response table for signal to noise ratios (smaller is better) Level A B C D E 1 -47.68 -49.10 -48.51 -49.87 -50.02 2 -50.03 -49.84 -50.76 -49.61 -49.87 3 -51.54 -50.31 -49.99 -49.78 -49.37 Delta 3.86 1.21 2.25 0.26 0.65 Rank 1 3 2 5 4 3.5 Confirmation experiment At the end of this study, we carried out a confirma- tory experiment as the final test to validate the estimated results obtained during the erosive analysis of the WD/CP/PF composite. Therefore, the experimental results were verified with the estimated results using the confirmation test. This test was conducted to predict the erosion rate caused by a new set of erosive-process- parameter levels (A2B1C1E3). To predict the S/N ratio, the following equation can be used: S/Np = (A2-T) + (B1-T) + (C1-T) + (E3-T) + T (2) where T is the overall experimental average of the S/N ratio and S/Np is the value of the predicted S/N ratio. The comparison results for the predicted and experimen- tal S/N ratio of the optimum process parameters are given in Table 6. The difference between the predicted and experimental S/N ratio is 0.51, i.e., an error of 1.1 %. It proves that the model can predict the erosion rate with a reasonable accuracy. Table 6: Comparison results of predicted and experimental signal- to-noise ratio of optimal process parameters Optimal erosive process parameters Predicted Experimental Difference Parameter level A2B1C1E3 A2B1C1E3 Predicted- Experimental Erosion rate mg/kg 191.5 189.8 1.7 Signal-to-noise ratio dB -46.17 -45.66 4 CONCLUSIONS Mechanical properties of a WD/CP/PF composite were analyzed based on the particle content. The results show that the tensile strength of the WD/PF composite increased with the addition of 10 % mass fraction of CP particles, but decreased with the addition of 20 % mass fraction of CP particles. The flexural properties of the WD/PF composite increased with the increase in CP particles. The impact strength of the WD/PF composite also increased with the addition of 10 % mass fraction of CP particles and decreased with further addition of CP particles. The steady-state erosion analysis was carried out for eight different impingement angles on the WD/CP/PF composite. The composite with the lower particle content shows a semi-brittle erosive behavior with a higher erosion wear at the 60° impingement angle. On the other hand, the composite with the higher particle content showed a fully brittle nature of the erosive beha- vior with a higher erosion wear at the 60° impingement angle. From the erosive analysis of the WD/CP/PF com- posite, the process parameters like the particle content, impingement angle and impact velocity are found to be the most significant parameters influencing the erosion rate. The standoff distance shows a moderate influence on the erosion rate, while the erodent size shows a less significant influence on the erosion rate. A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... 810 Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 5 REFERENCES 1 G. K. Mani, J. B. B. Rayappan, D. K. Bisoyi, Synthesis and cha- racterization of kapok fibers and its composites, J. of Appl. Sci., 12 (2012), 1661–1665, doi:10.3923/jas.2012.1661.1665 2 J. A. Khan, M. A. Khan, R. Islam, Mechanical, thermal and degra- dation properties of jute fabric – reinforced polypropylene composites: Effect of potassium permanganate as oxidizing agent, Polym. Compos., 34 (2013), 671–680, doi:10.1002/pc.22470 3 S. Singh, D. Deepak, L. Aggarwal, V. K. Gupta, Tensile and flexural behavior of hemp fiber reinforced virgin recycled HDPE matrix com- posites, Procedia Mater. Sci., 6 (2014), 1696–1702, doi:10.1016/ j.mspro.2014.07.155 4 I. V. Surendra, K. V. Rao, K. V. P. P. Chandu, Fabrication and investi- gation of mechanical properties of sisal, jute & okra natural fiber reinforced hybrid polymer composites, Int. J. Eng. Trends Technol., 19 (2015), 116–120, doi: 10.14445/22315381/IJETT-V19P220 5 D. Kurniawan, B. S. Kim, H. Y. Lee, J. Y. Lim, Effects of repetitive processing, wood content, and coupling agent on the mechanical, thermal, and water absorption properties of wood/polypropylene green composites, J. Adhes. Sci. Technol., 27 (2013), 1301–1312, doi:10.1080/01694243.2012.695948 6 E. Muñoz, J. A. García-Manrique, Water absorption behaviour and its effect on the mechanical properties of flax fibre reinforced bioepoxy composites, Int. J. Polym. Sci., (2015), 1–10, doi:10.1155/ 2015/390275 7 M. G. A. Selvan, A. Athijayamani, Mechanical properties of fragrant screwpine fiber reinforced unsaturated polyester composite: Effect of fiber length, fiber treatment and water absorption, Fibers Polym., 17 (2016), 104–116, doi:10.1007/s12221-016-5593-x 8 S. I. Durowaye, G. I. Lawal, M. A. Akande, V. O. Durowaye, Me- chanical properties of particulate coconut shell and palm fruit poly- ester composites, Int. J. Mater. Eng., 4 (2014), 141–147, doi:10.5923/j.ijme.20140404.04 9 L. Netra, S. Thomas, C. K. Das, R. Adhikari, Analysis of morpho- logical and mechanical behaviours of bamboo flour reinforced polypropylene composites, Nepal J. Sci. Technol., 13 (2012), 95–100, doi:10.3126/njst.v13i1.7447 10 M. A. M. M. Idrus, S. Hamdan, M. R. Rahman, M. S. Islam, Treated tropical wood sawdust-polypropylene polymer composite: mecha- nical and morphological study, J. Biomater. Nano-Biotechnol., 2 (2011), 435–444, doi:10.4236/jbnb.2011.24053 11 R. Panneerdhass, A. Gnanavelbabu, K. Rajkumar, Mechanical pro- perties of luffa fiber and ground nut reinforced epoxy polymer hybrid composites, Procedia Eng., 97 (2014), 2042–2051, doi:10.1016/ j.proeng.2014.12.447 12 ASTM D 638-10, Standard test method for tensile properties of plastics, Annual Book of ASTM Standards, 08.01 (2010), 1–16, ASTM International, West Conshohocken 13 ASTM D 790–10, Standard test methods for flexural properties of un-reinforced and reinforced plastics and electrical insulating materials, Annual Book of ASTM Standards, 08.01 (2010), 1–11, ASTM International, West Conshohocken 14 ISO 180:2000, Plastics – determination of Izod impact strength, third edition, ISO Central Secretariat, Switzerland, 2000 15 U. S. Rao, L. L. R. Rodrigues, Influence of machining parameters on tool wear in drilling of GFRP composites –Taguchi analysis and ANOVA methodology, Proc. of the Inter. Conf. on Advances in Mechanical and Robotics Engineering – MRE 2014, 25–29, doi:10.15224/978-1-63248-002-6-84 A. SUJIN JOSE: EFFECTS OF AN ADDITION OF COIR-PITH PARTICLES ON THE MECHANICAL AND ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 805–811 811 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 813–822 OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL USING ARTIFICIAL NEURAL NETWORK COUPLED WITH GENETIC ALGORITHM OPTIMALNA DOL@INA PODPORE ([ABLONE, VODILA) PRI TERMI^NEM VRTANJU GALVANIZIRANEGA JEKLA Z UPORABO UMETNE NEVRONSKE MRE@E IN GENETSKEGA ALGORITMA Navasingh Rajesh Jesudoss Hynes1, Ramar Kumar1, Jebaraj Angela Jennifa Sujana2 1Mepco Schlenk Engineering College, Department of Mechanical Engineering, Sivakasi, Vindhunagan, 626005 Tamil Nadu, India 2Mepco Schlenk Engineering College, Department of Information Technology, Sivakasi, Vindhunagan, 626005 Tamil Nadu, India findhynes@yahoo.co.in, mepcokumar@gmail.com, ang_jenefa@mepcoeng.ac.in Prejem rokopisa – received: 2016-09-27; sprejem za objavo – accepted for publication: 2017-01-24 doi:10.17222/mit.2016.290 Thermal drilling is a novel sheet-metal-hole-making technique that utilizes the heat produced at the interface of the rotating conical tool and workpiece in order to soften the workpiece and pierce a hole into it. In this work, experiments with thermal drilling of galvanized steel were conducted based on the Taguchi L27 orthogonal array. Significant process parameters such as rotational speed, tool angle and workpiece thickness were varied during the experimentation. In thermal drilling, the thermal-drill tool pushes aside a large amount of workpiece material to form a sleeve, which is often referred to as the bushing length. A predictive model for the bushing length was developed using a feed-forward artificial neural network based on experi- mental data. As the bushing length is closely associated with the tapping process, the influences of the input process parameters play a vital role in fastening galvanized steel with threaded fasteners in diverse engineering applications. The optimization problem was solved by implementing a genetic algorithm under constraint limits to maximize the bushing length. Further, a confirmation test was conducted with the intention to compare the optimum value and its corresponding bushing length predicted by the genetic algorithm. Good agreement was observed between the predicted and the experimental values. Keywords: thermal drilling, artificial neural network, genetic algorithm, galvanized steel, bushing length Termi~no vrtanje je nova, za vrtanje lukenj v plo~evino, uporabljena tehnika, ki izkori{~a toploto, proizvedeno na povr{ini vrte~e se konice orodja na obdelovancu z namenom, da ga zmeh~a in vanj naredi luknjo. V delu so bili izvedeni preizkusi na osnovi metode Taguchi L27 z ortogonalno matriko s termi~nim vrtanjem galvaniziranega jekla. Pomembni parametri postopka, kot so: hitrost vrtenja, kot orodja in debelina obdelovanca, so se med eksperimentiranjem spreminjali. Pri toplotnem vrtanju, vrtanje orodja potisne stran ve~ materiala obdelovanca tako, da se tvori navarek (rokav) okoli luknje, ki se pogosto omenja kot dol`ina {ablone. Napovedni model za dol`ino {ablone, je razvit z uporabo umetne nevronske mre`e, ki temelji na znanstvenih podatkih. Ker je dol`ina {ablone precej povezana s procesom izdelave navoja, vplivi teh vhodnih procesnih parametrov igrajo klju~no vlogo pri pritrditvi galvaniziranega jekla z navojem pritrdilnih elementov v razli~nih in`enirskih aplikacijah. Problem optimizacije je bil re{en z implementacijo genetskega algoritma na podlagi omejitev za pove~anje dol`ine {ablone. Ugotovljeno je bilo dobro ujemanje med napovedano in eksperimentalno vrednostjo. Klju~ne besede: termi~no vrtanje, umetna nevronska mre`a, genetski algoritem, galvanizirano jeklo, dol`ina {ablone 1 INTRODUCTION Thermal drilling is an emerging hole-making process with significant breakthroughs in many drilling situ- ations in both automobile and aerospace applications. This process uses the heat energy produced at the inter- face in order to soften and then pierce the workpiece.1 Moreover, this heat energy enhances the flow ability of the workpiece material, which is extruded onto both the front and back sides of a drilled hole. Finally, the extruded or deformed material forms a bushing shape, which surrounds the drilled hole.2 The length of the bushing formed on the workpiece after a thermal-drilling operation is called the bushing length. Achieving a sufficient bushing length is very important since the bushing length can increase the threading depth and the clamp-load-bearing capability in various engineering applications. Contrary to the conventional drilling process, in thermal drilling, there is no chip and wastage of the ma- terial, since all the deformed material contributes to developing a bushing. As it abolishes chip generation, it can be seen as a clean, eco-friendly and chipless hole- making technique. When joining metal sheets or thin- walled structures by drilling holes using the conven- tional-drilling process, just one or two threads can be made in them; however, reliability is a matter of great concern in such a situation. Alternatively, weld nuts and threaded rivet nuts are used, but due to thermal distortion, they get jammed and twisted during the fastening of the structures.3 On the other hand, during the thermal-drilling process, the bushing formed is about 3 to 4 times thicker than the workpiece. The bushing thus allows a greater contact area and high structural rigidity of fasteners in joining situations.4 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 813 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 519.6:620.1:67.017 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)813(2017) N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 813–822 OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL USING ARTIFICIAL NEURAL NETWORK COUPLED WITH GENETIC ALGORITHM OPTIMALNA DOL@INA PODPORE ([ABLONE, VODILA) PRI TERMI^NEM VRTANJU GALVANIZIRANEGA JEKLA Z UPORABO UMETNE NEVRONSKE MRE@E IN GENETSKEGA ALGORITMA Navasingh Rajesh Jesudoss Hynes1, Ramar Kumar1, Jebaraj Angela Jennifa Sujana2 1Mepco Schlenk Engineering College, Department of Mechanical Engineering, Sivakasi, Vindhunagan, 626005 Tamil Nadu, India 2Mepco Schlenk Engineering College, Department of Information Technology, Sivakasi, Vindhunagan, 626005 Tamil Nadu, India findhynes@yahoo.co.in, mepcokumar@gmail.com, ang_jenefa@mepcoeng.ac.in Prejem rokopisa – received: 2016-09-27; sprejem za objavo – accepted for publication: 2017-01-24 doi:10.17222/mit.2016.290 Thermal drilling is a novel sheet-metal-hole-making technique that utilizes the heat produced at the interface of the rotating conical tool and workpiece in order to soften the workpiece and pierce a hole into it. In this work, experiments with thermal drilling of galvanized steel were conducted based on the Taguchi L27 orthogonal array. Significant process parameters such as rotational speed, tool angle and workpiece thickness were varied during the experimentation. In thermal drilling, the thermal-drill tool pushes aside a large amount of workpiece material to form a sleeve, which is often referred to as the bushing length. A predictive model for the bushing length was developed using a feed-forward artificial neural network based on experi- mental data. As the bushing length is closely associated with the tapping process, the influences of the input process parameters play a vital role in fastening galvanized steel with threaded fasteners in diverse engineering applications. The optimization problem was solved by implementing a genetic algorithm under constraint limits to maximize the bushing length. Further, a confirmation test was conducted with the intention to compare the optimum value and its corresponding bushing length predicted by the genetic algorithm. Good agreement was observed between the predicted and the experimental values. Keywords: thermal drilling, artificial neural network, genetic algorithm, galvanized steel, bushing length Termi~no vrtanje je nova, za vrtanje lukenj v plo~evino, uporabljena tehnika, ki izkori{~a toploto, proizvedeno na povr{ini vrte~e se konice orodja na obdelovancu z namenom, da ga zmeh~a in vanj naredi luknjo. V delu so bili izvedeni preizkusi na osnovi metode Taguchi L27 z ortogonalno matriko s termi~nim vrtanjem galvaniziranega jekla. Pomembni parametri postopka, kot so: hitrost vrtenja, kot orodja in debelina obdelovanca, so se med eksperimentiranjem spreminjali. Pri toplotnem vrtanju, vrtanje orodja potisne stran ve~ materiala obdelovanca tako, da se tvori navarek (rokav) okoli luknje, ki se pogosto omenja kot dol`ina {ablone. Napovedni model za dol`ino {ablone, je razvit z uporabo umetne nevronske mre`e, ki temelji na znanstvenih podatkih. Ker je dol`ina {ablone precej povezana s procesom izdelave navoja, vplivi teh vhodnih procesnih parametrov igrajo klju~no vlogo pri pritrditvi galvaniziranega jekla z navojem pritrdilnih elementov v razli~nih in`enirskih aplikacijah. Problem optimizacije je bil re{en z implementacijo genetskega algoritma na podlagi omejitev za pove~anje dol`ine {ablone. Ugotovljeno je bilo dobro ujemanje med napovedano in eksperimentalno vrednostjo. Klju~ne besede: termi~no vrtanje, umetna nevronska mre`a, genetski algoritem, galvanizirano jeklo, dol`ina {ablone 1 INTRODUCTION Thermal drilling is an emerging hole-making process with significant breakthroughs in many drilling situ- ations in both automobile and aerospace applications. This process uses the heat energy produced at the inter- face in order to soften and then pierce the workpiece.1 Moreover, this heat energy enhances the flow ability of the workpiece material, which is extruded onto both the front and back sides of a drilled hole. Finally, the extruded or deformed material forms a bushing shape, which surrounds the drilled hole.2 The length of the bushing formed on the workpiece after a thermal-drilling operation is called the bushing length. Achieving a sufficient bushing length is very important since the bushing length can increase the threading depth and the clamp-load-bearing capability in various engineering applications. Contrary to the conventional drilling process, in thermal drilling, there is no chip and wastage of the ma- terial, since all the deformed material contributes to developing a bushing. As it abolishes chip generation, it can be seen as a clean, eco-friendly and chipless hole- making technique. When joining metal sheets or thin- walled structures by drilling holes using the conven- tional-drilling process, just one or two threads can be made in them; however, reliability is a matter of great concern in such a situation. Alternatively, weld nuts and threaded rivet nuts are used, but due to thermal distortion, they get jammed and twisted during the fastening of the structures.3 On the other hand, during the thermal-drilling process, the bushing formed is about 3 to 4 times thicker than the workpiece. The bushing thus allows a greater contact area and high structural rigidity of fasteners in joining situations.4 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 813 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 519.6:620.1:67.017 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)813(2017) A. H. Streppel and H. J. J Kals5 suggested that the thermal-drilling process could be applied to different materials like carbon steels, stainless steel, copper, brass and aluminium. They carried out experimentation on the thermal drilling of aluminium. Due to strong pre-strain hardening, the process resulted in bad quality of the bushing. M. Kerkhofs and M. Van Stappen6 compared the performance of (Ti, Al) N coated tungsten carbide thermal drills with uncoated drills. They reported that the coated thermal drills allowed a longer tool life than the uncoated flow drills. S. F. Miller et al.7 conducted experiments with the thermal-drilling process on steel, aluminum and titanium. They studied the properties such as the hardness and microstructure of drilled holes. They concluded that in thermal drilling large deformation and high frictional heat are generated, resulting in the changes in the material properties and microstructure of a workpiece. They reported that the thermal conductivity of the material drilled affects the quality and integrity of the hole. S. F. Miller et al.8 conducted an experimental and numerical analysis of a thermal-drilling process on the AISI 1020 cold-rolled carbon steel under a constant feed rate. Moreover, they developed two classic models for the thermal-drilling process. They predicted the tempe- rature distribution using a finite-element-based thermal model and in another attempt, they predicted the thrust force and torque using the basic principles of physics. S. F. Miller et al.9 thermally drilled into aluminium and magnesium alloys using tungsten carbide and titanium carbide in a cobalt matrix under different rotational speeds and feed rates. They analyzed the thrust force, torque, energy, power and peak power required for drilling. They also evaluated the formation of the bushing shape of cast metals. The bushings produced during the thermal drilling of brittle cast metals demon- strated a radial fracture. In their conclusions, they suggested to pre-heat a cast metal workpiece and create a high rotational speed for obtaining a cylindrically shaped bushing without a significant radial fracture. S. F. Miller et al.10 developed a three-dimensional finite-element model for the thermal-drilling process in order to evaluate the plastic strain and deformation of a workpiece. S. F. Miller et al.11 investigated thermal-drill tool-wear characteristics during the thermal drilling of an AISI 1015 carbon steel workpiece. Furthermore, the thrust force and torque were analyzed. They concluded that the carbide tool is durable and it demonstrates the least tool wear after drilling 11.000 holes. However, progressively severe abrasive grooving on the tool tip was observed. S. M. Lee et al.12 carried out thermal drilling into the IN-713LC super alloy under different rotation speeds and feed rates. They investigated the material properties such as hardness, roundness, and surface roughness of the holes drilled into the IN-713LC super alloy. They reported that the hardness is higher near the wall of a drilled hole and it decreases with the increasing distance from the edge of the hole. In addition to that, higher rotational speeds and feed rates demon- strate better roundness and lower surface roughness. H. M. Chow et al.13 optimized the process parameters of a thermally drilled austenite stainless-steel workpiece using the Taguchi method.14 They considered the input parameters such as friction angle, friction/contact-area ratio, feed rate and drilling speed, and studied their in- fluence on the response parameters like surface rough- ness. Moreover, the hardness and microstructural aspects of drilled holes were studied. It was observed that the surrounding area of a drilled hole acquired a fine grain size and a compact structure with a higher microhardness than that of the area away from the drilled region. S. M. Lee et al.15 employed thermal drilling for the AISI 304 stainless steel using tungsten carbide drills with and without coating. Their results illustrated that at the same rotational speed and for the same number of holes drilled, the coated drills experienced less tool wear than the uncoated drills. Furthermore, they investigated the changes in the relationship between the drilled-surface temperature, tool wear and axial thrust force. W. L. Ku et al.16 optimized the parameters of thermal drilling into austenite stainless steel using the Taguchi method. They studied the effects of the friction angle, friction/contact-area ratio, feed rate and rotational speed on the response-quality characteristics such as the surface roughness and bushing length. They revealed that at optimized drilling conditions, the bushing length of a drilled hole was nearly three times longer than the plate thickness, and a mirror-like quality wall surface of the drilled hole could be obtained. M. Folea et al.17 investi- gated the thermal drilling of a maraging-steel workpiece. The authors reported that the temperature was the most important factor in the thermal-drilling process. They also revealed that a greater quality of the holes drilled into maraging steel could be achieved with higher rota- tional speeds. T. K. Mehmet et al.18 studied the effects of the thermal-drilling parameters such as friction angle, friction/contact-area ratio, feed rate and rotational speed on workpiece temperature, thrust force and torque in thermal drilling of the ST12 material. They revealed that the thrust force and torque reduce with the increasing rotational speed and increase with the friction angle, feed rate and friction/contact-area ratio. D. Biermann and Y. Liu19 demonstrated the feasibi- lity of the thermal drilling of a magnesium wrought alloy and analyzed the thrust forces and torque. They measured the process temperature online and examined the strength of the joint through tapping and thread forming. B. B. Mehmet20 experimentally and numerically investigated the thermal drilling of the AISI 1020 steel. In their study, an analytical model for the torque, axial power and heat-transfer coefficient was developed. Good agreement was observed between experimental and numerical values. N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 814 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Similar work was performed by P. Krasauskas et al.21 on the thermal drilling into the AISI 304 steel. P. D. Pantawane and B. B. Ahuja22 carried out regression modelling of the thermal drilling of the AISI 1015 steel using the Taguchi method. They studied the effects of the tool-diameter ratio and rotational speed on the material thickness and the effect of the feed on the response parameters including the thrust force, torque and surface roughness of drilled holes. G. Somasundaram et al.23 carried thermal drilling on the Al-SiC composite material and studied the roundness errors on drilled holes. They considered the input process parameters such as the composition of workpiece, workpiece thick- ness, rotational speed and feed rate. They developed an empirical relation between the process parameters and the roundness error using the response surface methodo- logy. Artificial neural network (ANN) is a biologically in- spired computer program designed to simulate the way, in which the human brain processes information.24 ANN is widely applied in modeling many machining ope- rations like turning, milling and drilling. S. R. Karnik and V. N. Gaitone25 used a multilayer feed-forward ANN to predict the influence of the process parameters such as the cutting speed, feed, drill diameter, point angle and lip-clearance angle on the burr height and burr thickness in drilling the AISI 316L stainless steel. R. S. Mamilla et al.26 applied a multilayer ANN for modeling an abrasive flow-finishing process on the AISI 1040 and AISI 4340 steel. S. Assarzadeh and M. Ghoreishi27 successfully used a multilayer ANN for modeling the metal-removal rate and surface roughness in an electrical discharge machining process involving the BD3 steel. O. Babur et al.28 developed an ANN model according to experimental measurement data for the end milling of Inconel 718. They coupled a genetic algorithm (GA) with an ANN to forecast the surface roughness. S. Sarkar et al.29 pro- duced a multilayer feed-forward-ANN model to predict the process parameters of the machining of  titanium aluminide with a wire-electrical-discharge machine. A. K. Singh et al.30 used a multilayer feed-forward ANN to predict the flank wear of high-speed steel drill bits for drilling holes into a copper workpiece. The literature review on thermal drilling reveals that experimental investigations and numerical simulations of the effects of mechanical and physical properties on the thermal-drilling process was mainly carried out on copper, brass, aluminium, magnesium and stainless-steel alloys, maraging steel, ST12 steel and IN-713LC super alloy. Galvanized steel is being widely used for car body structures and it is a good candidate for the implemen- tation of the thermal-drilling process. The application of galvanized steel is widely used in the areas of roofing material, doors, ship’s ducts and panels, electrical- appliance automobile parts, etc.31 In the present work, thermal drilling was carried out on galvanized steel. Based on experimental measurement data, an ANN model was developed to predict the for- mation of the bushing length in thermal drilling of galvanized steel. This model considers rotational speed, tool angle and workpiece thickness as significant ther- mal-drilling process parameters. Then the optimum values of these drilling parameters were computed, with the aim of achieving the maximum bushing length, by solving the optimization problem implementing a GA. Thus, the present work allows us to achieve the maxi- mum bushing length, which is highly desirable for the subsequent tapping and joining in automotive applica- tions. A schematic diagram of the combined ANN-GA optimization24 is shown in Figure 1. It indicates the stages of the GA process and its relation with the ANN process. 2 THERMAL DRILLING 2.1 Physical description of the process Figure 2 shows a schematic representation of the thermal-drilling process. Based on the thermal-drill geometry, there are four steps involved in thermal drill- ing. Initially, the center point zone of a drill approaches and pierces the workpiece. Secondly, the produced heat softens the workpiece as a result of the friction between the thermal drill and the workpiece. Thirdly, the softened material is pushed downward and the drill moves for- ward to form the bushing using the cylindrical zone of the tool. Fourthly, the extruded burr on the workpiece surface is pressed to form a boss with the shoulder zone N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 815 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 1: Flow chart of the integrated ANN-GA optimization process A. H. Streppel and H. J. J Kals5 suggested that the thermal-drilling process could be applied to different materials like carbon steels, stainless steel, copper, brass and aluminium. They carried out experimentation on the thermal drilling of aluminium. Due to strong pre-strain hardening, the process resulted in bad quality of the bushing. M. Kerkhofs and M. Van Stappen6 compared the performance of (Ti, Al) N coated tungsten carbide thermal drills with uncoated drills. They reported that the coated thermal drills allowed a longer tool life than the uncoated flow drills. S. F. Miller et al.7 conducted experiments with the thermal-drilling process on steel, aluminum and titanium. They studied the properties such as the hardness and microstructure of drilled holes. They concluded that in thermal drilling large deformation and high frictional heat are generated, resulting in the changes in the material properties and microstructure of a workpiece. They reported that the thermal conductivity of the material drilled affects the quality and integrity of the hole. S. F. Miller et al.8 conducted an experimental and numerical analysis of a thermal-drilling process on the AISI 1020 cold-rolled carbon steel under a constant feed rate. Moreover, they developed two classic models for the thermal-drilling process. They predicted the tempe- rature distribution using a finite-element-based thermal model and in another attempt, they predicted the thrust force and torque using the basic principles of physics. S. F. Miller et al.9 thermally drilled into aluminium and magnesium alloys using tungsten carbide and titanium carbide in a cobalt matrix under different rotational speeds and feed rates. They analyzed the thrust force, torque, energy, power and peak power required for drilling. They also evaluated the formation of the bushing shape of cast metals. The bushings produced during the thermal drilling of brittle cast metals demon- strated a radial fracture. In their conclusions, they suggested to pre-heat a cast metal workpiece and create a high rotational speed for obtaining a cylindrically shaped bushing without a significant radial fracture. S. F. Miller et al.10 developed a three-dimensional finite-element model for the thermal-drilling process in order to evaluate the plastic strain and deformation of a workpiece. S. F. Miller et al.11 investigated thermal-drill tool-wear characteristics during the thermal drilling of an AISI 1015 carbon steel workpiece. Furthermore, the thrust force and torque were analyzed. They concluded that the carbide tool is durable and it demonstrates the least tool wear after drilling 11.000 holes. However, progressively severe abrasive grooving on the tool tip was observed. S. M. Lee et al.12 carried out thermal drilling into the IN-713LC super alloy under different rotation speeds and feed rates. They investigated the material properties such as hardness, roundness, and surface roughness of the holes drilled into the IN-713LC super alloy. They reported that the hardness is higher near the wall of a drilled hole and it decreases with the increasing distance from the edge of the hole. In addition to that, higher rotational speeds and feed rates demon- strate better roundness and lower surface roughness. H. M. Chow et al.13 optimized the process parameters of a thermally drilled austenite stainless-steel workpiece using the Taguchi method.14 They considered the input parameters such as friction angle, friction/contact-area ratio, feed rate and drilling speed, and studied their in- fluence on the response parameters like surface rough- ness. Moreover, the hardness and microstructural aspects of drilled holes were studied. It was observed that the surrounding area of a drilled hole acquired a fine grain size and a compact structure with a higher microhardness than that of the area away from the drilled region. S. M. Lee et al.15 employed thermal drilling for the AISI 304 stainless steel using tungsten carbide drills with and without coating. Their results illustrated that at the same rotational speed and for the same number of holes drilled, the coated drills experienced less tool wear than the uncoated drills. Furthermore, they investigated the changes in the relationship between the drilled-surface temperature, tool wear and axial thrust force. W. L. Ku et al.16 optimized the parameters of thermal drilling into austenite stainless steel using the Taguchi method. They studied the effects of the friction angle, friction/contact-area ratio, feed rate and rotational speed on the response-quality characteristics such as the surface roughness and bushing length. They revealed that at optimized drilling conditions, the bushing length of a drilled hole was nearly three times longer than the plate thickness, and a mirror-like quality wall surface of the drilled hole could be obtained. M. Folea et al.17 investi- gated the thermal drilling of a maraging-steel workpiece. The authors reported that the temperature was the most important factor in the thermal-drilling process. They also revealed that a greater quality of the holes drilled into maraging steel could be achieved with higher rota- tional speeds. T. K. Mehmet et al.18 studied the effects of the thermal-drilling parameters such as friction angle, friction/contact-area ratio, feed rate and rotational speed on workpiece temperature, thrust force and torque in thermal drilling of the ST12 material. They revealed that the thrust force and torque reduce with the increasing rotational speed and increase with the friction angle, feed rate and friction/contact-area ratio. D. Biermann and Y. Liu19 demonstrated the feasibi- lity of the thermal drilling of a magnesium wrought alloy and analyzed the thrust forces and torque. They measured the process temperature online and examined the strength of the joint through tapping and thread forming. B. B. Mehmet20 experimentally and numerically investigated the thermal drilling of the AISI 1020 steel. In their study, an analytical model for the torque, axial power and heat-transfer coefficient was developed. Good agreement was observed between experimental and numerical values. N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 814 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Similar work was performed by P. Krasauskas et al.21 on the thermal drilling into the AISI 304 steel. P. D. Pantawane and B. B. Ahuja22 carried out regression modelling of the thermal drilling of the AISI 1015 steel using the Taguchi method. They studied the effects of the tool-diameter ratio and rotational speed on the material thickness and the effect of the feed on the response parameters including the thrust force, torque and surface roughness of drilled holes. G. Somasundaram et al.23 carried thermal drilling on the Al-SiC composite material and studied the roundness errors on drilled holes. They considered the input process parameters such as the composition of workpiece, workpiece thick- ness, rotational speed and feed rate. They developed an empirical relation between the process parameters and the roundness error using the response surface methodo- logy. Artificial neural network (ANN) is a biologically in- spired computer program designed to simulate the way, in which the human brain processes information.24 ANN is widely applied in modeling many machining ope- rations like turning, milling and drilling. S. R. Karnik and V. N. Gaitone25 used a multilayer feed-forward ANN to predict the influence of the process parameters such as the cutting speed, feed, drill diameter, point angle and lip-clearance angle on the burr height and burr thickness in drilling the AISI 316L stainless steel. R. S. Mamilla et al.26 applied a multilayer ANN for modeling an abrasive flow-finishing process on the AISI 1040 and AISI 4340 steel. S. Assarzadeh and M. Ghoreishi27 successfully used a multilayer ANN for modeling the metal-removal rate and surface roughness in an electrical discharge machining process involving the BD3 steel. O. Babur et al.28 developed an ANN model according to experimental measurement data for the end milling of Inconel 718. They coupled a genetic algorithm (GA) with an ANN to forecast the surface roughness. S. Sarkar et al.29 pro- duced a multilayer feed-forward-ANN model to predict the process parameters of the machining of  titanium aluminide with a wire-electrical-discharge machine. A. K. Singh et al.30 used a multilayer feed-forward ANN to predict the flank wear of high-speed steel drill bits for drilling holes into a copper workpiece. The literature review on thermal drilling reveals that experimental investigations and numerical simulations of the effects of mechanical and physical properties on the thermal-drilling process was mainly carried out on copper, brass, aluminium, magnesium and stainless-steel alloys, maraging steel, ST12 steel and IN-713LC super alloy. Galvanized steel is being widely used for car body structures and it is a good candidate for the implemen- tation of the thermal-drilling process. The application of galvanized steel is widely used in the areas of roofing material, doors, ship’s ducts and panels, electrical- appliance automobile parts, etc.31 In the present work, thermal drilling was carried out on galvanized steel. Based on experimental measurement data, an ANN model was developed to predict the for- mation of the bushing length in thermal drilling of galvanized steel. This model considers rotational speed, tool angle and workpiece thickness as significant ther- mal-drilling process parameters. Then the optimum values of these drilling parameters were computed, with the aim of achieving the maximum bushing length, by solving the optimization problem implementing a GA. Thus, the present work allows us to achieve the maxi- mum bushing length, which is highly desirable for the subsequent tapping and joining in automotive applica- tions. A schematic diagram of the combined ANN-GA optimization24 is shown in Figure 1. It indicates the stages of the GA process and its relation with the ANN process. 2 THERMAL DRILLING 2.1 Physical description of the process Figure 2 shows a schematic representation of the thermal-drilling process. Based on the thermal-drill geometry, there are four steps involved in thermal drill- ing. Initially, the center point zone of a drill approaches and pierces the workpiece. Secondly, the produced heat softens the workpiece as a result of the friction between the thermal drill and the workpiece. Thirdly, the softened material is pushed downward and the drill moves for- ward to form the bushing using the cylindrical zone of the tool. Fourthly, the extruded burr on the workpiece surface is pressed to form a boss with the shoulder zone N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 815 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 1: Flow chart of the integrated ANN-GA optimization process of the thermal drill. Finally, the thermal drill retracts leaving a hole with a bushing length.37 2.2 Formation of the bushing length The initial volume of the material (Vi) available to produce a hole by thermal drilling is given in Equation (1).33 As shown in Figure 2, D1 and Pt represent the inner-hole diameter in mm and the thickness of the workpiece in mm, respectively. V D Pi = π 4 1 2 t (1) During thermal drilling, the volume of the material to be displaced (Vf) in order to produce a bushing is given in Equation (2).33 As shown in Figure 2, D2 and L represent the outer-bushing diameter in mm and the bushing length in mm, respectively. [ ]V D D Lf = − π 4 2 2 1 2 (2) It is known that the initial approximation of the volume of the material for producing a hole is the same as the volume of the material to be displaced during the formation of the bushing. Therefore, Equations (1) and (2) are assumed to be equal. Equating Equations (1) and (2), Equation (3) is obtained: [ ]π π 4 41 2 2 2 1 2D P D D Lt = − (3) While rearranging the terms in Equation (3), the bushing length (L) can be determined as shown in Equation (4): [ ]L D D D P= − 1 2 2 2 1 2 t (4) 3 EXPERIMENTAL PART In the present work, the experiments were carried out at in a vertical drilling machine. A hot-dip-galvanized steel workpiece with thickness values of 1, 1.5, 2 mm and with dimensions of 50 mm × 150 mm was used. 6-mm-diameter thermal drills with different tool angles were used in the experimentation. They were machined out of high-speed steel rods, having the required dimensions as shown in Figure 3. The feed rate of the thermal-drilling process is kept constant at 240 mm/min. The chemical composition of the galvanized steel is given in Table 1. Axial forces were measured using a digital drilling-tool dynamometer. Temperature measure- ments were done on-line using a non-contact-type infra-red thermometer and the temperature profile of the process was obtained using the DATA TEMP MX software with a time increment of 0.016 s. The micro- structure of the holes drilled into galvanized steel was measured using a scanning electron microscope (Carl Zeiss EVO18, Germany). Three different levels of the input process parameters were used, as shown in Table 2. Based on the proper selection of the input process parameters, the desirable output parameters can be obtained during the experimentation. The selected input process parameters such as the rotational speed, tool N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 816 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: Obtained thermal drills with different tool angles: a) 35 de- gree tool angle b) 37.5 degree tool angle, c) 45 degree tool angle Table 1: Chemical composition of galvanized steel (w/%) C Si S P Mn Ni Cr 0.003 0.006 0.005 0.018 0.173 0.011 0.031 Mo V Cu Al Nb Zn Ti Fe 0.001 0.002 0.017 0.035 0.001 0.004 0.05 Balance Table 2: Input process parameters and their levels Factors Levels 1 2 3 Rotational speed (min–1) 1944 2772 3600 Tool angle (°) 30 37.5 45 Workpiece thickness (mm) 1 1.5 2Figure 2: Schematic representation of the thermal-drilling process angle and workpiece thickness, and the output parameter, the bushing length of the thermally drilled galvanized steel sheet, were analyzed in this work. Based on the Taguchi’s L27 orthogonal array, experiments were con- ducted at three levels of the drilling-process parameters and the corresponding bushing length was measured using a digital vernier caliper (Model No CD-12C, Mitu- toya Corporation, Japan). All the experimental values of the bushing length are displayed in Table 3. Samples of the thermally drilled holes are shown in Figure 4. During the thermal drilling, the burr is extruded onto the top surface, and it is leveled when it comes in contact with the shoulder of the thermal-drill tool as shown in Figure 5a. The formation of the bushing at the rear side of the galvanized steel plate is shown in Figure 5b. It can be used to act as a cylindrical sleeve bearing and it can be taped to fabricate an internal screw. Additionally, it widely reduces the complexity of joining components in the aerospace and automotive industries. Table 3: Experimental results of the thermal-drilling process Expt No. Rotational speed (min–1) Tool angle (°) Workpiece thickness (mm) Bushing length (mm) 1 1944 30 1 2.54 2 1944 30 1.5 3.42 3 1944 30 2 4.38 4 1944 37.5 1 3.24 5 1944 37.5 1.5 3.21 6 1944 37.5 2 4.41 7 1944 45 1 2.97 8 1944 45 1.5 4.87 9 1944 45 2 5.26 10 2772 30 1 2.67 11 2772 30 1.5 3.84 12 2772 30 2 4.02 13 2772 37.5 1 2.91 14 2772 37.5 1.5 3.83 15 2772 37.5 2 4.42 16 2772 45 1 2.94 17 2772 45 1.5 4.31 18 2772 45 2 5.57 19 3600 30 1 3.55 20 3600 30 1.5 4.64 21 3600 30 2 5.84 22 3600 37.5 1 2.78 23 3600 37.5 1.5 3.85 24 3600 37.5 2 5.92 25 3600 45 1 2.42 26 3600 45 1.5 4.42 27 3600 45 2 5.61 4 ARTIFICIAL-NEURAL-NETWORK MODELING Recently, the use of the artificial intelligence perfor- mance has been remarkable in the entire engineering field.24 To understand and manage any industrial course of action, modeling and optimization of the process are essential. An accurate control is a requirement necessary to accomplish better quality and productivity. ANN plays a significant role in forecasting the solutions for non- linear problems in all engineering fields. Using statistical techniques, numerous researchers endeavored to develop a model based on experimental data. In general, a neural network represents a network of many processors ope- rating in parallel. Each processor contains a small size of the local memory. Then, these units are coupled through communication channels, which typically carry numeric data. An excellent instance of a biological neural net- work is the brain of a human. It comprises the most demanding and controlling arrangement, in which learning along with training controls the behavior of a human to take action to solve any problem met in day-by-day life. ANN can be successfully employed to predict the output parameters for any given input parameters, based on the training set for a given complex problem.28 In the N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 817 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: Appearance of a hole made by thermal drilling: a) top view and b) rear view Figure 4: Holes made by thermal drilling in galvanized steel plates of the thermal drill. Finally, the thermal drill retracts leaving a hole with a bushing length.37 2.2 Formation of the bushing length The initial volume of the material (Vi) available to produce a hole by thermal drilling is given in Equation (1).33 As shown in Figure 2, D1 and Pt represent the inner-hole diameter in mm and the thickness of the workpiece in mm, respectively. V D Pi = π 4 1 2 t (1) During thermal drilling, the volume of the material to be displaced (Vf) in order to produce a bushing is given in Equation (2).33 As shown in Figure 2, D2 and L represent the outer-bushing diameter in mm and the bushing length in mm, respectively. [ ]V D D Lf = − π 4 2 2 1 2 (2) It is known that the initial approximation of the volume of the material for producing a hole is the same as the volume of the material to be displaced during the formation of the bushing. Therefore, Equations (1) and (2) are assumed to be equal. Equating Equations (1) and (2), Equation (3) is obtained: [ ]π π 4 41 2 2 2 1 2D P D D Lt = − (3) While rearranging the terms in Equation (3), the bushing length (L) can be determined as shown in Equation (4): [ ]L D D D P= − 1 2 2 2 1 2 t (4) 3 EXPERIMENTAL PART In the present work, the experiments were carried out at in a vertical drilling machine. A hot-dip-galvanized steel workpiece with thickness values of 1, 1.5, 2 mm and with dimensions of 50 mm × 150 mm was used. 6-mm-diameter thermal drills with different tool angles were used in the experimentation. They were machined out of high-speed steel rods, having the required dimensions as shown in Figure 3. The feed rate of the thermal-drilling process is kept constant at 240 mm/min. The chemical composition of the galvanized steel is given in Table 1. Axial forces were measured using a digital drilling-tool dynamometer. Temperature measure- ments were done on-line using a non-contact-type infra-red thermometer and the temperature profile of the process was obtained using the DATA TEMP MX software with a time increment of 0.016 s. The micro- structure of the holes drilled into galvanized steel was measured using a scanning electron microscope (Carl Zeiss EVO18, Germany). Three different levels of the input process parameters were used, as shown in Table 2. Based on the proper selection of the input process parameters, the desirable output parameters can be obtained during the experimentation. The selected input process parameters such as the rotational speed, tool N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 816 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: Obtained thermal drills with different tool angles: a) 35 de- gree tool angle b) 37.5 degree tool angle, c) 45 degree tool angle Table 1: Chemical composition of galvanized steel (w/%) C Si S P Mn Ni Cr 0.003 0.006 0.005 0.018 0.173 0.011 0.031 Mo V Cu Al Nb Zn Ti Fe 0.001 0.002 0.017 0.035 0.001 0.004 0.05 Balance Table 2: Input process parameters and their levels Factors Levels 1 2 3 Rotational speed (min–1) 1944 2772 3600 Tool angle (°) 30 37.5 45 Workpiece thickness (mm) 1 1.5 2Figure 2: Schematic representation of the thermal-drilling process angle and workpiece thickness, and the output parameter, the bushing length of the thermally drilled galvanized steel sheet, were analyzed in this work. Based on the Taguchi’s L27 orthogonal array, experiments were con- ducted at three levels of the drilling-process parameters and the corresponding bushing length was measured using a digital vernier caliper (Model No CD-12C, Mitu- toya Corporation, Japan). All the experimental values of the bushing length are displayed in Table 3. Samples of the thermally drilled holes are shown in Figure 4. During the thermal drilling, the burr is extruded onto the top surface, and it is leveled when it comes in contact with the shoulder of the thermal-drill tool as shown in Figure 5a. The formation of the bushing at the rear side of the galvanized steel plate is shown in Figure 5b. It can be used to act as a cylindrical sleeve bearing and it can be taped to fabricate an internal screw. Additionally, it widely reduces the complexity of joining components in the aerospace and automotive industries. Table 3: Experimental results of the thermal-drilling process Expt No. Rotational speed (min–1) Tool angle (°) Workpiece thickness (mm) Bushing length (mm) 1 1944 30 1 2.54 2 1944 30 1.5 3.42 3 1944 30 2 4.38 4 1944 37.5 1 3.24 5 1944 37.5 1.5 3.21 6 1944 37.5 2 4.41 7 1944 45 1 2.97 8 1944 45 1.5 4.87 9 1944 45 2 5.26 10 2772 30 1 2.67 11 2772 30 1.5 3.84 12 2772 30 2 4.02 13 2772 37.5 1 2.91 14 2772 37.5 1.5 3.83 15 2772 37.5 2 4.42 16 2772 45 1 2.94 17 2772 45 1.5 4.31 18 2772 45 2 5.57 19 3600 30 1 3.55 20 3600 30 1.5 4.64 21 3600 30 2 5.84 22 3600 37.5 1 2.78 23 3600 37.5 1.5 3.85 24 3600 37.5 2 5.92 25 3600 45 1 2.42 26 3600 45 1.5 4.42 27 3600 45 2 5.61 4 ARTIFICIAL-NEURAL-NETWORK MODELING Recently, the use of the artificial intelligence perfor- mance has been remarkable in the entire engineering field.24 To understand and manage any industrial course of action, modeling and optimization of the process are essential. An accurate control is a requirement necessary to accomplish better quality and productivity. ANN plays a significant role in forecasting the solutions for non- linear problems in all engineering fields. Using statistical techniques, numerous researchers endeavored to develop a model based on experimental data. In general, a neural network represents a network of many processors ope- rating in parallel. Each processor contains a small size of the local memory. Then, these units are coupled through communication channels, which typically carry numeric data. An excellent instance of a biological neural net- work is the brain of a human. It comprises the most demanding and controlling arrangement, in which learning along with training controls the behavior of a human to take action to solve any problem met in day-by-day life. ANN can be successfully employed to predict the output parameters for any given input parameters, based on the training set for a given complex problem.28 In the N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 817 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: Appearance of a hole made by thermal drilling: a) top view and b) rear view Figure 4: Holes made by thermal drilling in galvanized steel plates present work, an ANN is developed using MATLAB version 7.10, neural network toolbox, with the intention of predicting the bushing length as a function of three input process parameters including the rotational speed, tool angle and workpiece thickness. Figure 6 shows the neural-network architecture of the bushing length ob- tained with thermal drilling. The input data is provided through input layer neurons, and it is fed forward through the hidden layer. Then the response is obtained at the output layer neurons. The neurons are linked by weights that take part in the learning process. The weights help in attaining the optimum solution and escape from the local minima. To establish the optimum structural design, we apply the trial-and-error method with the appropriate activation function and the most excellent training algorithm. A number of models were created and tested. A hyperbolic tangent sigmoid transfer function is considered as the hidden-layer activation function, and it is given in Equation (5): f x x e x ( ) ( )= = − −−tansig 2 1 12 (5) The most important selection criteria used for the suitable model are the mean squared error and the regression-coefficient value. The mean sum of the squared error (MSE) is the average squared difference between the predicted value and the actual experimental value as shown in Equation (6). A lower value of the MSE indicates a better model. The regression coefficient (R) was measured by correlating the predicted and the actual values as given in Equation (7). An R value of 1 indicates a close relationship between the predicted and actual experimental value, i.e., a very small error. MSE n A Pi i i = −∑1 2 (6) R A P P i i i n i i n= − − ⎛ ⎝ ⎜ ⎜ ⎜ ⎞ ⎠ ⎟ ⎟ ⎟ = = ∑ ∑ 1 2 1 2 1 (7) where Ai is the actual experimental value, Pi is the predicted value and n is the number of patterns. With the intention of measuring the accuracy of the prediction model, the percentage of error is calculated with Equation (8): z E P E = − ×100 (8) where z = % of prediction error, E = experimental value, P = predicted value 5 RESULTS AND DISCUSSION The results of the ANN coupled with the GA34 are utilized to predict and optimize the bushing length based on the input process parameters such as the rotational speed, tool angle and workpiece thickness in the ther- mal-drilling process, and they are discussed below. 5.1 Prediction of the bushing length by the ANN Table 4 displays the parameters used with this modeling technique. The developed ANN model of the bushing length was trained by means of the selected parameters. During the training process, a 19 set input for training trials were presented each time and the resultant output was acquired. The number of the hidden-layer neurons varied from 1 to 20. It was deter- mined based on trial and error, through step-by-step increasing the number of neurons and examining their result against the forecast value. The final neuron number of the hidden layer obtained was 10. The ulti- N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 818 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 7: Correlation between the predicted values of the ANN model and the experimental data for the prediction of the bushing length using: a) training data, b) validation data, c) test data and d) entire data Figure 6: ANN modeling of the bushing length in thermal drilling mate weights between the input layer and hidden layer (w1, w2, w3), the bias (b) and weights between the hidden layer and the output layer (w4) are shown in Table 5. Finally, the lowest MSE value achieved was 0.00415 when the configuration of the ANN was 3-10-1. There are three input neurons, each indicating one input variable form the input layer, 10 neurons in the hidden layer and the output layer with one neuron correspond to the bushing length. Table 4: Parameters used in the ANN modeling Input variables Rotational speed, tool angle,workpiece thickness Output variable Bushing length Network type Feed-forward backpropagation Algorithm Levenberg-Marquardt back-propagation algorithm (trainlm) Activation function Hyperbolic tangent sigmoidtransfer function Data division 70 % (training); 15 %(validation); 15 % (testing) Number of neurons in the hidden layer 10 Table 5: Final weights in between the layers No. of neurons Weights Input to hidden layer Hidden tooutput layer w1 w2 w3 b w4 1 -0.682 -1.572 1.929 3.529 0.680 2 -0.865 1.620 -2.115 2.483 -0.576 3 3.045 0.934 -0.191 -1.179 0.111 4 1.722 -2.097 1.179 -1.184 -0.253 5 -1.826 -1.470 1.863 0.273 -0.183 6 -0.027 -0.448 3.025 -0.131 1.113 7 -0.619 2.955 0.332 -0.694 0.268 8 0.855 -1.318 -2.373 1.895 0.036 9 1.645 0.306 -2.767 2.418 0.939 10 1.822 -1.597 -1.613 3.188 -0.638 The output performance of the established ANN is examined on the basis of the regression-correlation coefficient (R value) between the predicted values and the experimental values. The predicted responses of the ANN model are in excellent conformity with the expe- rimental values, i.e., the correlation regression coeffi- cients of 100 % for the training set, 98.41 % for the validation set and 94.80 % for the testing set are achieved as shown in Figure 7. It was concluded that the feed-forward back-propagation algorithm of the configu- ration of 3-10-1 gives the most excellent results for the prediction of the bushing length as shown in Table 6. The graphical representation of the experimental and predicted values of the bushing length by the ANN is shown in Figure 8. It shows that the ANN prediction values are very close to the experimental values of the bushing length. The average prediction error between the experimental and predicted values of the bushing length is 1.843 %. Therefore, it is demonstrated that the developed ANN model is suitable for forecasting the bushing length of drilled holes in thermal drilling of galvanized steel, having the highest accuracy of 98.157 %. N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 819 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 8: Correlation between experimental and predicted values Table 6: Comparison of experimental and predicted values of the bushing length Expt No. Bushing length (mm) Prediction error (%)Experimental value Predicted ANN value 1 2.54 2.485 2.165 2 3.42 3.427 0.204 3 4.38 4.252 2.922 4 3.24 2.618 19.197 5 3.21 3.336 3.925 6 4.41 4.425 0.340 7 2.97 2.843 4.276 8 4.87 4.677 3.963 9 5.26 5.517 4.886 10 2.67 2.385 10.674 11 3.84 3.714 3.281 12 4.02 4.275 6.343 13 2.91 2.905 0.172 14 3.83 3.814 0.418 15 4.42 4.615 4.412 16 2.94 2.964 0.816 17 4.31 4.284 0.603 18 5.57 5.599 0.521 19 3.55 3.738 5.296 20 4.64 4.585 1.185 21 5.84 5.893 0.907 22 2.78 2.760 0.719 23 3.85 3.875 0.649 24 5.92 5.836 1.419 25 2.42 2.755 13.843 26 4.42 4.450 0.679 27 5.61 6.007 7.077 present work, an ANN is developed using MATLAB version 7.10, neural network toolbox, with the intention of predicting the bushing length as a function of three input process parameters including the rotational speed, tool angle and workpiece thickness. Figure 6 shows the neural-network architecture of the bushing length ob- tained with thermal drilling. The input data is provided through input layer neurons, and it is fed forward through the hidden layer. Then the response is obtained at the output layer neurons. The neurons are linked by weights that take part in the learning process. The weights help in attaining the optimum solution and escape from the local minima. To establish the optimum structural design, we apply the trial-and-error method with the appropriate activation function and the most excellent training algorithm. A number of models were created and tested. A hyperbolic tangent sigmoid transfer function is considered as the hidden-layer activation function, and it is given in Equation (5): f x x e x ( ) ( )= = − −−tansig 2 1 12 (5) The most important selection criteria used for the suitable model are the mean squared error and the regression-coefficient value. The mean sum of the squared error (MSE) is the average squared difference between the predicted value and the actual experimental value as shown in Equation (6). A lower value of the MSE indicates a better model. The regression coefficient (R) was measured by correlating the predicted and the actual values as given in Equation (7). An R value of 1 indicates a close relationship between the predicted and actual experimental value, i.e., a very small error. MSE n A Pi i i = −∑1 2 (6) R A P P i i i n i i n= − − ⎛ ⎝ ⎜ ⎜ ⎜ ⎞ ⎠ ⎟ ⎟ ⎟ = = ∑ ∑ 1 2 1 2 1 (7) where Ai is the actual experimental value, Pi is the predicted value and n is the number of patterns. With the intention of measuring the accuracy of the prediction model, the percentage of error is calculated with Equation (8): z E P E = − ×100 (8) where z = % of prediction error, E = experimental value, P = predicted value 5 RESULTS AND DISCUSSION The results of the ANN coupled with the GA34 are utilized to predict and optimize the bushing length based on the input process parameters such as the rotational speed, tool angle and workpiece thickness in the ther- mal-drilling process, and they are discussed below. 5.1 Prediction of the bushing length by the ANN Table 4 displays the parameters used with this modeling technique. The developed ANN model of the bushing length was trained by means of the selected parameters. During the training process, a 19 set input for training trials were presented each time and the resultant output was acquired. The number of the hidden-layer neurons varied from 1 to 20. It was deter- mined based on trial and error, through step-by-step increasing the number of neurons and examining their result against the forecast value. The final neuron number of the hidden layer obtained was 10. The ulti- N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 818 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 7: Correlation between the predicted values of the ANN model and the experimental data for the prediction of the bushing length using: a) training data, b) validation data, c) test data and d) entire data Figure 6: ANN modeling of the bushing length in thermal drilling mate weights between the input layer and hidden layer (w1, w2, w3), the bias (b) and weights between the hidden layer and the output layer (w4) are shown in Table 5. Finally, the lowest MSE value achieved was 0.00415 when the configuration of the ANN was 3-10-1. There are three input neurons, each indicating one input variable form the input layer, 10 neurons in the hidden layer and the output layer with one neuron correspond to the bushing length. Table 4: Parameters used in the ANN modeling Input variables Rotational speed, tool angle,workpiece thickness Output variable Bushing length Network type Feed-forward backpropagation Algorithm Levenberg-Marquardt back-propagation algorithm (trainlm) Activation function Hyperbolic tangent sigmoidtransfer function Data division 70 % (training); 15 %(validation); 15 % (testing) Number of neurons in the hidden layer 10 Table 5: Final weights in between the layers No. of neurons Weights Input to hidden layer Hidden tooutput layer w1 w2 w3 b w4 1 -0.682 -1.572 1.929 3.529 0.680 2 -0.865 1.620 -2.115 2.483 -0.576 3 3.045 0.934 -0.191 -1.179 0.111 4 1.722 -2.097 1.179 -1.184 -0.253 5 -1.826 -1.470 1.863 0.273 -0.183 6 -0.027 -0.448 3.025 -0.131 1.113 7 -0.619 2.955 0.332 -0.694 0.268 8 0.855 -1.318 -2.373 1.895 0.036 9 1.645 0.306 -2.767 2.418 0.939 10 1.822 -1.597 -1.613 3.188 -0.638 The output performance of the established ANN is examined on the basis of the regression-correlation coefficient (R value) between the predicted values and the experimental values. The predicted responses of the ANN model are in excellent conformity with the expe- rimental values, i.e., the correlation regression coeffi- cients of 100 % for the training set, 98.41 % for the validation set and 94.80 % for the testing set are achieved as shown in Figure 7. It was concluded that the feed-forward back-propagation algorithm of the configu- ration of 3-10-1 gives the most excellent results for the prediction of the bushing length as shown in Table 6. The graphical representation of the experimental and predicted values of the bushing length by the ANN is shown in Figure 8. It shows that the ANN prediction values are very close to the experimental values of the bushing length. The average prediction error between the experimental and predicted values of the bushing length is 1.843 %. Therefore, it is demonstrated that the developed ANN model is suitable for forecasting the bushing length of drilled holes in thermal drilling of galvanized steel, having the highest accuracy of 98.157 %. N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 819 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 8: Correlation between experimental and predicted values Table 6: Comparison of experimental and predicted values of the bushing length Expt No. Bushing length (mm) Prediction error (%)Experimental value Predicted ANN value 1 2.54 2.485 2.165 2 3.42 3.427 0.204 3 4.38 4.252 2.922 4 3.24 2.618 19.197 5 3.21 3.336 3.925 6 4.41 4.425 0.340 7 2.97 2.843 4.276 8 4.87 4.677 3.963 9 5.26 5.517 4.886 10 2.67 2.385 10.674 11 3.84 3.714 3.281 12 4.02 4.275 6.343 13 2.91 2.905 0.172 14 3.83 3.814 0.418 15 4.42 4.615 4.412 16 2.94 2.964 0.816 17 4.31 4.284 0.603 18 5.57 5.599 0.521 19 3.55 3.738 5.296 20 4.64 4.585 1.185 21 5.84 5.893 0.907 22 2.78 2.760 0.719 23 3.85 3.875 0.649 24 5.92 5.836 1.419 25 2.42 2.755 13.843 26 4.42 4.450 0.679 27 5.61 6.007 7.077 5.2 Optimization of the bushing length using a GA The optimization of the thermal drilling process is performed using the GA in MATLAB toolbox, with the intention of enhancing the effectiveness of the drilling process in order to achieve high structural rigidity for fasteners in joining situations.35,36 This algorithm makes a binary coding system to characterize the variables such as rotational speed (RS), tool angle (TA) and workpiece thickness (WT). All of the process variables are symbo- lized by a ten-bit binary equivalent. In the chromosome, the process variables are represented as a substring. The GA employs different types of crossover and mutation operators to predict the maximum values of the bushing length. At this time, the second-order mathematical model is considered as an objective function with the aim of maximizing the output bushing length (BL). In this model, the rotational speed (RS), tool angle (TA) and workpiece thickness (WT) are considered as the input parameters. The aim of the optimization is to maximize the bushing length; however, usually, a GA is used to achieve the minimum function value for a mini- mization problem. Hence, in the present situation, the objective function was converted into a minimization problem. A unity negative factor is multiplied to the objective function in the larger-the-best-type bushing length (L) of the responses characteristic of thermal drilling to make them minimize the type objective. The parameters used in the GA technique are shown in Table 7. Equation (9) is a developed regression model of the bushing length and it is used as the fitness function or objective function for this problem. It is developed with the knowledge gained form ANN values.37,38 Table 7: GA parameters Population type Double vector Population size 100 Scaling function Rank Selection function Roulette wheel Elite count 2 Crossover fraction 0.8 Mutation function Adaptive feasible Crossover function Heuristic Stall generations 50 Maximum bushing length (L) = 6.56203 – 4.64305 × 10–4 × RS – 0.210894 × TA – 1.04080 × WT + 3.54118 × 10–7 × RS2 – 0.00413827 – TA2 – 0.282222 × WT2 – 5.82394 × 10–5 × RS × TA + 0.000668277 × RS × WT + 0.0584444 × TA × WT (9) is subjected to constrained variables such as: 1944 min–1 = RS = 3600 min–1 (RS – rotational speed) 30° = TA = 45° (TA – tool angle) 1 mm = WT = 2 mm (WT – workpiece thickness). Figure 9 shows the bushing-length fitness-function plot obtained with the genetic algorithm. The negative sign from the final result can be eliminated to get the maximum value of the bushing length. The optimum result is achieved at the 52nd iteration of the genetic algorithm. It is found that the mean fitness value is 5.7472 mm, whereas the best fitness value is 5.7584 mm. Table 8 presents the optimized and experimental value of the bushing length. Very close agreement between the optimized and experimental values of the bushing length is obtained. It confirms the potential applicability of these GA techniques for the industry-related problems. Figure 10 shows the cross-section of the bushing length of a hole drilled in galvanized steel under the condition of the optimal process parameter. The bushing length can offer a longer contact-surface area, which can uphold a shaft securely, and the drilled-hole inner surface has a thermally affected zone. Table 8: Comparison of predicted and experimental values of the bushing length Optimum drilling input process parameters Output para- meters Rota- tional speed (min–1) Tool angle (°) Work- piece thickness (mm) Bushing length (mm) Parameters optimized with genetic algorithm 3552.461 45 2 5.758 Experimentally used parameters 3600 45 2 5.6 N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 820 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 9: Fitness-function plot of the bushing length from the genetic algorithm Figure 10: Cross-section of a thermally drilled hole 5.3 Microstructural investigation Figure 11 shows a macrograph obtained with scanning electron microscopy at a magnification of 17×. The scanning electron microscope was utilized to observe and investigate the inside surface of a hole drilled in galvanized steel. During the thermal-drilling operation, the rubbing action between the interface of the thermal drill and the galvanized-steel workpiece led to a rise in the temperature of up to 798 °C as recorded on-line during the experimentation. Due to such fric- tional heating and application of the axial tensile force, Luders bands were formed to relieve the high amount of the internal stresses experienced in the galvanized steel.39,40 Luders bands started to form at the tail end of the bushing at a distance of 914 μm, after the piercing of a hole with the center zone of the thermal drill. Until a distance of 1725 μm, stretcher-strain markings such as Luders bands are observed in the micrograph shown in Figure 11. The width of each band thus formed is different because of non-uniform yielding of the galva- nized steel during the contact with the cylindrical zone of the thermal drill. 6 CONCLUSIONS The present work demonstrates the possibility of thermal drilling of galvanized steel that has tremendous applications in the domain of automobile and aerospace engineering. Owing to its importance, the mechanism and formation of the bushing length is studied. Due to frictional heating and applied axial force, internal stresses tend to increase in that region and subsequently lead to the formation of Luders bands at the tail end of the bushing. The formation of these stretcher-strain marks is due to discontinuous non-uni- form yielding of the galvanized steel. The relationship between the input parameters such as the rotational speed, tool angle and workpiece thick- ness, and the output parameters like the bushing length is modeled through an ANN technique. The developed ANN model is appropriately incorporated with the GA to optimize the thermal-drilling process parameters. A good correlation was observed between the experimental measurements and the predicted optimum values. This shows that the ANN model combined with the GA can be successfully applied to find the optimum conditions for achieving the maximum bushing length in the thermal drilling of galvanized steel. The modeling and optimization are valid for one material and coating only. For different materials, the building of a new ANN model is required, and the gene- tic optimization is to be performed again. Acknowledgements The authors are grateful for the financial grant by the Mepco Schlenk Engineering College (Autonomous), Sivakasi, under the Students Project Scheme (Letter No. OF/EDC/2840/2015-2016 dated 24.10.2015) for establi- shing an experimental set-up. The authors acknowledge the support and encouragement by Dr. S. Arivazhagan, Principal and Dr. P. Nagaraj, Head of Mechanical Engineering, towards this work. 7 REFERENCES 1 N. R. J. Hynes, M. V. Maheshwaran, Numerical analysis on thermal drilling of aluminum metal matrix composite, AIP. Conf. Proc., 1728 (2016), 1–5, doi:10.1063/1.4946597 2 O. Cebeli, D. Zulkuf, Investigate the friction drilling of aluminium alloys according to the thermal conductivity, Tem. J., 2 (2013) 1, 93–101 3 S. F. Miller, Experimental analysis and numerical modeling of the friction drilling process, Thesis, University of Michigan, 2006 4 N. R. J. Hynes, M. Muthukumaran, N. Rakesh, C. K. Gurubaran, Numerical analysis in friction drilling of AISI1020 steel and AA 6061 T6 alloy, Recent Advances in Environmental and Earth Sciences and Economics, 39 (2015), 145–149 5 A. H. Streppel, H. J. J Kals, Flow drilling: a preliminary analysis of a new bush-making operation, CIRP Ann. Manuf. Techn., 32 (1983) 1, 167–171, doi:10.1016/S0007-8506(07)63383-6 N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 821 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 11: Microstructure of a thermally drilled hole 5.2 Optimization of the bushing length using a GA The optimization of the thermal drilling process is performed using the GA in MATLAB toolbox, with the intention of enhancing the effectiveness of the drilling process in order to achieve high structural rigidity for fasteners in joining situations.35,36 This algorithm makes a binary coding system to characterize the variables such as rotational speed (RS), tool angle (TA) and workpiece thickness (WT). All of the process variables are symbo- lized by a ten-bit binary equivalent. In the chromosome, the process variables are represented as a substring. The GA employs different types of crossover and mutation operators to predict the maximum values of the bushing length. At this time, the second-order mathematical model is considered as an objective function with the aim of maximizing the output bushing length (BL). In this model, the rotational speed (RS), tool angle (TA) and workpiece thickness (WT) are considered as the input parameters. The aim of the optimization is to maximize the bushing length; however, usually, a GA is used to achieve the minimum function value for a mini- mization problem. Hence, in the present situation, the objective function was converted into a minimization problem. A unity negative factor is multiplied to the objective function in the larger-the-best-type bushing length (L) of the responses characteristic of thermal drilling to make them minimize the type objective. The parameters used in the GA technique are shown in Table 7. Equation (9) is a developed regression model of the bushing length and it is used as the fitness function or objective function for this problem. It is developed with the knowledge gained form ANN values.37,38 Table 7: GA parameters Population type Double vector Population size 100 Scaling function Rank Selection function Roulette wheel Elite count 2 Crossover fraction 0.8 Mutation function Adaptive feasible Crossover function Heuristic Stall generations 50 Maximum bushing length (L) = 6.56203 – 4.64305 × 10–4 × RS – 0.210894 × TA – 1.04080 × WT + 3.54118 × 10–7 × RS2 – 0.00413827 – TA2 – 0.282222 × WT2 – 5.82394 × 10–5 × RS × TA + 0.000668277 × RS × WT + 0.0584444 × TA × WT (9) is subjected to constrained variables such as: 1944 min–1 = RS = 3600 min–1 (RS – rotational speed) 30° = TA = 45° (TA – tool angle) 1 mm = WT = 2 mm (WT – workpiece thickness). Figure 9 shows the bushing-length fitness-function plot obtained with the genetic algorithm. The negative sign from the final result can be eliminated to get the maximum value of the bushing length. The optimum result is achieved at the 52nd iteration of the genetic algorithm. It is found that the mean fitness value is 5.7472 mm, whereas the best fitness value is 5.7584 mm. Table 8 presents the optimized and experimental value of the bushing length. Very close agreement between the optimized and experimental values of the bushing length is obtained. It confirms the potential applicability of these GA techniques for the industry-related problems. Figure 10 shows the cross-section of the bushing length of a hole drilled in galvanized steel under the condition of the optimal process parameter. The bushing length can offer a longer contact-surface area, which can uphold a shaft securely, and the drilled-hole inner surface has a thermally affected zone. Table 8: Comparison of predicted and experimental values of the bushing length Optimum drilling input process parameters Output para- meters Rota- tional speed (min–1) Tool angle (°) Work- piece thickness (mm) Bushing length (mm) Parameters optimized with genetic algorithm 3552.461 45 2 5.758 Experimentally used parameters 3600 45 2 5.6 N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 820 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 9: Fitness-function plot of the bushing length from the genetic algorithm Figure 10: Cross-section of a thermally drilled hole 5.3 Microstructural investigation Figure 11 shows a macrograph obtained with scanning electron microscopy at a magnification of 17×. The scanning electron microscope was utilized to observe and investigate the inside surface of a hole drilled in galvanized steel. During the thermal-drilling operation, the rubbing action between the interface of the thermal drill and the galvanized-steel workpiece led to a rise in the temperature of up to 798 °C as recorded on-line during the experimentation. Due to such fric- tional heating and application of the axial tensile force, Luders bands were formed to relieve the high amount of the internal stresses experienced in the galvanized steel.39,40 Luders bands started to form at the tail end of the bushing at a distance of 914 μm, after the piercing of a hole with the center zone of the thermal drill. Until a distance of 1725 μm, stretcher-strain markings such as Luders bands are observed in the micrograph shown in Figure 11. The width of each band thus formed is different because of non-uniform yielding of the galva- nized steel during the contact with the cylindrical zone of the thermal drill. 6 CONCLUSIONS The present work demonstrates the possibility of thermal drilling of galvanized steel that has tremendous applications in the domain of automobile and aerospace engineering. Owing to its importance, the mechanism and formation of the bushing length is studied. Due to frictional heating and applied axial force, internal stresses tend to increase in that region and subsequently lead to the formation of Luders bands at the tail end of the bushing. The formation of these stretcher-strain marks is due to discontinuous non-uni- form yielding of the galvanized steel. The relationship between the input parameters such as the rotational speed, tool angle and workpiece thick- ness, and the output parameters like the bushing length is modeled through an ANN technique. The developed ANN model is appropriately incorporated with the GA to optimize the thermal-drilling process parameters. A good correlation was observed between the experimental measurements and the predicted optimum values. This shows that the ANN model combined with the GA can be successfully applied to find the optimum conditions for achieving the maximum bushing length in the thermal drilling of galvanized steel. The modeling and optimization are valid for one material and coating only. For different materials, the building of a new ANN model is required, and the gene- tic optimization is to be performed again. Acknowledgements The authors are grateful for the financial grant by the Mepco Schlenk Engineering College (Autonomous), Sivakasi, under the Students Project Scheme (Letter No. OF/EDC/2840/2015-2016 dated 24.10.2015) for establi- shing an experimental set-up. The authors acknowledge the support and encouragement by Dr. S. Arivazhagan, Principal and Dr. P. Nagaraj, Head of Mechanical Engineering, towards this work. 7 REFERENCES 1 N. R. J. Hynes, M. V. Maheshwaran, Numerical analysis on thermal drilling of aluminum metal matrix composite, AIP. Conf. Proc., 1728 (2016), 1–5, doi:10.1063/1.4946597 2 O. Cebeli, D. Zulkuf, Investigate the friction drilling of aluminium alloys according to the thermal conductivity, Tem. J., 2 (2013) 1, 93–101 3 S. F. Miller, Experimental analysis and numerical modeling of the friction drilling process, Thesis, University of Michigan, 2006 4 N. R. J. Hynes, M. Muthukumaran, N. Rakesh, C. K. Gurubaran, Numerical analysis in friction drilling of AISI1020 steel and AA 6061 T6 alloy, Recent Advances in Environmental and Earth Sciences and Economics, 39 (2015), 145–149 5 A. H. Streppel, H. J. J Kals, Flow drilling: a preliminary analysis of a new bush-making operation, CIRP Ann. Manuf. Techn., 32 (1983) 1, 167–171, doi:10.1016/S0007-8506(07)63383-6 N. R. J. 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Proc. Mech. Eng., (2015), 1–7, doi:10.1177/0954408915614300 21 P. Krasauskas, S. Kilikevi~ius, R. ^esnavi~ius, D. Pa~enga, Expe- rimental analysis and numerical simulation of the stainless AISI 304 steel friction drilling process, Mecha., 20 (2014) 6, 590–595, doi:10.5755/j01.mech.20.6.8664 22 P. D. Pantawane, B. B. Ahuja, Parametric analysis and modelling of friction drilling process on AISI 1015, Int. J. Mecha. Manuf. Sys., 7 (2014) 1, 60–79, doi:10.1504/IJMMS.2014.062771 23 G. Somasundaram, B. S. Rajendra, K. Palanikumar, Modeling and analysis of roundness error in friction drilling of aluminum silicon carbide metal matrix composite, J. Compos. Mater., 46 (2011) 2, 169–181, doi:10.1177/0021998311410493 24 S. Changyu, W. Lixia, C. Wei, W. Jinxing, Optimization for injection molding process conditions of the refrigeratory top cover using com- bination method of artificial neural network and genetic algorithms, Poly. Plast. Tech. Eng., 46 (2007), 105–112, doi:10.1080/ 03602550601152853 25 S. R. Karnik, V. N. Gaitone, Development of artificial neural network models to study the effect of process parameters on burr size in drilling, Int. J. Adv. Manuf. Technol., 39 (2008), 439–453, doi:10.1007/s00170-007-1231-5 26 R. S. Mamilla, S. Mondal, J. Ramkumar, V. K. Jain, Experimental investigations and modeling of drill bit-guided abrasive flow finishing (DBG-AFF) process, Int. J. Adv. Manuf. Technol., 42 (2009), 678–688, doi:10.1007/s00170-008-1642-y 27 S. Assarzadeh, M. Ghoreishi, Neural-network-based modeling and optimization of the electro-discharge machining process, Int. J. Adv. Manuf. Technol., 39 (2008), 488–500, doi:10.1007/s00170-007- 1235-1 28 O. Babur, O. Hasan, K. Hasan, Optimum surface roughness in end milling Inconel 718 by coupling neural network model and genetic algorithm, Int. J. Adv. Manuf. Technol., 27 (2005), 234–241, doi:10.1007/s00170-004-2175-7 29 S. Sarkar, S. Mitra, B. Bhattacharyya, Parametric optimisation of wire electrical discharge machining of ã titanium aluminide alloy through an artificial neural network model, Int. J. Adv. Manuf. Technol., 27 (2006), 501–508, doi:10.1007/s00170-004-2203-7 30 A. K. Singh, S. S. Panda, S. K. Pal, D. Chakraborty, Predicting drill wear using an artificial neural network, Int. J. Adv. Manuf. Technol., 28 (2006), 456–462, doi:10.1007/s00170-004-2376-0 31 S. M. Hamidinejad, F. Kolahan, A. H. Kokabi, The modeling and process analysis of resistance spot welding on galvanized steel sheets used in car body manufacturing, Mater. Des., 34 (2012), 759–767, doi:10.1016/j.matdes.2011.06.064 32 Y. Lieh-Dai, K. Wei-Liang, C. Han-Ming, W. Der-An, L. Yan- Cherng, Mar-M247, Haynes-230 & Inconel-718 study of machining characteristics for Ni-based super alloys on friction drilling, Adv. Mat. Res., 459 (2012), 632–637, doi:10.4028/www.scientific.net/ AMR.459.632 33 J. B. Peter, C. J. Brian, Q. Jun, Feasibility of thermally drilling automotive alloy sheet, castings, and hydro formed shapes, (2007), http://web.ornl.gov/info/reports/2007/3445605662084.pdf, 08/2016 34 P. Sathiya, K. Panneerselvam, M. Y. Abdul Jaleel, Optimization of laser welding process parameters for super austenitic stainless steel using artificial neural networks and genetic algorithm, Mater. Des., 36 (2012), 490–498, doi:10.1016/j.matdes.2011.11.028 35 K. Girish, S. S. Kuldip, Predictive modelling and optimization of machining parameters to minimize surface roughness using artificial neural network coupled with genetic algorithm, Proc. CIRP., 31 (2015), 453–458, doi:10.1016/j.procir.2015.03.043 36 S. S. Kuldip, S. Sachin, K. Girish, Optimization of machining parameters to minimize surface roughness using integrated ANN-GA approach, Proc. CIRP., 29 (2015), 305–310, doi:10.1016/j.procir. 2015.02.002 37 D. S. Nagesh, G. L. Datta, Genetic algorithm for optimization of welding variables for height to width ratio and application of ANN for prediction of bead geometry for TIG welding process, App. Soft. Comp., 10 (2010), 897–907, doi:10.1016/j.asoc.2009.10.007 38 M. Z. Azlan, H. Habibollah, S. Safian, Genetic algorithm and simulated annealing to estimate optimal process parameters of the abrasive water jet machining, Eng. Comp., 27 (2011), 251–259, doi:10.1007/s00366-010-0195-5 39 V. S. Ananthan, E. O. Hall, Macroscopic aspects of Luders band deformation in mild steel, Actametall. Mater., 39 (1991) 12, 3153–3160, doi:10.1016/0956-7151(91)90049-7 40 F. H. Julian, K. Stelios, On the effect of Luders bands on the bending of steel tubes. Part II: Analysis, Int. J. Solids. Struct., 48 (2011), 3285–3298, doi:10.1016/j.ijsolstr.2011.07.012 N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 822 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... 823–829 GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX IN A DYE-SENSITIZED SOLAR CELL @ELIRNI POLISAHARID KOT ELEKTROLITNA OSNOVA V SOLARNIH CELICAH, OB^UTLJIVIH NA BARVILA Jose Paolo Bantang1, Drexel Camacho1,2 1De La Salle University, Chemistry Department, 2401 Taft Avenue, 1004 Manila, Philippines 2De La Salle University, Organic Materials and Interfaces Unit, CENSER, 2401 Taft Avenue, 1004 Manila, Philippines drexel.camacho@dlsu.edu.ph Prejem rokopisa – received: 2016-09-30; sprejem za objavo – accepted for publication: 2017-03-16 doi:10.17222/mit.2016.294 Hydrophilic polysaccharide, -carrageenan, was utilized as the polymer matrix in gel-electrolyte systems for dye-sensitized solar-cell (DSSC) applications. The influence of the solvent system was investigated to optimize the solubility of -carrageenan and tetrabutylammonium-iodide (TBAI)/I2 electrolytes by minimizing the water content because of its unfavorable effect on DSSCs. We report herein that two solvent systems, a water/acetonitrile mixed solvent and DMSO, were found to effectively dissolve the components. The composite natures of the -carrageenan-electrolyte systems in these solvents were confirmed with an FTIR analysis. The presence of -carrageenan did not impede the electrochemical properties of the electrolytes, as confirmed with cyclic voltammetry, electrochemical impedance spectroscopy and linear sweep voltammetry. The incorporation of the gel electrolytes in DSSCs showed that the DMSO system exhibited better solar-cell efficiency compared to the mixed-solvent system. Keywords: dye-sensitized solar cell, -carrageenan, gel electrolyte, electrochemical impedance spectroscopy, ionic conductivity Hidrofilni polisaharid, -karagen, je bil uporabljen kot polimerna osnova v `elirno elektrolitskih sistemih za uporabo v barvno ob~utljivih son~nih celicah (angl. DSSC). Da bi izbolj{ali topnost -karagena in elektrolitov tetrabutilamonijevega iodida (TBAI)/I2 z zmanj{anjem vsebnosti vode, zaradi njenega ne`elenega u~inka na DSSC, je bila preiskovana ob~utljivost sistema topil. ^lanek poro~a, da sta bila najdena dva nova sistema topil, me{anica topila voda/acetonitril in DMSO, ki u~inkovito raztopita komponente. Lastnosti oz. obna{anje kompozitov elektrolitskega sistema -karagen v teh raztopinah, so bile potrjene z FTIR-analizo. Prisotnost -karagenskega elektrolitnega sistem ni predstavljala ovire za imepdan~no spektroskopijo in lienarno "sweep" voltametrijo. Vklju~itev gel-elektrolitov v DSSC je pokazala, da DMSO-sistem ka`e bolj{o solarno celi~no u~inkovitost kot sistem me{anih topil. Klju~ne besede: barvno ob~utljive son~ne celice, rastlinska `elatina -karagen, gel-elektrolit, elektrokemi~na impedan~na spekroskopija, ionska prevodnost 1 INTRODUCTION Gel-polymer electrolytes are promising materials for a potential incorporation in electrochemical devices1 such as dye-sensitized solar cells (DSC). This is because their mechanical behavior is that of solids, yet their inter- nal structure is flexible and their conductivity behavior resembles that of a liquid state allowing a good elec- trode/electrolyte contact.2,3 Moreover, its ease of fabrica- tion allows more tunability in designing systems for specific applications. Natural polysaccharide is a potential matrix for polymer electrolytes owing to its hydrophilic and gel-forming capacity that traps the solvent together with the redox couple inside the polymer matrix. The high water-retention property of polysaccharide-based poly- mer-electrolyte systems was shown to promote good ionic conductivities and thermal stability.4,5 Agarose,6,7 cellulose and its derivatives8,9 and -carrageenan10 are the gel-forming natural polysaccharides that have been used as polymer-electrolyte (PE) systems for dye-sensi- tized-solar-cell (DSSC) applications. To maximize the water-retention property of these polysaccharide-based gel systems, an aqueous medium is necessary. However, we can see that water disturbs the interfacial attachment of dyes to TiO2. Thus, it is highly desirable to develop a polysaccharide-based electrolyte system in a smaller amount of water or in a non-aqueous medium. -Carrageenan, which is composed mainly of disaccharide units of -D-galactopyranose with either an -D-galactopyranose or 3,6-anhydrogalactose is a pro- mising polysaccharide matrix for electrolytes due to its hydrophilic, linear and sulfated properties. The electro- static interactions of ions with hydroxyl groups and sulfate groups in the main chain are essential to the conduction mechanism of a polysaccharide electrolyte system.11,12 The promising features of -carrageenan as a polymer-electrolyte system can be potentially applied to DSSCs. Solid-state DSSCs using carrageenan-gel elec- trolyte systems were fabricated by soaking the aqueous polymer gel with the redox electrolytes11 and forming thin polymer membranes13 leading to decent efficiencies. To date, studies on the preparation and characterization of carrageenan-electrolyte systems have been limited Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 823 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 67.017:621.3.035.22:544.623 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)823(2017) 6 M. Kerkhofs, M. Van Stappen, The performance of (Ti, Al) N-coated flow drills, Surf. Coat. Technol., 68 (1994) 69, 741–746 7 S. F. Miller, P. J. Blau, A. J. Shih, Microstructural alterations associated with friction drilling of steel, aluminum, and titanium, J. Mater. Eng. Perform., 14 (2005), 647–653, doi:10.1361/ 105994905X64558 8 S. F. Miller, R. Li, H. Wang, A. J. Shih, Experimental and numerical analysis of the friction drilling process, J. Manuf. Sci. Eng., 128 (2006), 802–810, doi:10.1115/1.2193554 9 S. F. Miller, J. Tao, A. J. Shih, Friction drilling of cast metals, Int. J. Mach. Tools Manuf., 46 (2006), 1526–1535, doi:10.1016/ j.ijmachtools.2005.09.003 10 S. F. Miller, A. J. Shih, Thermo-mechanical finite element modeling of the friction drilling process, J. Manuf. Sci. Eng., 129 (2007), 531–538 11 S. F. Miller, P. J. Blau, A. J. Shih, Tool wear in friction drilling, Int. J. Mach. Tools Manuf., 47 (2007), 1636–1645, doi:10.1016/ j.ijmachtools.2006.10.009 12 S. M. Lee, H. M. Chow, B. H. Yan, Friction drilling of IN-713LC cast superalloy, Mater. Manuf. Proces., 22 (2007), 893–897, doi:10.1080/10426910701451697 13 H. M. Chow, S. M. Lee, L. D. Yang, Machining characteristic study of friction drilling on AISI 304 stainless steel, J. Mater. Process. Technol., 207 (2008), 180–186, doi:10.1016/j.jmatprotec.2007. 12.064 14 B. Mehmet, C. Ibrahim, G. Mustafa, O. Feridun, Application of the taguchi method to optimize the cutting conditions in hard turning of a ring bore, Mater. Tech., 49 (2015) 5, 765–772, doi:10.17222/mit. 2014.246 15 S. M. Lee, H. M. Chow, F. Y. Huang, B. H. Yan, Friction drilling of austenitic stainless steel by uncoated and PVD AlCrN- and TiAlN-coated tungsten carbide tools, Int. J. Mach. Tools Manuf., 49 (2009), 81–88,doi:10.1016/j.ijmachtools.2008.07.012 16 W. L. Ku, C. L. Hung, S. M. Lee, H. M. Chow, Optimization in ther- mal friction drilling for SUS 304 stainless steel, Int. J. Adv. Manuf. Technol., 53 (2011), 935–944, doi:10.1007/s00170-010-2899-5 17 M. Folea, D. Schlegel, E. Gete, C. Langlade, A. Roman, Preliminary tests on flow drilling of maraging steels, Acad. J. Manuf. Eng. 10 (2012) 4, 42–47 18 T. K. Mehmet, A. Alaattin, B. Bertan, K. A. Hamza, An experimental study on friction drilling of ST12 steel, T. Can. Soc. Mech. Eng., 38 (2014), 319–329 19 D. Biermann, Y. Liu, Innovative flow drilling on magnesium wrought alloy AZ31, Proc. CIRP., 18 (2014), 209–214, doi:10.1016/j.procir. 2014.06.133 20 B. B. Mehmet, G. Kadir, G. Arif, Three-dimensional finite element model of friction drilling process in hot forming processes, J. Proc. Mech. Eng., (2015), 1–7, doi:10.1177/0954408915614300 21 P. Krasauskas, S. Kilikevi~ius, R. ^esnavi~ius, D. Pa~enga, Expe- rimental analysis and numerical simulation of the stainless AISI 304 steel friction drilling process, Mecha., 20 (2014) 6, 590–595, doi:10.5755/j01.mech.20.6.8664 22 P. D. Pantawane, B. B. Ahuja, Parametric analysis and modelling of friction drilling process on AISI 1015, Int. J. Mecha. Manuf. Sys., 7 (2014) 1, 60–79, doi:10.1504/IJMMS.2014.062771 23 G. Somasundaram, B. S. Rajendra, K. Palanikumar, Modeling and analysis of roundness error in friction drilling of aluminum silicon carbide metal matrix composite, J. Compos. Mater., 46 (2011) 2, 169–181, doi:10.1177/0021998311410493 24 S. Changyu, W. Lixia, C. Wei, W. Jinxing, Optimization for injection molding process conditions of the refrigeratory top cover using com- bination method of artificial neural network and genetic algorithms, Poly. Plast. Tech. Eng., 46 (2007), 105–112, doi:10.1080/ 03602550601152853 25 S. R. Karnik, V. N. Gaitone, Development of artificial neural network models to study the effect of process parameters on burr size in drilling, Int. J. Adv. Manuf. Technol., 39 (2008), 439–453, doi:10.1007/s00170-007-1231-5 26 R. S. Mamilla, S. Mondal, J. Ramkumar, V. K. Jain, Experimental investigations and modeling of drill bit-guided abrasive flow finishing (DBG-AFF) process, Int. J. Adv. Manuf. Technol., 42 (2009), 678–688, doi:10.1007/s00170-008-1642-y 27 S. Assarzadeh, M. Ghoreishi, Neural-network-based modeling and optimization of the electro-discharge machining process, Int. J. Adv. Manuf. Technol., 39 (2008), 488–500, doi:10.1007/s00170-007- 1235-1 28 O. Babur, O. Hasan, K. Hasan, Optimum surface roughness in end milling Inconel 718 by coupling neural network model and genetic algorithm, Int. J. Adv. Manuf. Technol., 27 (2005), 234–241, doi:10.1007/s00170-004-2175-7 29 S. Sarkar, S. Mitra, B. Bhattacharyya, Parametric optimisation of wire electrical discharge machining of ã titanium aluminide alloy through an artificial neural network model, Int. J. Adv. Manuf. Technol., 27 (2006), 501–508, doi:10.1007/s00170-004-2203-7 30 A. K. Singh, S. S. Panda, S. K. Pal, D. Chakraborty, Predicting drill wear using an artificial neural network, Int. J. Adv. Manuf. Technol., 28 (2006), 456–462, doi:10.1007/s00170-004-2376-0 31 S. M. Hamidinejad, F. Kolahan, A. H. Kokabi, The modeling and process analysis of resistance spot welding on galvanized steel sheets used in car body manufacturing, Mater. Des., 34 (2012), 759–767, doi:10.1016/j.matdes.2011.06.064 32 Y. Lieh-Dai, K. Wei-Liang, C. Han-Ming, W. Der-An, L. Yan- Cherng, Mar-M247, Haynes-230 & Inconel-718 study of machining characteristics for Ni-based super alloys on friction drilling, Adv. Mat. Res., 459 (2012), 632–637, doi:10.4028/www.scientific.net/ AMR.459.632 33 J. B. Peter, C. J. Brian, Q. Jun, Feasibility of thermally drilling automotive alloy sheet, castings, and hydro formed shapes, (2007), http://web.ornl.gov/info/reports/2007/3445605662084.pdf, 08/2016 34 P. Sathiya, K. Panneerselvam, M. Y. Abdul Jaleel, Optimization of laser welding process parameters for super austenitic stainless steel using artificial neural networks and genetic algorithm, Mater. Des., 36 (2012), 490–498, doi:10.1016/j.matdes.2011.11.028 35 K. Girish, S. S. Kuldip, Predictive modelling and optimization of machining parameters to minimize surface roughness using artificial neural network coupled with genetic algorithm, Proc. CIRP., 31 (2015), 453–458, doi:10.1016/j.procir.2015.03.043 36 S. S. Kuldip, S. Sachin, K. Girish, Optimization of machining parameters to minimize surface roughness using integrated ANN-GA approach, Proc. CIRP., 29 (2015), 305–310, doi:10.1016/j.procir. 2015.02.002 37 D. S. Nagesh, G. L. Datta, Genetic algorithm for optimization of welding variables for height to width ratio and application of ANN for prediction of bead geometry for TIG welding process, App. Soft. Comp., 10 (2010), 897–907, doi:10.1016/j.asoc.2009.10.007 38 M. Z. Azlan, H. Habibollah, S. Safian, Genetic algorithm and simulated annealing to estimate optimal process parameters of the abrasive water jet machining, Eng. Comp., 27 (2011), 251–259, doi:10.1007/s00366-010-0195-5 39 V. S. Ananthan, E. O. Hall, Macroscopic aspects of Luders band deformation in mild steel, Actametall. Mater., 39 (1991) 12, 3153–3160, doi:10.1016/0956-7151(91)90049-7 40 F. H. Julian, K. Stelios, On the effect of Luders bands on the bending of steel tubes. Part II: Analysis, Int. J. Solids. Struct., 48 (2011), 3285–3298, doi:10.1016/j.ijsolstr.2011.07.012 N. R. J. HYNES et al.: OPTIMUM BUSHING LENGTH IN THERMAL DRILLING OF GALVANIZED STEEL ... 822 Materiali in tehnologije / Materials and technology 51 (2017) 5, 813–822 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... 823–829 GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX IN A DYE-SENSITIZED SOLAR CELL @ELIRNI POLISAHARID KOT ELEKTROLITNA OSNOVA V SOLARNIH CELICAH, OB^UTLJIVIH NA BARVILA Jose Paolo Bantang1, Drexel Camacho1,2 1De La Salle University, Chemistry Department, 2401 Taft Avenue, 1004 Manila, Philippines 2De La Salle University, Organic Materials and Interfaces Unit, CENSER, 2401 Taft Avenue, 1004 Manila, Philippines drexel.camacho@dlsu.edu.ph Prejem rokopisa – received: 2016-09-30; sprejem za objavo – accepted for publication: 2017-03-16 doi:10.17222/mit.2016.294 Hydrophilic polysaccharide, -carrageenan, was utilized as the polymer matrix in gel-electrolyte systems for dye-sensitized solar-cell (DSSC) applications. The influence of the solvent system was investigated to optimize the solubility of -carrageenan and tetrabutylammonium-iodide (TBAI)/I2 electrolytes by minimizing the water content because of its unfavorable effect on DSSCs. We report herein that two solvent systems, a water/acetonitrile mixed solvent and DMSO, were found to effectively dissolve the components. The composite natures of the -carrageenan-electrolyte systems in these solvents were confirmed with an FTIR analysis. The presence of -carrageenan did not impede the electrochemical properties of the electrolytes, as confirmed with cyclic voltammetry, electrochemical impedance spectroscopy and linear sweep voltammetry. The incorporation of the gel electrolytes in DSSCs showed that the DMSO system exhibited better solar-cell efficiency compared to the mixed-solvent system. Keywords: dye-sensitized solar cell, -carrageenan, gel electrolyte, electrochemical impedance spectroscopy, ionic conductivity Hidrofilni polisaharid, -karagen, je bil uporabljen kot polimerna osnova v `elirno elektrolitskih sistemih za uporabo v barvno ob~utljivih son~nih celicah (angl. DSSC). Da bi izbolj{ali topnost -karagena in elektrolitov tetrabutilamonijevega iodida (TBAI)/I2 z zmanj{anjem vsebnosti vode, zaradi njenega ne`elenega u~inka na DSSC, je bila preiskovana ob~utljivost sistema topil. ^lanek poro~a, da sta bila najdena dva nova sistema topil, me{anica topila voda/acetonitril in DMSO, ki u~inkovito raztopita komponente. Lastnosti oz. obna{anje kompozitov elektrolitskega sistema -karagen v teh raztopinah, so bile potrjene z FTIR-analizo. Prisotnost -karagenskega elektrolitnega sistem ni predstavljala ovire za imepdan~no spektroskopijo in lienarno "sweep" voltametrijo. Vklju~itev gel-elektrolitov v DSSC je pokazala, da DMSO-sistem ka`e bolj{o solarno celi~no u~inkovitost kot sistem me{anih topil. Klju~ne besede: barvno ob~utljive son~ne celice, rastlinska `elatina -karagen, gel-elektrolit, elektrokemi~na impedan~na spekroskopija, ionska prevodnost 1 INTRODUCTION Gel-polymer electrolytes are promising materials for a potential incorporation in electrochemical devices1 such as dye-sensitized solar cells (DSC). This is because their mechanical behavior is that of solids, yet their inter- nal structure is flexible and their conductivity behavior resembles that of a liquid state allowing a good elec- trode/electrolyte contact.2,3 Moreover, its ease of fabrica- tion allows more tunability in designing systems for specific applications. Natural polysaccharide is a potential matrix for polymer electrolytes owing to its hydrophilic and gel-forming capacity that traps the solvent together with the redox couple inside the polymer matrix. The high water-retention property of polysaccharide-based poly- mer-electrolyte systems was shown to promote good ionic conductivities and thermal stability.4,5 Agarose,6,7 cellulose and its derivatives8,9 and -carrageenan10 are the gel-forming natural polysaccharides that have been used as polymer-electrolyte (PE) systems for dye-sensi- tized-solar-cell (DSSC) applications. To maximize the water-retention property of these polysaccharide-based gel systems, an aqueous medium is necessary. However, we can see that water disturbs the interfacial attachment of dyes to TiO2. Thus, it is highly desirable to develop a polysaccharide-based electrolyte system in a smaller amount of water or in a non-aqueous medium. -Carrageenan, which is composed mainly of disaccharide units of -D-galactopyranose with either an -D-galactopyranose or 3,6-anhydrogalactose is a pro- mising polysaccharide matrix for electrolytes due to its hydrophilic, linear and sulfated properties. The electro- static interactions of ions with hydroxyl groups and sulfate groups in the main chain are essential to the conduction mechanism of a polysaccharide electrolyte system.11,12 The promising features of -carrageenan as a polymer-electrolyte system can be potentially applied to DSSCs. Solid-state DSSCs using carrageenan-gel elec- trolyte systems were fabricated by soaking the aqueous polymer gel with the redox electrolytes11 and forming thin polymer membranes13 leading to decent efficiencies. To date, studies on the preparation and characterization of carrageenan-electrolyte systems have been limited Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 823 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 67.017:621.3.035.22:544.623 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)823(2017) only to pure aqueous systems. A significant amount of water present in the electrolyte systems applied to DSSCs causes a decrease in the photocurrent of the device associated with dye desorption, iodate formation and a decrease in the electron lifetime.14 Hence, the amount of water must be controlled, if not eliminated. It is the objective of this paper to investigate the influence of the solvent system on carrageenan electrolyte and its performance in a DSSC. This paper reports, for the first time, on a novel carrageenan-based electrolyte system with an iodide/tri-iodide redox couple prepared using a mixed solvent system and a non-aqueous solvent. 2 MATERIALS AND METHODS 2.1 Optimization of the -carrageenan electrolyte sys- tem Different -carrageenan (Shemberg, Philippines) biopolymer gel electrolytes were optimized by varying the ratio of water/acetonitrile and the amount of polymer with a constant amount of tetrabutylammonium iodide (TBAI, Sigma Aldrich, a reagent grade of 98 %; 20 % w/w of TBAI with respect to -carrageenan) and iodine (I2, Aldrich, = 99.99 % metal basis). The molar ratio of TBAI/I2 was 4:1. The ratio of water/ACN was optimized by dissolving -carrageenan (0.2 g) with TBAI/I2 in different solvent systems consisting of 100 % water, 3:1 water/ACN, 1:1 water/ACN, 1:3 water/ACN, and 100 % ACN at 80 °C until a homogeneous solution was formed. Thin films were formed by casting solutions on plastic petri dishes and slowly drying in air. Different biopoly- mer electrolytes were prepared using the same procedure by varying the concentration of -carrageenan (0, 0.5, 1.0, 1.5 and 2 % w/v). Using the optimized parameters, -carrageenan gel electrolytes containing 20 % w/w (with respect to the amount of -carrageenan) of other salts such as potassium iodide (Sigma Aldrich) and trimethylsulfonium iodide (TMSI, Sigma Aldrich) were also prepared. A liquid-state polymer electrolyte and a gel-state polymer electrolyte based on -carrageenan were pre- pared by dissolving 1 % w/v or 2 % w/v of the poly- saccharide in 2 mL of dimethylsulfoxide at 70 °C. Tetrabutylammonium iodide (0.5 M), I2 (0.05 M) and LiI (0.1 M) were added to the -carrageenan solution. As the control, the same composition was used without adding the biopolymer. 2.2 Fabrication of dye-sensitized solar cells (DSSCs) The photoanode was prepared spreading a TiO2 paste (Solaronix, Ti-Nanoxide T/SP) on a fluorine-doped tin-oxide (FTO) conducting glass (TCO22-7) with the doctor-blading method. The active area of the photo- anode was fixed to 1 cm × 1 cm with adhesive tapes. The deposited paste was sintered at 450 °C for 30 min and cooled down slowly to room temperature. The deposited TiO2 film was soaked in a dye solution containing a ruthenium dye (Ruthenizer 535-bisTBA, Solaronix) in ethanol for 24 h. The counter electrode was prepared by sintering an FTO conducting glass coated with a pla- tinum solution (Platisol 41121, Solaronix) at 450 °C for 30 mins. The photoanode and counter electrode were sealed with a hot-melt thermoplastic material called Surlyn (DuPont, thickness of 60 μm), also acting as a spacer. A dye-sensitized solar cell was fabricated by inserting a hot gel-electrolyte solution in a pre-drilled hole on the counter electrode. Three different solar cells were fabricated containing different iodide salts (KI, TMSI, and TBAI) in a water/ACN -carrageenan matrix. A control DSSC was prepared using an acetonitrile- based liquid electrolyte (Iodolyte AN50, Solaronix). The same process was used for the fabrication of a DSSC incorporating gel electrolytes in DMSO. 2.3 Characterization The surface morphology and thickness of the thin films were characterized using scanning electron micro- scopy (JEOL JSM-5310). The electrical conductivity was measured using the 4-point probe Van der Pauw method. In brief, four wires were attached to the corners of a 1 cm × 1 cm thin film using a silver paste. The voltage at a constant current was measured. The resistivity and electrical conductivity were calculated using the following equations: Resistivity, s d V l V l f R R = + ⎛ ⎝ ⎜ ⎞ ⎠ ⎟ ⎛ ⎝ ⎜ ⎞ ⎠ ⎟ π ln 2 1 2 2 2 1 2 and Electrical conductivity,  = 1/s, where d, V, I and f R /R( )1 2 refer to the thickness, potential, current and van der Pauw function, respectively. The redox properties of the -carrageenan gel electrolyte were characterized using 3-electrode cyclic voltammetry (Powerlab/4SP potentiostat). The electrodes composed of the Ag/AgCl reference electrode, plati- num-wire counter electrode and glassy-carbon working electrode were pierced through the gel. Electrochemical impedance spectroscopy (Metrohm Autolab potentiostat PGSTAT128N) was obtained by sandwiching the gel electrolyte in between two platinum-coated FTO conducting-glass electrodes. The active area was fixed to 1 × 1 cm2 and the thickness of the spacer was about 50 μm. The frequency was set from 106 Hz to 10–1 Hz with an AC applied voltage of 10 mV. The ionic conductivity () of the gel electrolyte was calculated from the solution resistance (Rs) using Equation (3):  = L ARs (3) The solution resistance (Rs) is determined as the intercept of the Nyquist plot against the real part of the impedance or the horizontal axis. The variables L and A correspond to the thickness of the spacer and the active area, respectively. The solar-cell parameters such as the J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... 824 Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS open-circuit voltage, short-circuit current, fill factor and efficiency of the fabricated dye-sensitized solar cells with -carrageenan gel electrolyte were assessed using a 100 mW/cm2 AM 1.5 solar simulator (Abet Technolo- gies) coupled with a potentiostat to obtain a current- voltage (IV) curve. 3 RESULTS AND DISCUSSIONS 3.1 Carrageenan-electrolyte system in a mixed solvent system Acetonitrile (ACN), which is a common organic solvent used for preparing liquid electrolytes for DSSC applications, does not dissolve -carrageenan. Thus, to form -carrageenan electrolytes containing a redox couple I–/I3– (from TBAI and I2), the amount of water is crucial to dissolve the hydrophilic polysaccharide.15 Using water only as the solvent, a black film was formed with observable colorless crystals (Figure 1a) indicating the unsuccessful dissolution of both TBAI salt and iodine. Likewise, crystallization of the salt was formed on the surface of the film prepared using a 3:1 v/v water/ACN system (Figure 1b). The salt crystallization is attributed to the relative insolubility of the salt in a water-rich solvent. Moreover, the presence of the salt and iodine crystals on the film surface indicates that it fails to dissociate in the polymer matrix, leading to an un- successful incorporation of the charge-carrier redox couple I–/I3–. For a 1:1 v/v water/ACN mixed-solvent system, a uniform orange polymer-electrolyte film with- out visible crystals was formed (Figure 1c), indicating good dissolution and incorporation of the I–/I3– redox couple in the polymer matrix. No films were cast for the 1:3 v/v water/acetonitrile solvent systems and in 100 % acetonitrile because of the relative insolubility of carrageenan in an acetonitrile-rich solvent. The concentration of carrageenan in the polymer electrolyte containing 20 % w/w TBAI (w.r.t. -carra- geenan) in the 1:1 H2O/acetonitrile mixed solvent affects the property and the ionic conductivity of the electrolyte. A liquid polymer electrolyte is formed when the con- centrations of carrageenan are in a range of 0.1–0.5 % w/v. Increasing the carrageenan concentration to  1.0 % w/v gave a gel polymer electrolyte. Consequently, the ionic conductivity of the polymer electrolytes changes in response to its composite form. The ionic conductivity (Figure 2) of -carrageenan-electrolyte systems in- creases as the carrageenan concentration increases from 0.1 to 1.0 % w/v. The increase in the ionic conductivity at low -carrageenan concentrations is attributed to high free-ion concentrations due to their interaction with the functional group present in -carrageenan.16. The ionic conductivity decreases at carrageenan concentrations of 1.0–2.0 % w/v. The increase in the polymer concentra- tion causes chain entanglements, impeding the move- ment of the ions, thus, decreasing its ionic conductivity.16 Different iodide salts were incorporated in -carrageenan to prepare the gel electrolytes. Films were formed and the IR spectra of the electrolyte films con- tained typical functional-group vibrations of -carra- geenan.17 The bands at 1272 cm–1 were identified as an O=S=O symmetric vibration that would respond in the presence of a cation. A shift in the frequency was observed in the O=S=O symmetric vibration compared with the pure k-carrageenan reference (Table 1, entry 1). This shift is attributed to the electrostatic interaction between the negatively charged sulfate ions in the carrageenan chains and the positively charged cations from the iodide salt.4 The effect of the addition of iodide salts generally increased the electrical conductivity (Table 1, entry 2) of the films formed, as compared with J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 825 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Effect of -carrageenan concentration on ionic conductivity of carrageenan/TBAI/I2 systems Figure 1: Optical images of 1 % w/v -carrageenan films mixed with 40 % w/w TBAI (w.r.t. -carrageenan) in 4:1 mole ratio TBAI /I2 electrolyte in varying solvents a) in 100 % water, b) in 3:1 H2O/acetonitrile, c) in 1:1 H2O/acetonitrile, d) polymer gel electrolytes containing -carrageenan and TBAI/LiI/I2 electrolyte: left: 1 % w/v -carrageenan in DMSO (liquid state), right: 2 % w/v -carrageenan in DMSO (gel-state) only to pure aqueous systems. A significant amount of water present in the electrolyte systems applied to DSSCs causes a decrease in the photocurrent of the device associated with dye desorption, iodate formation and a decrease in the electron lifetime.14 Hence, the amount of water must be controlled, if not eliminated. It is the objective of this paper to investigate the influence of the solvent system on carrageenan electrolyte and its performance in a DSSC. This paper reports, for the first time, on a novel carrageenan-based electrolyte system with an iodide/tri-iodide redox couple prepared using a mixed solvent system and a non-aqueous solvent. 2 MATERIALS AND METHODS 2.1 Optimization of the -carrageenan electrolyte sys- tem Different -carrageenan (Shemberg, Philippines) biopolymer gel electrolytes were optimized by varying the ratio of water/acetonitrile and the amount of polymer with a constant amount of tetrabutylammonium iodide (TBAI, Sigma Aldrich, a reagent grade of 98 %; 20 % w/w of TBAI with respect to -carrageenan) and iodine (I2, Aldrich, = 99.99 % metal basis). The molar ratio of TBAI/I2 was 4:1. The ratio of water/ACN was optimized by dissolving -carrageenan (0.2 g) with TBAI/I2 in different solvent systems consisting of 100 % water, 3:1 water/ACN, 1:1 water/ACN, 1:3 water/ACN, and 100 % ACN at 80 °C until a homogeneous solution was formed. Thin films were formed by casting solutions on plastic petri dishes and slowly drying in air. Different biopoly- mer electrolytes were prepared using the same procedure by varying the concentration of -carrageenan (0, 0.5, 1.0, 1.5 and 2 % w/v). Using the optimized parameters, -carrageenan gel electrolytes containing 20 % w/w (with respect to the amount of -carrageenan) of other salts such as potassium iodide (Sigma Aldrich) and trimethylsulfonium iodide (TMSI, Sigma Aldrich) were also prepared. A liquid-state polymer electrolyte and a gel-state polymer electrolyte based on -carrageenan were pre- pared by dissolving 1 % w/v or 2 % w/v of the poly- saccharide in 2 mL of dimethylsulfoxide at 70 °C. Tetrabutylammonium iodide (0.5 M), I2 (0.05 M) and LiI (0.1 M) were added to the -carrageenan solution. As the control, the same composition was used without adding the biopolymer. 2.2 Fabrication of dye-sensitized solar cells (DSSCs) The photoanode was prepared spreading a TiO2 paste (Solaronix, Ti-Nanoxide T/SP) on a fluorine-doped tin-oxide (FTO) conducting glass (TCO22-7) with the doctor-blading method. The active area of the photo- anode was fixed to 1 cm × 1 cm with adhesive tapes. The deposited paste was sintered at 450 °C for 30 min and cooled down slowly to room temperature. The deposited TiO2 film was soaked in a dye solution containing a ruthenium dye (Ruthenizer 535-bisTBA, Solaronix) in ethanol for 24 h. The counter electrode was prepared by sintering an FTO conducting glass coated with a pla- tinum solution (Platisol 41121, Solaronix) at 450 °C for 30 mins. The photoanode and counter electrode were sealed with a hot-melt thermoplastic material called Surlyn (DuPont, thickness of 60 μm), also acting as a spacer. A dye-sensitized solar cell was fabricated by inserting a hot gel-electrolyte solution in a pre-drilled hole on the counter electrode. Three different solar cells were fabricated containing different iodide salts (KI, TMSI, and TBAI) in a water/ACN -carrageenan matrix. A control DSSC was prepared using an acetonitrile- based liquid electrolyte (Iodolyte AN50, Solaronix). The same process was used for the fabrication of a DSSC incorporating gel electrolytes in DMSO. 2.3 Characterization The surface morphology and thickness of the thin films were characterized using scanning electron micro- scopy (JEOL JSM-5310). The electrical conductivity was measured using the 4-point probe Van der Pauw method. In brief, four wires were attached to the corners of a 1 cm × 1 cm thin film using a silver paste. The voltage at a constant current was measured. The resistivity and electrical conductivity were calculated using the following equations: Resistivity, s d V l V l f R R = + ⎛ ⎝ ⎜ ⎞ ⎠ ⎟ ⎛ ⎝ ⎜ ⎞ ⎠ ⎟ π ln 2 1 2 2 2 1 2 and Electrical conductivity,  = 1/s, where d, V, I and f R /R( )1 2 refer to the thickness, potential, current and van der Pauw function, respectively. The redox properties of the -carrageenan gel electrolyte were characterized using 3-electrode cyclic voltammetry (Powerlab/4SP potentiostat). The electrodes composed of the Ag/AgCl reference electrode, plati- num-wire counter electrode and glassy-carbon working electrode were pierced through the gel. Electrochemical impedance spectroscopy (Metrohm Autolab potentiostat PGSTAT128N) was obtained by sandwiching the gel electrolyte in between two platinum-coated FTO conducting-glass electrodes. The active area was fixed to 1 × 1 cm2 and the thickness of the spacer was about 50 μm. The frequency was set from 106 Hz to 10–1 Hz with an AC applied voltage of 10 mV. The ionic conductivity () of the gel electrolyte was calculated from the solution resistance (Rs) using Equation (3):  = L ARs (3) The solution resistance (Rs) is determined as the intercept of the Nyquist plot against the real part of the impedance or the horizontal axis. The variables L and A correspond to the thickness of the spacer and the active area, respectively. The solar-cell parameters such as the J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... 824 Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS open-circuit voltage, short-circuit current, fill factor and efficiency of the fabricated dye-sensitized solar cells with -carrageenan gel electrolyte were assessed using a 100 mW/cm2 AM 1.5 solar simulator (Abet Technolo- gies) coupled with a potentiostat to obtain a current- voltage (IV) curve. 3 RESULTS AND DISCUSSIONS 3.1 Carrageenan-electrolyte system in a mixed solvent system Acetonitrile (ACN), which is a common organic solvent used for preparing liquid electrolytes for DSSC applications, does not dissolve -carrageenan. Thus, to form -carrageenan electrolytes containing a redox couple I–/I3– (from TBAI and I2), the amount of water is crucial to dissolve the hydrophilic polysaccharide.15 Using water only as the solvent, a black film was formed with observable colorless crystals (Figure 1a) indicating the unsuccessful dissolution of both TBAI salt and iodine. Likewise, crystallization of the salt was formed on the surface of the film prepared using a 3:1 v/v water/ACN system (Figure 1b). The salt crystallization is attributed to the relative insolubility of the salt in a water-rich solvent. Moreover, the presence of the salt and iodine crystals on the film surface indicates that it fails to dissociate in the polymer matrix, leading to an un- successful incorporation of the charge-carrier redox couple I–/I3–. For a 1:1 v/v water/ACN mixed-solvent system, a uniform orange polymer-electrolyte film with- out visible crystals was formed (Figure 1c), indicating good dissolution and incorporation of the I–/I3– redox couple in the polymer matrix. No films were cast for the 1:3 v/v water/acetonitrile solvent systems and in 100 % acetonitrile because of the relative insolubility of carrageenan in an acetonitrile-rich solvent. The concentration of carrageenan in the polymer electrolyte containing 20 % w/w TBAI (w.r.t. -carra- geenan) in the 1:1 H2O/acetonitrile mixed solvent affects the property and the ionic conductivity of the electrolyte. A liquid polymer electrolyte is formed when the con- centrations of carrageenan are in a range of 0.1–0.5 % w/v. Increasing the carrageenan concentration to  1.0 % w/v gave a gel polymer electrolyte. Consequently, the ionic conductivity of the polymer electrolytes changes in response to its composite form. The ionic conductivity (Figure 2) of -carrageenan-electrolyte systems in- creases as the carrageenan concentration increases from 0.1 to 1.0 % w/v. The increase in the ionic conductivity at low -carrageenan concentrations is attributed to high free-ion concentrations due to their interaction with the functional group present in -carrageenan.16. The ionic conductivity decreases at carrageenan concentrations of 1.0–2.0 % w/v. The increase in the polymer concentra- tion causes chain entanglements, impeding the move- ment of the ions, thus, decreasing its ionic conductivity.16 Different iodide salts were incorporated in -carrageenan to prepare the gel electrolytes. Films were formed and the IR spectra of the electrolyte films con- tained typical functional-group vibrations of -carra- geenan.17 The bands at 1272 cm–1 were identified as an O=S=O symmetric vibration that would respond in the presence of a cation. A shift in the frequency was observed in the O=S=O symmetric vibration compared with the pure k-carrageenan reference (Table 1, entry 1). This shift is attributed to the electrostatic interaction between the negatively charged sulfate ions in the carrageenan chains and the positively charged cations from the iodide salt.4 The effect of the addition of iodide salts generally increased the electrical conductivity (Table 1, entry 2) of the films formed, as compared with J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 825 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Effect of -carrageenan concentration on ionic conductivity of carrageenan/TBAI/I2 systems Figure 1: Optical images of 1 % w/v -carrageenan films mixed with 40 % w/w TBAI (w.r.t. -carrageenan) in 4:1 mole ratio TBAI /I2 electrolyte in varying solvents a) in 100 % water, b) in 3:1 H2O/acetonitrile, c) in 1:1 H2O/acetonitrile, d) polymer gel electrolytes containing -carrageenan and TBAI/LiI/I2 electrolyte: left: 1 % w/v -carrageenan in DMSO (liquid state), right: 2 % w/v -carrageenan in DMSO (gel-state) the reference pure -carrageenan due to the addition and incorporation of the redox species. The ionic conduc- tivities (Table 1, entry 3) of the gel electrolytes are generally higher than that of the acetonitrile-based liquid electrolyte, indicating a facile transport of the redox couple in the carrageenan matrix. Cyclic voltammetry of the gel electrolytes with and without -carrageenan (Figure 3) showed that the addition of the polymer does not impede the electrochemical behavior of the I–/I3– redox couple in the mixed water/acetonitrile system. Two redox processes were observed, indicating the migration of I–/I3– in the 3D matrix of -carrageenan.18 Solar cells were fabricated using the conventional set-up composed of a TiO2/ruthenium dye as the photo- anode and a platinum counter electrode sandwiching the -carrageenan-gel electrolyte systems in water/ACN. I-V characteristics of the DSSCs (Table 2 and Figure 4a) containing the -carrageenan-gel electrolytes showed poor efficiency (<0.10 %) as compared to the liquid electrolyte (2.33 %), attributed to a much lower short- circuit current, which can be associated with the pre- sence of a significant amount of water in the electrolyte system. The presence of water causes dye detachment due to the weakening of the dye-TiO2 linkage, resulting in a limited amount of electrons injected from the excited state of the dye and consequently leading to the lowering of the short-circuit current.16 The obtained efficiencies of different -carrageenan-gel electrolytes with a mixture of acetonitrile and water suggest that imparting a large amount of water (50 %) in the electrolyte system for J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... 826 Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 4: I-V curves of the DSSCs: a) containing different iodide salts (KI, TMSI and TBAI) in the water/ACN -carrageenan matrix and b) containing -carrageenan-based electrolytes in DMSO under a 100 mW/cm2 illumination Table 2: Characteristics of the fabricated DSSCs under a 100 mW/cm2 illumination Electrolyte Voc (V) Isc(mA/cm2) FF  ( %) Iodolyte (control) 0.695 6.444 0.521 2.333 KI:I2 0.746 0.100 0.819 0.061 TMSI:I2 0.776 0.0414 0.152 0.005 TBAI:I2 0.778 0.0817 0.765 0.049 Figure 3: Electrochemical-behavior (cyclic voltametry) analysis of the -carrageenan-gel electrolytes containing different salts; control: LiI/I2 in 1:1 v/v water/ACN without carrageenan Table 1: FTIR and conductivity of the -carrageenan electrolytes containing different salts Entry Pure - carrageenan(reference) Iodolyte AN50 (control) KI/I2 TMSI/I2 TBAI/I2 in -carrageenan 1 FTIR O=S=O vibration (cm–1) 1272 – 1261 1274 1263 2 Electrical conductivity of polymer-electrolyte filmsa (S cm–1) 0.127 – 0.281 0.189 0.326 3 Ionic conductivity of polymer-electrolyte gelsb (x10–4 S cm–1) – 1.66 2.52 3.05 2.17 aMeasured with 4-point probe Van der Pauw technique bMeasured with electrochemical impedance spectroscopy (EIS) DSSCs is still detrimental to the performance of a solar- cell device. 3.2 Carrageenan-electrolyte system in a non-aqueous solvent Carrageenan is not known to dissolve in purely non- aqueous solvents. However, in the course of our investigation of electrolyte systems, we observed that it dissolves in dimethylsulfoxide (DMSO).19,20 We investi- gated the solubility properties of -carrageenan in DMSO, dissolving different amounts of the biopolymer at 60 °C. A liquid solution is formed with 1 % w/v -carrageenan (a solubility of 0.01 g/mL DMSO at room temperature and at 60 °C). Increasing the -carrageenan concentration from 1.5 to 3.5 % w/v allows the dissolu- tion of the biopolymer in DMSO, forming a transparent solution at 60 °C. However, upon cooling down at room temperature, it forms into a stable gel (Figure 1d). A further increase in the concentration of -carrageenan of 4–5 % w/v leads to the formation of a stable gel even at 60 °C. Other organic solvents such as ethanol, methanol, acetonitrile, tetrahydrofuran, dimethylformamide and methoxyethanol did not show good solubility. The advantages of using DMSO are its relatively low vapor pressure and relatively non-volatile nature, which are useful for maintaining its gel state. Moreover, this orga- nic solvent shows a relatively low toxicity and is con- sidered environmentally benign.6 The addition of electro- lyte systems TBAI and I2 turned the -carrageenan DMSO solution into a dark-brown one indicating the incorporation of the electrolytes. The IR spectrum of the polymer-electrolyte system showed functional groups typical of -carrageenan. A shift in the O=S=O symme- tric vibration from typical 1272 for pure -carrageenan to 1226 cm–1 for carrageenan gel electrolyte is attributed J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 827 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: EIS spectrum of a DSSC containing -carrageenan-based electrolytes in DMSO; inset: equivalent circuit model (R1 = series resistance, R2 = TiO2/dye/electrolyte interface resistance, CPE = constant phase element) Figure 5: a) EIS spectra of electrolytes measured using a symme- trical-cell set-up between two platinized FTO conducting glasses in a frequency range of 1 MHz to 0.1 Hz, b) linear sweep voltammogram of the electrolytes measured between two platinized FTO conducting glasses swept between –1 V to 1 V. Scan rate: 10 mV/s Table 3: I-V electrochemical properties of the carrageenan electrolyte in DMSO measured using two platinized FTO conducting glasses and I-V characteristics of DSSCs under a 100 mW/cm2 illumination* Electrochemical properties of electrolyte I-V characteristics of DSSC Electrolytes in DMSO Rb ** (Ù) RPt (Ù) Ionic con- ductivity, ×10–4 (S/cm) Diffusivity of I3–, ×10–9 (cm2 s–1) Short-circuit current (mA/cm2) Open-circuit voltage (V) Fill factor Efficiency (%) Control (without -carrageenan) 21.28 218.13 2.82 1.48 0.690 0.781 0.410 0.221 1% w/v -carrageenan 25.33 296.41 2.37 1.41 0.960 0.763 0.483 0.354 2 % w/v -carrageenan 25.25 338.46 2.38 1.11 0.983 0.751 0.488 0.360 *The electrolyte used was composed of tetrabutylammonium iodide (0.5 M), I2, (0.05 M) and LiI (0.1 M) **Rb corresponds to the solution resistance, which corresponds to the intercept at the real part of the EIS spectrum the reference pure -carrageenan due to the addition and incorporation of the redox species. The ionic conduc- tivities (Table 1, entry 3) of the gel electrolytes are generally higher than that of the acetonitrile-based liquid electrolyte, indicating a facile transport of the redox couple in the carrageenan matrix. Cyclic voltammetry of the gel electrolytes with and without -carrageenan (Figure 3) showed that the addition of the polymer does not impede the electrochemical behavior of the I–/I3– redox couple in the mixed water/acetonitrile system. Two redox processes were observed, indicating the migration of I–/I3– in the 3D matrix of -carrageenan.18 Solar cells were fabricated using the conventional set-up composed of a TiO2/ruthenium dye as the photo- anode and a platinum counter electrode sandwiching the -carrageenan-gel electrolyte systems in water/ACN. I-V characteristics of the DSSCs (Table 2 and Figure 4a) containing the -carrageenan-gel electrolytes showed poor efficiency (<0.10 %) as compared to the liquid electrolyte (2.33 %), attributed to a much lower short- circuit current, which can be associated with the pre- sence of a significant amount of water in the electrolyte system. The presence of water causes dye detachment due to the weakening of the dye-TiO2 linkage, resulting in a limited amount of electrons injected from the excited state of the dye and consequently leading to the lowering of the short-circuit current.16 The obtained efficiencies of different -carrageenan-gel electrolytes with a mixture of acetonitrile and water suggest that imparting a large amount of water (50 %) in the electrolyte system for J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... 826 Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 4: I-V curves of the DSSCs: a) containing different iodide salts (KI, TMSI and TBAI) in the water/ACN -carrageenan matrix and b) containing -carrageenan-based electrolytes in DMSO under a 100 mW/cm2 illumination Table 2: Characteristics of the fabricated DSSCs under a 100 mW/cm2 illumination Electrolyte Voc (V) Isc(mA/cm2) FF  ( %) Iodolyte (control) 0.695 6.444 0.521 2.333 KI:I2 0.746 0.100 0.819 0.061 TMSI:I2 0.776 0.0414 0.152 0.005 TBAI:I2 0.778 0.0817 0.765 0.049 Figure 3: Electrochemical-behavior (cyclic voltametry) analysis of the -carrageenan-gel electrolytes containing different salts; control: LiI/I2 in 1:1 v/v water/ACN without carrageenan Table 1: FTIR and conductivity of the -carrageenan electrolytes containing different salts Entry Pure - carrageenan(reference) Iodolyte AN50 (control) KI/I2 TMSI/I2 TBAI/I2 in -carrageenan 1 FTIR O=S=O vibration (cm–1) 1272 – 1261 1274 1263 2 Electrical conductivity of polymer-electrolyte filmsa (S cm–1) 0.127 – 0.281 0.189 0.326 3 Ionic conductivity of polymer-electrolyte gelsb (x10–4 S cm–1) – 1.66 2.52 3.05 2.17 aMeasured with 4-point probe Van der Pauw technique bMeasured with electrochemical impedance spectroscopy (EIS) DSSCs is still detrimental to the performance of a solar- cell device. 3.2 Carrageenan-electrolyte system in a non-aqueous solvent Carrageenan is not known to dissolve in purely non- aqueous solvents. However, in the course of our investigation of electrolyte systems, we observed that it dissolves in dimethylsulfoxide (DMSO).19,20 We investi- gated the solubility properties of -carrageenan in DMSO, dissolving different amounts of the biopolymer at 60 °C. A liquid solution is formed with 1 % w/v -carrageenan (a solubility of 0.01 g/mL DMSO at room temperature and at 60 °C). Increasing the -carrageenan concentration from 1.5 to 3.5 % w/v allows the dissolu- tion of the biopolymer in DMSO, forming a transparent solution at 60 °C. However, upon cooling down at room temperature, it forms into a stable gel (Figure 1d). A further increase in the concentration of -carrageenan of 4–5 % w/v leads to the formation of a stable gel even at 60 °C. Other organic solvents such as ethanol, methanol, acetonitrile, tetrahydrofuran, dimethylformamide and methoxyethanol did not show good solubility. The advantages of using DMSO are its relatively low vapor pressure and relatively non-volatile nature, which are useful for maintaining its gel state. Moreover, this orga- nic solvent shows a relatively low toxicity and is con- sidered environmentally benign.6 The addition of electro- lyte systems TBAI and I2 turned the -carrageenan DMSO solution into a dark-brown one indicating the incorporation of the electrolytes. The IR spectrum of the polymer-electrolyte system showed functional groups typical of -carrageenan. A shift in the O=S=O symme- tric vibration from typical 1272 for pure -carrageenan to 1226 cm–1 for carrageenan gel electrolyte is attributed J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 827 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: EIS spectrum of a DSSC containing -carrageenan-based electrolytes in DMSO; inset: equivalent circuit model (R1 = series resistance, R2 = TiO2/dye/electrolyte interface resistance, CPE = constant phase element) Figure 5: a) EIS spectra of electrolytes measured using a symme- trical-cell set-up between two platinized FTO conducting glasses in a frequency range of 1 MHz to 0.1 Hz, b) linear sweep voltammogram of the electrolytes measured between two platinized FTO conducting glasses swept between –1 V to 1 V. Scan rate: 10 mV/s Table 3: I-V electrochemical properties of the carrageenan electrolyte in DMSO measured using two platinized FTO conducting glasses and I-V characteristics of DSSCs under a 100 mW/cm2 illumination* Electrochemical properties of electrolyte I-V characteristics of DSSC Electrolytes in DMSO Rb ** (Ù) RPt (Ù) Ionic con- ductivity, ×10–4 (S/cm) Diffusivity of I3–, ×10–9 (cm2 s–1) Short-circuit current (mA/cm2) Open-circuit voltage (V) Fill factor Efficiency (%) Control (without -carrageenan) 21.28 218.13 2.82 1.48 0.690 0.781 0.410 0.221 1% w/v -carrageenan 25.33 296.41 2.37 1.41 0.960 0.763 0.483 0.354 2 % w/v -carrageenan 25.25 338.46 2.38 1.11 0.983 0.751 0.488 0.360 *The electrolyte used was composed of tetrabutylammonium iodide (0.5 M), I2, (0.05 M) and LiI (0.1 M) **Rb corresponds to the solution resistance, which corresponds to the intercept at the real part of the EIS spectrum to the interaction of the cation with the sulfate functional groups.5 The ionic conductivity and diffusivity of the tri-iodide ion (Table 3) in DMSO showed that the addition of -carrageenan decreases the ionic conduc- tivity of the electrolyte system in comparison to the DMSO-based liquid-electrolyte system. EIS spectra (Figure 5a) of the electrolytes in DMSO show a broader semi-circle arc upon the addition of -carrageenan, which is associated with an increase in the charge-transfer resistance at the platinum/electrolyte interface (RPt). A higher RPt indicates a less efficient electro-activity at the Pt/electrolyte interface attributed to a slower ionic conductivity.21 The concentration of iodide ions is in excess in comparison with the amount of tri-iodide ions. Hence, the limiting current density of the gel electrolytes is associated with the diffusion coeffi- cient of the tri-iodide ions. The limiting current densities of the electrolytes (Figure 5b) show negative and positive values due to the transport of charge carriers from one electrode to the other through the whole elec- trolyte system.8 A slow diffusion of the tri-iodide ions is consistent with the slow ionic conductivity of the carrageenan-electrolyte system. The slow tri-iodide ion diffusion in the -carrageenan electrolytes resulted in a higher recombination.22 These results were consistent with the decreasing trend in the open-circuit voltage of the electrolytes with -carrageenan. The photovoltaic performance of the -carrageenan-based dye-sensitized solar cells (Figure 4b and Table 3) showed that the addition of -carrageenan to the I–/I3– redox electrolyte system in dimethylsulfoxide decreases the open-circuit voltage and increases the short-circuit current of a solar cell. The improvement in the short-circuit current of the DSSCs with a -carrageenan/DMSO electrolyte system is attributed to the ability of the hydroxyl functional group and negatively charged sulfate ions of -carra- geenan to allow dissociation of the ions in the polymer matrix through the electrostatic interaction with the cations. The enhancement in the short-circuit current of the -carrageenan/DMSO electrolytes is the major contribution to the improvement of the overall efficiency of the dye-sensitized solar cells. Good dissociation of ions in the polymer matrix increases the concentration of I–, its counterion, and I3– leading to an increase in the short-circuit current.23 A typical liquid-electrolyte-based DSSC show an EIS spectrum consisting of three semi-circle arcs corres- ponding to the interfacial resistances occurring during the electrochemical reaction at the platinum/electrolyte interface, during the charge-transfer process at the TiO2/dye/electrolyte interface and Warburg diffusion of I–/I3– in the electrolyte. A frequency-response-analyzer (FRA) impedance analysis of the -carrageenan/DMSO DSSC (Figure 6) showed a unique single semi-circle arc, associated with the charge-recombination process occurring at the TiO2/dye/electrolyte interface (R2). The other parameters such as R1 and CPE corresponded to the series resistance and constant phase element, respec- tively. A decreasing trend in the semi-circle arc radius was observed when the -carrageenan concentration was increased. A quantitative approach of the recombination resistance was done by fitting the Nyquist plot with its corresponding equivalent circuit model. Chi-square values for the fitted equivalent circuit model range from 0.02–0.07, indicating good value fitting. The charge- recombination resistances (R2) of the DSSCs without -carrageenan (control), with 1 % w/v -carrageenan (liquid state) and 2 % w/v -carrageenan (gel state), obtained from the plot are 515.19 , 317.88  and 292.86 , respectively. A higher R2 value signifies lower charge recombination between the photoanode conduc- tion-band electrons and I3– ions at the TiO2/dye/elec- trolyte interface resulting in a higher Voc.24 The higher charge recombination of the polymer electrolytes with -carrageenan, compared to the electrolyte without the polymer, is generally attributed to the slow ionic con- ductivity of the ions in the gel matrix of carrageenan (Table 3). Moreover, the -carrageenan electrolyte pene- trates poorly into the TiO2 layer due to its polymeric nature contributing to the less efficient charge-transfer process.21 The results were consistent with the photo- voltaic characteristics of the fabricated dye-sensitized solar cells. 4 CONCLUSIONS A hydrophilic polysaccharide, -carrageenan, was successfully used as the polymer matrix for a gel-elec- trolyte system in a mixed solvent of water and aceto- nitrile for dye-sensitized solar cells. The solvent system improved the dissolution of the components and the electrolyte properties. Efficiencies of less than 0.1 % of the fabricated DSSCs are attributed to the significant amount of water. A novel gel-polymer electrolyte con- sisting of -carrageenan electrolyte gel in DMSO was developed for DSSCs. Removal of the water from the polymer-electrolyte system improved the DSSC per- formance from 0.049 to 0.36 %, a seven-fold increase in the magnitude. The presence of -carrageenan in the DMSO-based gel-electrolyte system helps improve the solar-cell efficiency from 0.221 to 0.360 %. Acknowledgments The authors are grateful to Dr. Erwin Enriquez, Dr. Arnel Salvador and Dr. Armando Somintac for the use of the solar simulator and Ms. Anna San Esteban for guidance in solar measurements. This work was supported by research grants from DOST-PCIEERD of the Republic of the Philippines; DOST-SEI-ASTHRDP and DLSU-PhD Fellowship Grants to J. P. O Bantang; DLSU Research Faculty Grant, DLSU Science Foundation and DLSU-URCO to D. H. Camacho. J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... 828 Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 5 REFERENCES 1 J. R. MacCallum, C. A. Vincent, Polymer Electrolytes Reviews-l, Elsevier Applied Science, London, 1989 2 K. N. Kumar, T. Sreekanth, M. J. Reddy, U. V. S. Rao, Study of transport and electrochemical cell characteristics of PVP:NaClO3 polymer electrolyte system, J. Power Sources, 101 (2001), 130–133 doi:10.1016/S0378-7753(01)00658-9 3 T. M. W. J. Bandara, P. Ekanayake, M. A. K. L. Dissanayake, I. Albinsson, B. E. Mellander, A polymer electrolyte containing ionic liquid for possible applications in photoelectrochemical solar cells, J. Solid State Electrochem., 14 (2010), 1221–1226, doi:10.1007/ s10008-009-0951-x 4 X. Guo, P. Yi, Y. Yang, J. Cui, S. Xiao, W. Wang, Effects of surfactants on agarose-based magnetic polymer electrolyte for dye-sensitized solar cells, Electrochim. Acta, 90 (2013), 524–529, doi:10.1016/j.electacta.2012.12.028 5 R. Singh, A. R. Polu, B. Bhattacharya, H.-W. Rhee, C. Varlikli, P. K. Singh, Perspectives for solid biopolymer electrolytes in dye sensitized solar cell and battery application, Renew. Sustainable Energy Rev., 65 (2016), 1098–1117, doi:10.1016/j.rser.2016.06.026 6 W. Wang, X. Guo, Y. Yang, Lithium iodide effect on the electrochemical behavior of agarose based polymer electrolyte for dye-sensitized solar cell, Electrochim. Acta, 56 (2011), 7347–7351, doi:10.1016/j.electacta.2011.06.032 7 H.-L. Hsu, W.-T. Hsu, J. Leu, Effects of environmentally benign solvents in the agarose gel electrolytes on dye-sensitized solar cells, Electrochim. Acta, 56 (2011), 5904–5909, doi:10.1016/j.electacta. 2011.04.117 8 P. Li, Y. Zhang, W. Fa, Y. Zhang, B. Huang, Synthesis of a grafted cellulose gel electrolyte in an ionic liquid ([Bmim]I) for dye-sensitized solar cells, Carbohydr. Polym., 86 (2011), 1216–1220, doi:10.1016/j.carbpol.2011.06.017 9 X. Huang, Y. Liu, J. Deng, B. Yi, X. Yu, P. Shen, S. Tan, A novel polymer gel electrolyte based on cyanoethylated cellulose for dye-sensitized solar cells, Electrochim. Acta, 80 (2012), 219–226, doi:10.1016/j.electacta.2012.07.014 10 M. Kaneko, T. Hoshi, Y, Kaburagi, H. Ueno, Solid type dye-sensi- tized solar cell using polysaccharide containing the redox electrolyte solution, J. Electroanal. Chem., 572 (2004), 21–27, doi:10.1016/ j.jelechem.2004.05.021 11 H. Ueno, M. Kaneko, Investigation of a nanostructured poly- saccharide solid medium for electrochemistry, J. Electroanal. Chem., 568 (2004), 87–92, doi:10.1016/j.jelechem.2003.12.044 12 N. N. Mobarak, N. Ramli, A. Ahmad, M. Y. A. Rahman, Chemical interaction and conductivity of carboxymethyl -carrageenan based green polymer electrolyte, Solid State Ionics, 224 (2012), 51–57, doi:10.1016/j.ssi.2012.07.010 13 F. Bella, N. N. Mobarak, F. N. Jumaah, A. Ahmad, From seaweeds to biopolymeric electrolytes for third generation solar cells: An intriguing approach, Electrochim. Acta, 151 (2015), 306–311, doi:10.1016/j.electacta.2014.11.058 14 C. Law, S. C. Pathirana, X. Li, A. Y. Anderson, P. R. F. Barnes, A. Listorti, T. H. Ghaddar, B. C. O’Regan, Water-Based Electrolytes for Dye-Sensitized Solar Cells, Adv. Mater., 22 (2010), 4505–4509, doi:10.1002/adma.201001703 15 D. H. Camacho, S. J. M. Tambio, M. I. A. Oliveros, Carrageenan- Ionic Liquid Composite: Development of Polysaccharide-Based Solid Electrolyte System, The Manila J. of Science, 6 (2011), 8–15 16 H.-L. Lu, Y.-H. Lee, S.-T. Huang, C. Su, T. C.-K. Yang, Influences of water in bis-benzimidazole-derivative electrolyte additives to the degradation of the dye-sensitized solar cells, Sol. Energy Mater. Sol. Cells, 95 (2011), 158–162, doi:10.1016/j.solmat.2010.02.018 17 C. Tranquilan-Aranilla, N. Nagasawa, A. Bayquen, A. Dela Rosa, Synthesis and characterization of carboxymethyl derivatives of kappa-carrageenan, Carbohydr. Polym., 87 (2012), 1810–1816, doi:10.1016/j.carbpol.2011.10.009 18 S. Yuan, Q. Tang, B. He, P. Yang, Efficient quasi-solid-state dye-sen- sitized solar cells employing polyaniline and polypyrrole incorporated microporous conducting gel electrolytes, J. Power Sources, 254 (2014), 98–105, doi:10.1016/j.jpowsour.2013.12.112 19 S. Chan, J. P. Bantang, D. Camacho, Influence of Nanomaterial Fillers in Biopolymer Electrolyte System for Squaraine-Based Dye-Sensitized Solar Cells, Int. J. Electrochem. Sci., 10 (2015), 7696–7706 20 J. P. Bantang, D. Camacho, A Novel Biopolymer Gel Electrolyte System for DSSC Applications, Proceedings of the DLSU Research Congress, 3, 2015 21 S. Venkatesan, N. Obadja, T.-W. Chang, L.-T. Chen, Y.-L. Lee, Performance improvement of gel- and solid-state dye-sensitized solar cells by utilization of the blending effect of poly (vinylidene fluoride-cohexafluropropylene) and poly (acrylonitrile-co-vinyl acetate) co-polymers, J. Power Sources, 268 (2014), 77–81, doi:10.1016/j.jpowsour.2014.06.016 22 H.-L. Hsu, C.-F. Tien, Y.-T. Yang, J. Leu, Dye-sensitized solar cells based on agarose gel electrolytes using allylimidazolium iodides and environmentally benign solvents, Electrochim. Acta, 91 (2013), 208–213, doi:10.1016/j.electacta.2012.12.133 23 K. Hara, T. Horiguchi, T. Kinoshita, K. Sayama, H. Arakawa, Influence of electrolytes on the photovoltaic performance of organic dye-sensitizednanocrystalline TiO2 solar cells, Sol. Energy Mater. Sol. Cells, 70 (2001), 151–161, doi:10.1016/S0927-0248(01)00019-8 24 Y. Yang, J. Cui, P. Yi, X. Zheng, X. Guo, W. Wang, Effects of nanoparticle additives on the properties of agarose polymer electrolytes, J. Power Sources, 248 (2014), 988–993, doi:10.1016/ j.jpowsour.2013.10.016 J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 829 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS to the interaction of the cation with the sulfate functional groups.5 The ionic conductivity and diffusivity of the tri-iodide ion (Table 3) in DMSO showed that the addition of -carrageenan decreases the ionic conduc- tivity of the electrolyte system in comparison to the DMSO-based liquid-electrolyte system. EIS spectra (Figure 5a) of the electrolytes in DMSO show a broader semi-circle arc upon the addition of -carrageenan, which is associated with an increase in the charge-transfer resistance at the platinum/electrolyte interface (RPt). A higher RPt indicates a less efficient electro-activity at the Pt/electrolyte interface attributed to a slower ionic conductivity.21 The concentration of iodide ions is in excess in comparison with the amount of tri-iodide ions. Hence, the limiting current density of the gel electrolytes is associated with the diffusion coeffi- cient of the tri-iodide ions. The limiting current densities of the electrolytes (Figure 5b) show negative and positive values due to the transport of charge carriers from one electrode to the other through the whole elec- trolyte system.8 A slow diffusion of the tri-iodide ions is consistent with the slow ionic conductivity of the carrageenan-electrolyte system. The slow tri-iodide ion diffusion in the -carrageenan electrolytes resulted in a higher recombination.22 These results were consistent with the decreasing trend in the open-circuit voltage of the electrolytes with -carrageenan. The photovoltaic performance of the -carrageenan-based dye-sensitized solar cells (Figure 4b and Table 3) showed that the addition of -carrageenan to the I–/I3– redox electrolyte system in dimethylsulfoxide decreases the open-circuit voltage and increases the short-circuit current of a solar cell. The improvement in the short-circuit current of the DSSCs with a -carrageenan/DMSO electrolyte system is attributed to the ability of the hydroxyl functional group and negatively charged sulfate ions of -carra- geenan to allow dissociation of the ions in the polymer matrix through the electrostatic interaction with the cations. The enhancement in the short-circuit current of the -carrageenan/DMSO electrolytes is the major contribution to the improvement of the overall efficiency of the dye-sensitized solar cells. Good dissociation of ions in the polymer matrix increases the concentration of I–, its counterion, and I3– leading to an increase in the short-circuit current.23 A typical liquid-electrolyte-based DSSC show an EIS spectrum consisting of three semi-circle arcs corres- ponding to the interfacial resistances occurring during the electrochemical reaction at the platinum/electrolyte interface, during the charge-transfer process at the TiO2/dye/electrolyte interface and Warburg diffusion of I–/I3– in the electrolyte. A frequency-response-analyzer (FRA) impedance analysis of the -carrageenan/DMSO DSSC (Figure 6) showed a unique single semi-circle arc, associated with the charge-recombination process occurring at the TiO2/dye/electrolyte interface (R2). The other parameters such as R1 and CPE corresponded to the series resistance and constant phase element, respec- tively. A decreasing trend in the semi-circle arc radius was observed when the -carrageenan concentration was increased. A quantitative approach of the recombination resistance was done by fitting the Nyquist plot with its corresponding equivalent circuit model. Chi-square values for the fitted equivalent circuit model range from 0.02–0.07, indicating good value fitting. The charge- recombination resistances (R2) of the DSSCs without -carrageenan (control), with 1 % w/v -carrageenan (liquid state) and 2 % w/v -carrageenan (gel state), obtained from the plot are 515.19 , 317.88  and 292.86 , respectively. A higher R2 value signifies lower charge recombination between the photoanode conduc- tion-band electrons and I3– ions at the TiO2/dye/elec- trolyte interface resulting in a higher Voc.24 The higher charge recombination of the polymer electrolytes with -carrageenan, compared to the electrolyte without the polymer, is generally attributed to the slow ionic con- ductivity of the ions in the gel matrix of carrageenan (Table 3). Moreover, the -carrageenan electrolyte pene- trates poorly into the TiO2 layer due to its polymeric nature contributing to the less efficient charge-transfer process.21 The results were consistent with the photo- voltaic characteristics of the fabricated dye-sensitized solar cells. 4 CONCLUSIONS A hydrophilic polysaccharide, -carrageenan, was successfully used as the polymer matrix for a gel-elec- trolyte system in a mixed solvent of water and aceto- nitrile for dye-sensitized solar cells. The solvent system improved the dissolution of the components and the electrolyte properties. Efficiencies of less than 0.1 % of the fabricated DSSCs are attributed to the significant amount of water. A novel gel-polymer electrolyte con- sisting of -carrageenan electrolyte gel in DMSO was developed for DSSCs. Removal of the water from the polymer-electrolyte system improved the DSSC per- formance from 0.049 to 0.36 %, a seven-fold increase in the magnitude. The presence of -carrageenan in the DMSO-based gel-electrolyte system helps improve the solar-cell efficiency from 0.221 to 0.360 %. Acknowledgments The authors are grateful to Dr. Erwin Enriquez, Dr. Arnel Salvador and Dr. Armando Somintac for the use of the solar simulator and Ms. Anna San Esteban for guidance in solar measurements. This work was supported by research grants from DOST-PCIEERD of the Republic of the Philippines; DOST-SEI-ASTHRDP and DLSU-PhD Fellowship Grants to J. P. O Bantang; DLSU Research Faculty Grant, DLSU Science Foundation and DLSU-URCO to D. H. Camacho. J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... 828 Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 5 REFERENCES 1 J. R. MacCallum, C. A. Vincent, Polymer Electrolytes Reviews-l, Elsevier Applied Science, London, 1989 2 K. N. Kumar, T. Sreekanth, M. J. Reddy, U. V. S. Rao, Study of transport and electrochemical cell characteristics of PVP:NaClO3 polymer electrolyte system, J. Power Sources, 101 (2001), 130–133 doi:10.1016/S0378-7753(01)00658-9 3 T. M. W. J. Bandara, P. Ekanayake, M. A. K. L. Dissanayake, I. Albinsson, B. E. Mellander, A polymer electrolyte containing ionic liquid for possible applications in photoelectrochemical solar cells, J. Solid State Electrochem., 14 (2010), 1221–1226, doi:10.1007/ s10008-009-0951-x 4 X. Guo, P. Yi, Y. Yang, J. Cui, S. Xiao, W. Wang, Effects of surfactants on agarose-based magnetic polymer electrolyte for dye-sensitized solar cells, Electrochim. Acta, 90 (2013), 524–529, doi:10.1016/j.electacta.2012.12.028 5 R. Singh, A. R. Polu, B. Bhattacharya, H.-W. Rhee, C. Varlikli, P. K. Singh, Perspectives for solid biopolymer electrolytes in dye sensitized solar cell and battery application, Renew. Sustainable Energy Rev., 65 (2016), 1098–1117, doi:10.1016/j.rser.2016.06.026 6 W. Wang, X. Guo, Y. Yang, Lithium iodide effect on the electrochemical behavior of agarose based polymer electrolyte for dye-sensitized solar cell, Electrochim. Acta, 56 (2011), 7347–7351, doi:10.1016/j.electacta.2011.06.032 7 H.-L. Hsu, W.-T. Hsu, J. Leu, Effects of environmentally benign solvents in the agarose gel electrolytes on dye-sensitized solar cells, Electrochim. Acta, 56 (2011), 5904–5909, doi:10.1016/j.electacta. 2011.04.117 8 P. Li, Y. Zhang, W. Fa, Y. Zhang, B. Huang, Synthesis of a grafted cellulose gel electrolyte in an ionic liquid ([Bmim]I) for dye-sensitized solar cells, Carbohydr. Polym., 86 (2011), 1216–1220, doi:10.1016/j.carbpol.2011.06.017 9 X. Huang, Y. Liu, J. Deng, B. Yi, X. Yu, P. Shen, S. Tan, A novel polymer gel electrolyte based on cyanoethylated cellulose for dye-sensitized solar cells, Electrochim. Acta, 80 (2012), 219–226, doi:10.1016/j.electacta.2012.07.014 10 M. Kaneko, T. Hoshi, Y, Kaburagi, H. Ueno, Solid type dye-sensi- tized solar cell using polysaccharide containing the redox electrolyte solution, J. Electroanal. Chem., 572 (2004), 21–27, doi:10.1016/ j.jelechem.2004.05.021 11 H. Ueno, M. Kaneko, Investigation of a nanostructured poly- saccharide solid medium for electrochemistry, J. Electroanal. Chem., 568 (2004), 87–92, doi:10.1016/j.jelechem.2003.12.044 12 N. N. Mobarak, N. Ramli, A. Ahmad, M. Y. A. Rahman, Chemical interaction and conductivity of carboxymethyl -carrageenan based green polymer electrolyte, Solid State Ionics, 224 (2012), 51–57, doi:10.1016/j.ssi.2012.07.010 13 F. Bella, N. N. Mobarak, F. N. Jumaah, A. Ahmad, From seaweeds to biopolymeric electrolytes for third generation solar cells: An intriguing approach, Electrochim. Acta, 151 (2015), 306–311, doi:10.1016/j.electacta.2014.11.058 14 C. Law, S. C. Pathirana, X. Li, A. Y. Anderson, P. R. F. Barnes, A. Listorti, T. H. Ghaddar, B. C. O’Regan, Water-Based Electrolytes for Dye-Sensitized Solar Cells, Adv. Mater., 22 (2010), 4505–4509, doi:10.1002/adma.201001703 15 D. H. Camacho, S. J. M. Tambio, M. I. A. Oliveros, Carrageenan- Ionic Liquid Composite: Development of Polysaccharide-Based Solid Electrolyte System, The Manila J. of Science, 6 (2011), 8–15 16 H.-L. Lu, Y.-H. Lee, S.-T. Huang, C. Su, T. C.-K. Yang, Influences of water in bis-benzimidazole-derivative electrolyte additives to the degradation of the dye-sensitized solar cells, Sol. Energy Mater. Sol. Cells, 95 (2011), 158–162, doi:10.1016/j.solmat.2010.02.018 17 C. Tranquilan-Aranilla, N. Nagasawa, A. Bayquen, A. Dela Rosa, Synthesis and characterization of carboxymethyl derivatives of kappa-carrageenan, Carbohydr. Polym., 87 (2012), 1810–1816, doi:10.1016/j.carbpol.2011.10.009 18 S. Yuan, Q. Tang, B. He, P. Yang, Efficient quasi-solid-state dye-sen- sitized solar cells employing polyaniline and polypyrrole incorporated microporous conducting gel electrolytes, J. Power Sources, 254 (2014), 98–105, doi:10.1016/j.jpowsour.2013.12.112 19 S. Chan, J. P. Bantang, D. Camacho, Influence of Nanomaterial Fillers in Biopolymer Electrolyte System for Squaraine-Based Dye-Sensitized Solar Cells, Int. J. Electrochem. Sci., 10 (2015), 7696–7706 20 J. P. Bantang, D. Camacho, A Novel Biopolymer Gel Electrolyte System for DSSC Applications, Proceedings of the DLSU Research Congress, 3, 2015 21 S. Venkatesan, N. Obadja, T.-W. Chang, L.-T. Chen, Y.-L. Lee, Performance improvement of gel- and solid-state dye-sensitized solar cells by utilization of the blending effect of poly (vinylidene fluoride-cohexafluropropylene) and poly (acrylonitrile-co-vinyl acetate) co-polymers, J. Power Sources, 268 (2014), 77–81, doi:10.1016/j.jpowsour.2014.06.016 22 H.-L. Hsu, C.-F. Tien, Y.-T. Yang, J. Leu, Dye-sensitized solar cells based on agarose gel electrolytes using allylimidazolium iodides and environmentally benign solvents, Electrochim. Acta, 91 (2013), 208–213, doi:10.1016/j.electacta.2012.12.133 23 K. Hara, T. Horiguchi, T. Kinoshita, K. Sayama, H. Arakawa, Influence of electrolytes on the photovoltaic performance of organic dye-sensitizednanocrystalline TiO2 solar cells, Sol. Energy Mater. Sol. Cells, 70 (2001), 151–161, doi:10.1016/S0927-0248(01)00019-8 24 Y. Yang, J. Cui, P. Yi, X. Zheng, X. Guo, W. Wang, Effects of nanoparticle additives on the properties of agarose polymer electrolytes, J. Power Sources, 248 (2014), 988–993, doi:10.1016/ j.jpowsour.2013.10.016 J. P. BANTANG, D. CAMACHO: GELLING POLYSACCHARIDE AS THE ELECTROLYTE MATRIX ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 823–829 829 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... 831–836 DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES AND STRESS CORROSION RESISTANCE OF 7000 Al ALLOYS RAZVOJ TOPLOTNE OBDELAVE ZA IZBOLJ[ANJE MEHANSKIH LASTNOSTI IN NAPETOSTNO KOROZIJSKO ODPORNOST 7000 Al ZLITIN Mehdi Shakouri1, Mohammad Esmailian1, Saeed Shabestari2 1Advanced Material and Renewable Energy Department, Iranian Research Organization for Science and Technology (IROST), P. O. Box 3353-5111, Tehran, Iran 2Center of Excellence for High Strength Alloys Technology (CEHSAT), School of Metallurgy and Materials Engineering, Iran University of Science and Technology (IUST), Narmak, 16846, Tehran, Iran mshakoory@yahoo.com Prejem rokopisa – received: 2016-10-10; sprejem za objavo – accepted for publication: 2017-04-19 doi:10.17222/mit.2016.297 A retrogression and re-ageing (RRA) treatment is a three-step heat treatment that can improve both the mechanical strength and the corrosion resistance in aluminum alloys. In this work, the mechanical and stress corrosion properties under various ageing treatment conditions were investigated in an Al–8.5Zn–2.1Mg–2Cu–0.2Ag (w/%) alloy. The treatments were the T6 conventional method followed by a retrogression and re-ageing (RRA) treatment. Tensile test, scanning electron microscopy (SEM), energy-dispersive X-ray spectroscopy (EDS) and differential scanning calorimetry (DSC) were used to investigate the mechanical and stress corrosion cracking (SCC) properties. The results showed that both, the strength and the corrosion resistance criteria (SCR), which is defined as the ratio between the remaining strength percent in stressed and un-stressed conditions, improve after the RRA treatment. The tensile strength and SCR criteria for the T6 heat treatment were 547 MPa and 71 % initially and then increased to 612 MPa and 95 % for the RRA treatment, respectively. Moreover, the EDS results showed that in grain-boundary precipitates the Cu concentration is much higher for the RRA in comparison with the T6 treatment, but it is lower for the matrix precipitates in the RRA treatment. Keywords: aluminium alloys, stress corrosion resistance, precipitation, retrogression and re-ageing Obdelava z retrogresijo in ponovno o`ivitvijo (angl. RRA) je tristopenjska toplotna obdelava, s katero lahko izbolj{amo tako mehansko trdnost kot odpornosti proti koroziji pri aluminijevih zlitinah. V pri~ujo~em delu so bile raziskovane mehanske lastnosti in lastnosti odpornosti proti koroziji pri zlitini Al–8.5Zn–2.1Mg–2Cu–0.2Ag (w/%) pod razli~nimi pogoji obdelave staranja. Izvedeni so bili testi po konvencionalni T6 metodi, z upo{tevanjem RRA obdelave. Natezni preiskus, vrsti~na elektronska mikroskopija (SEM), energijsko disperzijska rentgenska spektroskopija (EDS) in DSC-testiranja so bili izvedeni, da bi raziskali mehanske in napetostno-korozijske lomne lastnosti (angl. SCC). Rezultati so pokazali, da se oba kriterija tako trdnost kot odpornost proti koroziji, ki definirata razmerje med procentom zaostale trdnosti pri napetostnih in nenapetostnih pogojih, po RRA-obdelavi izbolj{ata. Klju~ne besede: aluminijeve zlitine, odpornost proti koroziji, oborine, retrogresija in ponovno staranje 1 INTRODUCTION The Al–Zn–Mg–Cu series aluminum alloys are preci- pitation-hardening alloys that are used extensively for light-weight structural applications, in particular in the aircraft industry. They have a combination of good strength and good stress corrosion resistance.1 The most popular heat treatment to gain the peak- aged strength (T6X temper) is a solution treatment, quenching, stretching (for stress relief purposes) followed by artificial aging at 120 °C for 24 h, but this temper is highly susceptible to SCC. In order to reduce this susceptibility, an over-aging treatment (T7X temper) is needed.2 This requirement becomes increasingly de- manding as the solute contents are increased in commer- cial alloys in order to improve the mechanical properties even further. To improve both the mechanical strength and corro- sion resistance, it has been proposed to utilize an RRA treatment. This three-step heat treatment has been shown to offer a stress corrosion resistance as good as that of a T7X heat treatment, while keeping a strength compar- able to that of a T6X temper.3,4 This type of heat treat- ment comprises three steps. First, an ageing step that leads to a T6 state. A second step (called retrogression or reversion) of short duration at high temperature dissolves part of the initially formed precipitates. Third heat treatment step at lower temperature leads to the desired microstructure.5 Stress corrosion cracking occurs under loading in a corrosive environment. Several investigations have re- ported that the SCC mechanism involves anodic dissolu- tion, hydrogen-induced cracking, passive film rupture, hydrogen embrittlement, magnesium segregation to the Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 831 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 621.78:67.017:620.193:669.715 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)831(2017) M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... 831–836 DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES AND STRESS CORROSION RESISTANCE OF 7000 Al ALLOYS RAZVOJ TOPLOTNE OBDELAVE ZA IZBOLJ[ANJE MEHANSKIH LASTNOSTI IN NAPETOSTNO KOROZIJSKO ODPORNOST 7000 Al ZLITIN Mehdi Shakouri1, Mohammad Esmailian1, Saeed Shabestari2 1Advanced Material and Renewable Energy Department, Iranian Research Organization for Science and Technology (IROST), P. O. Box 3353-5111, Tehran, Iran 2Center of Excellence for High Strength Alloys Technology (CEHSAT), School of Metallurgy and Materials Engineering, Iran University of Science and Technology (IUST), Narmak, 16846, Tehran, Iran mshakoory@yahoo.com Prejem rokopisa – received: 2016-10-10; sprejem za objavo – accepted for publication: 2017-04-19 doi:10.17222/mit.2016.297 A retrogression and re-ageing (RRA) treatment is a three-step heat treatment that can improve both the mechanical strength and the corrosion resistance in aluminum alloys. In this work, the mechanical and stress corrosion properties under various ageing treatment conditions were investigated in an Al–8.5Zn–2.1Mg–2Cu–0.2Ag (w/%) alloy. The treatments were the T6 conventional method followed by a retrogression and re-ageing (RRA) treatment. Tensile test, scanning electron microscopy (SEM), energy-dispersive X-ray spectroscopy (EDS) and differential scanning calorimetry (DSC) were used to investigate the mechanical and stress corrosion cracking (SCC) properties. The results showed that both, the strength and the corrosion resistance criteria (SCR), which is defined as the ratio between the remaining strength percent in stressed and un-stressed conditions, improve after the RRA treatment. The tensile strength and SCR criteria for the T6 heat treatment were 547 MPa and 71 % initially and then increased to 612 MPa and 95 % for the RRA treatment, respectively. Moreover, the EDS results showed that in grain-boundary precipitates the Cu concentration is much higher for the RRA in comparison with the T6 treatment, but it is lower for the matrix precipitates in the RRA treatment. Keywords: aluminium alloys, stress corrosion resistance, precipitation, retrogression and re-ageing Obdelava z retrogresijo in ponovno o`ivitvijo (angl. RRA) je tristopenjska toplotna obdelava, s katero lahko izbolj{amo tako mehansko trdnost kot odpornosti proti koroziji pri aluminijevih zlitinah. V pri~ujo~em delu so bile raziskovane mehanske lastnosti in lastnosti odpornosti proti koroziji pri zlitini Al–8.5Zn–2.1Mg–2Cu–0.2Ag (w/%) pod razli~nimi pogoji obdelave staranja. Izvedeni so bili testi po konvencionalni T6 metodi, z upo{tevanjem RRA obdelave. Natezni preiskus, vrsti~na elektronska mikroskopija (SEM), energijsko disperzijska rentgenska spektroskopija (EDS) in DSC-testiranja so bili izvedeni, da bi raziskali mehanske in napetostno-korozijske lomne lastnosti (angl. SCC). Rezultati so pokazali, da se oba kriterija tako trdnost kot odpornost proti koroziji, ki definirata razmerje med procentom zaostale trdnosti pri napetostnih in nenapetostnih pogojih, po RRA-obdelavi izbolj{ata. Klju~ne besede: aluminijeve zlitine, odpornost proti koroziji, oborine, retrogresija in ponovno staranje 1 INTRODUCTION The Al–Zn–Mg–Cu series aluminum alloys are preci- pitation-hardening alloys that are used extensively for light-weight structural applications, in particular in the aircraft industry. They have a combination of good strength and good stress corrosion resistance.1 The most popular heat treatment to gain the peak- aged strength (T6X temper) is a solution treatment, quenching, stretching (for stress relief purposes) followed by artificial aging at 120 °C for 24 h, but this temper is highly susceptible to SCC. In order to reduce this susceptibility, an over-aging treatment (T7X temper) is needed.2 This requirement becomes increasingly de- manding as the solute contents are increased in commer- cial alloys in order to improve the mechanical properties even further. To improve both the mechanical strength and corro- sion resistance, it has been proposed to utilize an RRA treatment. This three-step heat treatment has been shown to offer a stress corrosion resistance as good as that of a T7X heat treatment, while keeping a strength compar- able to that of a T6X temper.3,4 This type of heat treat- ment comprises three steps. First, an ageing step that leads to a T6 state. A second step (called retrogression or reversion) of short duration at high temperature dissolves part of the initially formed precipitates. Third heat treatment step at lower temperature leads to the desired microstructure.5 Stress corrosion cracking occurs under loading in a corrosive environment. Several investigations have re- ported that the SCC mechanism involves anodic dissolu- tion, hydrogen-induced cracking, passive film rupture, hydrogen embrittlement, magnesium segregation to the Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 831 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 621.78:67.017:620.193:669.715 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)831(2017) grain boundaries and a precipitate-free zone (PFZ) along the grain boundary.6–10 However, the microstructural characteristics of Al-Zn-Mg-Cu high strength aluminum alloys are well known to have a strong influence not only on the mechanical properties but also on the SCC sus- ceptibility. Larger grain-boundary precipitates can trap more atomic hydrogen to nucleate hydrogen bubbles, thereby decreasing the hydrogen concentration at grain boundaries below a critical value is considered to prevent intergranular SCC fracture.11 Furthermore, the cathodic grain-boundary precipitates grow by depleting solute atoms. Studies showed that this leads to the broadening of the anodic PFZ, which contained no strengthen pre- cipitation phase and as a result was soft and weak.12 The combination of tensile stress and anodic dissolution caused SCC. SCC resistance is of practical importance for the industrial applications of the Al–Zn–Mg–Cu series aluminum alloys. Various heat treatments in these alloys offer very different SCC properties. The SCR criteria are proposed to compare the SCC resistance. The aim of this paper is to investigate the mechanical strength and SCC of a 7000 series aluminum alloy after conventional T6 and RRA treatments. The SCR criteria have been defined to easily compare the SCC resistance of this alloy after various treatments. SEM investigations were used to determine the effect of the heat treatments on the alloy’s microstructure with the aim of studying its effect on SCC susceptibility. For comparison, the stress corrosion resistance of the alloy with different heat treatments was studied by the breaking load method according to ASTM G139 standard.13 2 EXPERIMENTAL PART The samples used in this study were received as 8-mm-thick sheets that were homogenized, hot rolled and heat treated after being alloyed and casted. The chemical composition of the alloy is shown in Table 1. Table 1: Chemical composition (in mass fractions, w/%) of fabricated alloy Alloy No. Zn Cu Mg Fe Si Zr Ag 1 8.5 2 2.1 0.18 0.16 0.20 0.19 An induction melting furnace used for melting and the melt was poured in a water-cooled copper mold. The as-cast specimens were homogenized at 460 °C for 24 h and hot rolled to about 33 % reduction. Hot rolling was performed with 40 min–1 rolling speed at 430 °C. The specimens were milled to tensile samples according to the ASTM E8M standard. A schematic showing the pre- paration of tensile samples from the specimens is given in Figure 1. A solution heat treatment was done at 471 °C for 6 h followed by a water quenching. T6 ageing was per- formed for 24 h at 120 °C (T6 temper). The scheme of the retrogression and re-ageing treatment is shown in Figure 2. Mechanical properties measurements were made at ambient temperature on the specimens machined according to the ASTM E8M-04 small size standard. The average of three tests was used for each result. The test’s strain rate was 10–3 /s. The SEM analysis was performed on a TESCAN scanning electron microscope. Thermal analysis was per- formed in a DSC 1 Mettler Toledo differential scanning calorimeter. Polished alloy disks with a diameter of 5 mm and 0.6 mm thick were sealed in aluminum pans and heated in a flowing argon atmosphere at a constant heating rate of 10 °C/min. The flowing rate of the argon was 100 mL/min. The stress corrosion tests were performed according to the ASTM G139 standard. The Neutral 3.5 % Sodium Chloride Solution was prepared in accordance with the requirements of the ASTM G44 standard. The ASTM G49 standard was used for preparation of direct tension stress corrosion test specimens. A 207 MPa stress for the cycle of 4 d was applied as per section 8-2 of the ASTM G139 standard. In order to eliminate the corrosion other than SCC (such as pitting), some specimens were tested unstressed. The ratio of remaining strength of stressed and unstressed samples represented in percent, were calculated as SCR criteria for different treatment. 3 RESULTS AND DISCUSSION 3.1 Mechanical strength The mechanical strength of the alloy was 547 MPa after the T6 treatment and was 612 MPa after the RRA treatment. The RRA treatment leads to an about 12 % increase in mechanical strength. While T6 treatment is carried out to ensure maximum strength; the RRA M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... 832 Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Scheme of the RRA treatment Figure 1: Schematic showing the preparation of tensile samples from the specimens treated samples show higher strength. It seems that the performed T6 treatment does not lead to precipitation of all the alloying elements and yet there are elements in the solution. To investigate this further, DSC analysis was conducted on the samples. 3.2 DSC results Figure 3 shows the DSC analysis for the T6 and RRA treatment. Unlike the RRA treatment, the T6 curve shows two exothermic peaks at temperatures between 200 °C and 245 °C. The RRA curve in this region has only one small peak at about 240 °C. These peaks are results of  and  precipitation from solid solution.14–16 In the RRA curve, there exists only one peak related to  precipitation. This means that there are still a lot of ele- ments in solid solution after T6 treatment suggesting that either time or temperature or both are not enough for diffusion and precipitation of the elements. As a result, the maximum strength has not achieved. 3.3 Stress corrosion cracking Stress corrosion tests were performed according to the ASTM G139 standard for direct tension stress corro- sion test method. Figure 4 shows the strength of T6 and RRA samples after exposure to corrosive 3.5 % sodium chloride solution without stressing and with a 207 MPa stress after 4 d. As expected, strength is dropped in both the T6 and RRA treatment after the SCC test. After the SCC test with 207 MPa constant stresses for 4 d in corro- sive solution, the strength of RRA treated samples decreased from 612 MPa to 545 MPa. Accordingly, the strength of the T6 treated samples dropped from 547 MPa to 367 MPa. Calculating the results in percentages, gives us a better understanding of the strength reduction. The test was conducted in corrosive solution in two stress conditions: un-stressed and with constant 207 MPa tension stress. The mechanical strength decreased for both conditions. In order to better demonstrate the SCC results, it is most useful to calculate the remaining strength after SCC test and subtract the effect of corro- sion methods other than SCC (such as pitting). To achieve this, the percentage of SCR ratio of remaining strength of stressed and unstressed samples is defined as stress corrosion resistance SCR. The SCR results of the alloy after the T6 and RRA heat treatments are shown in Figure 5. The remaining strengths of the unstressed samples after RRA and T6 treatments were 94.7 % and 93.3 %, respectively, which are not much different. It means that unstressed corrosion results for both treat- ments were almost identical. However, the remaining strength percentages in the constant stress condition were very different. The remaining strength percentage was 89 % for the RRA treatment and 68 % for the T6 treatment. The SCR for the RRA treatment was 95 % and for the T6 treatment it was 71 %. A significant im- M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 833 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: The SCR results of the alloy after T6 and RRA heat treatments Figure 3: DSC curves for T6 and RRA treatment Figure 4: Direct tension stress corrosion cracking test after T6 and RRA treatment grain boundaries and a precipitate-free zone (PFZ) along the grain boundary.6–10 However, the microstructural characteristics of Al-Zn-Mg-Cu high strength aluminum alloys are well known to have a strong influence not only on the mechanical properties but also on the SCC sus- ceptibility. Larger grain-boundary precipitates can trap more atomic hydrogen to nucleate hydrogen bubbles, thereby decreasing the hydrogen concentration at grain boundaries below a critical value is considered to prevent intergranular SCC fracture.11 Furthermore, the cathodic grain-boundary precipitates grow by depleting solute atoms. Studies showed that this leads to the broadening of the anodic PFZ, which contained no strengthen pre- cipitation phase and as a result was soft and weak.12 The combination of tensile stress and anodic dissolution caused SCC. SCC resistance is of practical importance for the industrial applications of the Al–Zn–Mg–Cu series aluminum alloys. Various heat treatments in these alloys offer very different SCC properties. The SCR criteria are proposed to compare the SCC resistance. The aim of this paper is to investigate the mechanical strength and SCC of a 7000 series aluminum alloy after conventional T6 and RRA treatments. The SCR criteria have been defined to easily compare the SCC resistance of this alloy after various treatments. SEM investigations were used to determine the effect of the heat treatments on the alloy’s microstructure with the aim of studying its effect on SCC susceptibility. For comparison, the stress corrosion resistance of the alloy with different heat treatments was studied by the breaking load method according to ASTM G139 standard.13 2 EXPERIMENTAL PART The samples used in this study were received as 8-mm-thick sheets that were homogenized, hot rolled and heat treated after being alloyed and casted. The chemical composition of the alloy is shown in Table 1. Table 1: Chemical composition (in mass fractions, w/%) of fabricated alloy Alloy No. Zn Cu Mg Fe Si Zr Ag 1 8.5 2 2.1 0.18 0.16 0.20 0.19 An induction melting furnace used for melting and the melt was poured in a water-cooled copper mold. The as-cast specimens were homogenized at 460 °C for 24 h and hot rolled to about 33 % reduction. Hot rolling was performed with 40 min–1 rolling speed at 430 °C. The specimens were milled to tensile samples according to the ASTM E8M standard. A schematic showing the pre- paration of tensile samples from the specimens is given in Figure 1. A solution heat treatment was done at 471 °C for 6 h followed by a water quenching. T6 ageing was per- formed for 24 h at 120 °C (T6 temper). The scheme of the retrogression and re-ageing treatment is shown in Figure 2. Mechanical properties measurements were made at ambient temperature on the specimens machined according to the ASTM E8M-04 small size standard. The average of three tests was used for each result. The test’s strain rate was 10–3 /s. The SEM analysis was performed on a TESCAN scanning electron microscope. Thermal analysis was per- formed in a DSC 1 Mettler Toledo differential scanning calorimeter. Polished alloy disks with a diameter of 5 mm and 0.6 mm thick were sealed in aluminum pans and heated in a flowing argon atmosphere at a constant heating rate of 10 °C/min. The flowing rate of the argon was 100 mL/min. The stress corrosion tests were performed according to the ASTM G139 standard. The Neutral 3.5 % Sodium Chloride Solution was prepared in accordance with the requirements of the ASTM G44 standard. The ASTM G49 standard was used for preparation of direct tension stress corrosion test specimens. A 207 MPa stress for the cycle of 4 d was applied as per section 8-2 of the ASTM G139 standard. In order to eliminate the corrosion other than SCC (such as pitting), some specimens were tested unstressed. The ratio of remaining strength of stressed and unstressed samples represented in percent, were calculated as SCR criteria for different treatment. 3 RESULTS AND DISCUSSION 3.1 Mechanical strength The mechanical strength of the alloy was 547 MPa after the T6 treatment and was 612 MPa after the RRA treatment. The RRA treatment leads to an about 12 % increase in mechanical strength. While T6 treatment is carried out to ensure maximum strength; the RRA M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... 832 Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Scheme of the RRA treatment Figure 1: Schematic showing the preparation of tensile samples from the specimens treated samples show higher strength. It seems that the performed T6 treatment does not lead to precipitation of all the alloying elements and yet there are elements in the solution. To investigate this further, DSC analysis was conducted on the samples. 3.2 DSC results Figure 3 shows the DSC analysis for the T6 and RRA treatment. Unlike the RRA treatment, the T6 curve shows two exothermic peaks at temperatures between 200 °C and 245 °C. The RRA curve in this region has only one small peak at about 240 °C. These peaks are results of  and  precipitation from solid solution.14–16 In the RRA curve, there exists only one peak related to  precipitation. This means that there are still a lot of ele- ments in solid solution after T6 treatment suggesting that either time or temperature or both are not enough for diffusion and precipitation of the elements. As a result, the maximum strength has not achieved. 3.3 Stress corrosion cracking Stress corrosion tests were performed according to the ASTM G139 standard for direct tension stress corro- sion test method. Figure 4 shows the strength of T6 and RRA samples after exposure to corrosive 3.5 % sodium chloride solution without stressing and with a 207 MPa stress after 4 d. As expected, strength is dropped in both the T6 and RRA treatment after the SCC test. After the SCC test with 207 MPa constant stresses for 4 d in corro- sive solution, the strength of RRA treated samples decreased from 612 MPa to 545 MPa. Accordingly, the strength of the T6 treated samples dropped from 547 MPa to 367 MPa. Calculating the results in percentages, gives us a better understanding of the strength reduction. The test was conducted in corrosive solution in two stress conditions: un-stressed and with constant 207 MPa tension stress. The mechanical strength decreased for both conditions. In order to better demonstrate the SCC results, it is most useful to calculate the remaining strength after SCC test and subtract the effect of corro- sion methods other than SCC (such as pitting). To achieve this, the percentage of SCR ratio of remaining strength of stressed and unstressed samples is defined as stress corrosion resistance SCR. The SCR results of the alloy after the T6 and RRA heat treatments are shown in Figure 5. The remaining strengths of the unstressed samples after RRA and T6 treatments were 94.7 % and 93.3 %, respectively, which are not much different. It means that unstressed corrosion results for both treat- ments were almost identical. However, the remaining strength percentages in the constant stress condition were very different. The remaining strength percentage was 89 % for the RRA treatment and 68 % for the T6 treatment. The SCR for the RRA treatment was 95 % and for the T6 treatment it was 71 %. A significant im- M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 833 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: The SCR results of the alloy after T6 and RRA heat treatments Figure 3: DSC curves for T6 and RRA treatment Figure 4: Direct tension stress corrosion cracking test after T6 and RRA treatment provement in the stress corrosion cracking has been achieved after the RRA treatment. 3.4 SEM and EDS analysis The SEM micrographs of the T6 and RRA samples are shown in Figures 6 and 7. Comparison of these two micrographs shows that, after the RRA treatment the grain-boundary precipitates have lost their continuity and divided to relatively larger particles. This is an essential factor for reducing the intergranular corrosion and consequently reducing the stress corrosion cracking. Two basic mechanisms of the SCC in 7000 series aluminum alloys have been pro- posed: anodic dissolution and cathodic dissolution (hydrogen embrittlement).6 The precipitates in the grain boundaries have the electrode potential different from the Al matrix. This would result in the anodic dissolution and form defects in a chloride solution. The hydrogen atoms produced in the crack tip also lead to the hydrogen embrittlement in the grain boundaries. The large size and distance of the grain-boundary particles could decrease the anodic dissolution speed. These particles can also act as the trapping sites for atomic hydrogen and transform M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... 834 Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 9: EDS analysis of RRA sample: a) matrix; b) grain boundary precipitates Figure 7: SEM micrograph of RRA sample showing less continuous precipitates along the grain boundary in comparison to the T6 sample Figure 6: SEM micrograph of T6 sample showing continuous precipitates along the grain boundary Figure 8: EDS analysis of T6 sample: a) matrix, b) grain boundary precipitates them to molecular hydrogen bubbles to reduce the con- centration of the atomic hydrogen at the grain boun- daries. Both the size and distance of the grain-boundary precipitates in RRA treated alloy were larger than that in T6 treated alloy, therefore, the SCR criteria after RRA treatment is higher than the T6 treatment. Another factor showing SCC resistance improvement after RRA treatment is the concentration changes of pre- cipitates and matrix. The Zn and Cu atoms have opposite effects on the electrochemical properties of alumi- num.17,18 In a corrosive medium, the  phase in grain boundary is very active and anodic with respect to the matrix.19 Cu enrichment of precipitates after RRA increases the electrode potential of the grain boundaries and decreases the activity of grain-boundary phases. Figures 9 and 10 show the EDS analysis of the grain- boundary precipitates and matrix after RRA and T6 treatment. Figure 8a shows the Cu-rich matrix after the T6 treatment while according to Figure 8b, the preci- pitates are Zn-reach. After RRA treatment, the matrix depleted from Cu atoms and the grain boundary preci- pitates were Cu-enrichment (Figures 9a to 9b). Large variations in concentration of solute atoms happen during the RRA treatment. Zn atoms have a higher diffu- sivity than Cu atoms and leave the precipitates first after the reversion stage.20 As a result, the Zn content of the matrix increases and a corresponding increase in the Cu precipitate content happens. This happens reversely in the T6 treatment, where Zn incorporates first the preci- pitates, resulting in a high Zn content of the precipitates and high Cu matrix content. The Zn and Cu atoms have opposite effects on the electrochemical properties of aluminum.21 A change in the corrosion behavior of the 7000 series aluminum alloy is related to changes in pre- cipitate composition and particularly to a Cu enrichment of the precipitates.22 During the RRA treatment, the precipitates enrich by Cu atoms (Figure 9b) and this factor as well as other factors improves the SCC resistance. 4 CONCLUSIONS The mechanical strength of an Al–8.5Zn–2.1Mg– 2Cu–0.2Ag (w/%) alloy after conventional T6 treatment was lower than the mechanical strength after RRA treatment. SCR is a useful criteria to compare the SCC in Al-Zn-Mg-Cu alloys after various heat treatments. The SCR of the Al–8.5Zn–2.1Mg–2Cu–0.2Ag (w/%) alloy after conventional T6 treatment was 71 % and was 95 % after RRA treatment. The SCC improved significantly after the RRA treatment. The SCC improvement after the RRA treatment related to Cu enrichment of the precipitates along with the discontinuities of the grain-boundary precipitates. 5 REFERENCES 1 T. Dursun, C. Soutis, Recent developments in advanced aircraft aluminium alloys, Materials & Design, 56 (2014), 862–871, doi:10.1016/j.matdes.2013.12.002 2 D. Liu, B. Xiong, F. Bian, Z. Li, X. Li, Y. Zhang, F. Wang, H. Liu, In situ studies of microstructure evolution and properties of an Al–7.5Zn–1.7Mg–1.4Cu–0.12Zr alloy during retrogression and reaging, Materials & Design, 56 (2014), 1020–1024, doi:10.1016/ j.matdes.2013.12.006 3 J. Li, N. Birbilis, C. Li, Z. Jia, B. Cai, Z. Zheng, Influence of retrogression temperature and time on the mechanical properties and exfoliation corrosion behavior of aluminium alloy AA7150, Materials Characterization, 60 (2009) 11, 1334–1341, doi:10.1016/ j.matchar.2009.06.007 4 R. Bucci, C. Warren, E. Starke, The Need for New Materials in Aging Aircraft Structures, Journal of Aircraft, 37 (2000) 1, 122–129, doi:10.2514/2.2571 5 D. Feng, X. M. Zhang, S. D. Liu, and Y. L. Deng, Non-isothermal retrogression and re-ageing treatment schedule for AA7055 thick plate, Materials & Design, 60 (2014), 208–217, doi:10.1016/ j.matdes.2014.03.064 6 D. Najjar, T. Magnin, T. J. Warner, Influence of critical surface defects and localized competition between anodic dissolution and hydrogen effects during stress corrosion cracking of a 7050 alumi- nium alloy, Materials Science and Engineering: A, 238 (1997) 2, 293–302, doi:10.1016/S0921-5093(97)00369-9 7 T. Burleigh, The postulated mechanisms for stress corrosion cracking of aluminum alloys: a review of the literature 1980-1989, Corrosion, 47 (1991) 2, 89–98, doi:10.5006/1.3585235 8 R. Jones, D. Baer, M. Danielson, J. Vetrano, Role of Mg in the stress corrosion cracking of an Al-Mg alloy, Metallurgical and Materials Transactions A, 32 (2001) 7, 1699–1711, doi:10.1007/s11661- 001-0148-0 9 E. Pugh, Progress toward understanding the stress corrosion problem, Corrosion, 41 (1985) 9, 517–526, doi:10.5006/1.3583022 10 A. Sedriks, J. Green, D. Novak, Corrosion processes and solution chemistry within stress corrosion cracks in aluminum alloys, Proceedings of NACE, 1971, 569–575 11 T. Tsai, T. Chuang, Role of grain size on the stress corrosion crack- ing of 7475 aluminum alloys, Materials Science and Engineering: A, 225 (1997) 1–2, 135–144, doi:10.1016/S0921-5093(96)10840-6 12 T. Pardoen, D. Dumont, A. Deschamps, Y. Brechet, Grain boundary versus transgranular ductile failure, Journal of the Mechanics and Physics of Solids, 51 (2003) 4, 637–665, doi:10.1016/S0022- 5096(02)00102-3 13 ASTM 139:2005(G)- Standard test method for determining stress- corrosion cracking resistance of heat-treatable aluminum alloy products using breaking load method, ASTM International, West Conshohocken 14 K. Ghosh, N. Gao, Determination of kinetic parameters from calorimetric study of solid state reactions in 7150 Al-Zn-Mg alloy, Transactions of Nonferrous Metals Society of China, 21 (2011) 6, 1199–1209, doi:10.1016/S1003-6326(11)60843-1 15 T. Marlaud, A. Deschamps, F. Bley, W. Lefebvre, B. Baroux, Influence of alloy composition and heat treatment on precipitate composition in Al–Zn–Mg–Cu alloys, Acta Materialia, 58 (2010) 1, 248–260, doi:10.1016/j.actamat.2009.09.003 16 J. G. Tang, H. Chen, X. M. Zhang, S. D. Liu, W. J. Liu, H. Ouyang, H. P. Li, Influence of quench-induced precipitation on aging behavior of Al-Zn-Mg-Cu alloy, Transactions of Nonferrous Metals Society of China, 22 (2012) 6, 1255–1263, doi:10.1016/S1003-6326(11) 61313-7 17 R. W. Revie, H. H. Uhlig, Uhlig’s corrosion handbook: vol. 51, John Wiley & Sons, Hoboken 2011 18 G. Silva, B. Rivolta, R. Gerosa, U. Derudi, Study of the SCC Behavior of 7075 Aluminum Alloy After One-Step Aging at 163° C, Journal of Materials Engineering and Performance, 22 (2013) 1, 210–214, doi:10.1007/s11665-012-0221-4 M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 835 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS provement in the stress corrosion cracking has been achieved after the RRA treatment. 3.4 SEM and EDS analysis The SEM micrographs of the T6 and RRA samples are shown in Figures 6 and 7. Comparison of these two micrographs shows that, after the RRA treatment the grain-boundary precipitates have lost their continuity and divided to relatively larger particles. This is an essential factor for reducing the intergranular corrosion and consequently reducing the stress corrosion cracking. Two basic mechanisms of the SCC in 7000 series aluminum alloys have been pro- posed: anodic dissolution and cathodic dissolution (hydrogen embrittlement).6 The precipitates in the grain boundaries have the electrode potential different from the Al matrix. This would result in the anodic dissolution and form defects in a chloride solution. The hydrogen atoms produced in the crack tip also lead to the hydrogen embrittlement in the grain boundaries. The large size and distance of the grain-boundary particles could decrease the anodic dissolution speed. These particles can also act as the trapping sites for atomic hydrogen and transform M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... 834 Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 9: EDS analysis of RRA sample: a) matrix; b) grain boundary precipitates Figure 7: SEM micrograph of RRA sample showing less continuous precipitates along the grain boundary in comparison to the T6 sample Figure 6: SEM micrograph of T6 sample showing continuous precipitates along the grain boundary Figure 8: EDS analysis of T6 sample: a) matrix, b) grain boundary precipitates them to molecular hydrogen bubbles to reduce the con- centration of the atomic hydrogen at the grain boun- daries. Both the size and distance of the grain-boundary precipitates in RRA treated alloy were larger than that in T6 treated alloy, therefore, the SCR criteria after RRA treatment is higher than the T6 treatment. Another factor showing SCC resistance improvement after RRA treatment is the concentration changes of pre- cipitates and matrix. The Zn and Cu atoms have opposite effects on the electrochemical properties of alumi- num.17,18 In a corrosive medium, the  phase in grain boundary is very active and anodic with respect to the matrix.19 Cu enrichment of precipitates after RRA increases the electrode potential of the grain boundaries and decreases the activity of grain-boundary phases. Figures 9 and 10 show the EDS analysis of the grain- boundary precipitates and matrix after RRA and T6 treatment. Figure 8a shows the Cu-rich matrix after the T6 treatment while according to Figure 8b, the preci- pitates are Zn-reach. After RRA treatment, the matrix depleted from Cu atoms and the grain boundary preci- pitates were Cu-enrichment (Figures 9a to 9b). Large variations in concentration of solute atoms happen during the RRA treatment. Zn atoms have a higher diffu- sivity than Cu atoms and leave the precipitates first after the reversion stage.20 As a result, the Zn content of the matrix increases and a corresponding increase in the Cu precipitate content happens. This happens reversely in the T6 treatment, where Zn incorporates first the preci- pitates, resulting in a high Zn content of the precipitates and high Cu matrix content. The Zn and Cu atoms have opposite effects on the electrochemical properties of aluminum.21 A change in the corrosion behavior of the 7000 series aluminum alloy is related to changes in pre- cipitate composition and particularly to a Cu enrichment of the precipitates.22 During the RRA treatment, the precipitates enrich by Cu atoms (Figure 9b) and this factor as well as other factors improves the SCC resistance. 4 CONCLUSIONS The mechanical strength of an Al–8.5Zn–2.1Mg– 2Cu–0.2Ag (w/%) alloy after conventional T6 treatment was lower than the mechanical strength after RRA treatment. SCR is a useful criteria to compare the SCC in Al-Zn-Mg-Cu alloys after various heat treatments. The SCR of the Al–8.5Zn–2.1Mg–2Cu–0.2Ag (w/%) alloy after conventional T6 treatment was 71 % and was 95 % after RRA treatment. The SCC improved significantly after the RRA treatment. The SCC improvement after the RRA treatment related to Cu enrichment of the precipitates along with the discontinuities of the grain-boundary precipitates. 5 REFERENCES 1 T. Dursun, C. Soutis, Recent developments in advanced aircraft aluminium alloys, Materials & Design, 56 (2014), 862–871, doi:10.1016/j.matdes.2013.12.002 2 D. Liu, B. Xiong, F. Bian, Z. Li, X. Li, Y. Zhang, F. Wang, H. Liu, In situ studies of microstructure evolution and properties of an Al–7.5Zn–1.7Mg–1.4Cu–0.12Zr alloy during retrogression and reaging, Materials & Design, 56 (2014), 1020–1024, doi:10.1016/ j.matdes.2013.12.006 3 J. Li, N. Birbilis, C. Li, Z. Jia, B. Cai, Z. Zheng, Influence of retrogression temperature and time on the mechanical properties and exfoliation corrosion behavior of aluminium alloy AA7150, Materials Characterization, 60 (2009) 11, 1334–1341, doi:10.1016/ j.matchar.2009.06.007 4 R. Bucci, C. Warren, E. Starke, The Need for New Materials in Aging Aircraft Structures, Journal of Aircraft, 37 (2000) 1, 122–129, doi:10.2514/2.2571 5 D. Feng, X. M. Zhang, S. D. Liu, and Y. L. Deng, Non-isothermal retrogression and re-ageing treatment schedule for AA7055 thick plate, Materials & Design, 60 (2014), 208–217, doi:10.1016/ j.matdes.2014.03.064 6 D. Najjar, T. Magnin, T. J. Warner, Influence of critical surface defects and localized competition between anodic dissolution and hydrogen effects during stress corrosion cracking of a 7050 alumi- nium alloy, Materials Science and Engineering: A, 238 (1997) 2, 293–302, doi:10.1016/S0921-5093(97)00369-9 7 T. Burleigh, The postulated mechanisms for stress corrosion cracking of aluminum alloys: a review of the literature 1980-1989, Corrosion, 47 (1991) 2, 89–98, doi:10.5006/1.3585235 8 R. Jones, D. Baer, M. Danielson, J. Vetrano, Role of Mg in the stress corrosion cracking of an Al-Mg alloy, Metallurgical and Materials Transactions A, 32 (2001) 7, 1699–1711, doi:10.1007/s11661- 001-0148-0 9 E. Pugh, Progress toward understanding the stress corrosion problem, Corrosion, 41 (1985) 9, 517–526, doi:10.5006/1.3583022 10 A. Sedriks, J. Green, D. Novak, Corrosion processes and solution chemistry within stress corrosion cracks in aluminum alloys, Proceedings of NACE, 1971, 569–575 11 T. Tsai, T. Chuang, Role of grain size on the stress corrosion crack- ing of 7475 aluminum alloys, Materials Science and Engineering: A, 225 (1997) 1–2, 135–144, doi:10.1016/S0921-5093(96)10840-6 12 T. Pardoen, D. Dumont, A. Deschamps, Y. Brechet, Grain boundary versus transgranular ductile failure, Journal of the Mechanics and Physics of Solids, 51 (2003) 4, 637–665, doi:10.1016/S0022- 5096(02)00102-3 13 ASTM 139:2005(G)- Standard test method for determining stress- corrosion cracking resistance of heat-treatable aluminum alloy products using breaking load method, ASTM International, West Conshohocken 14 K. Ghosh, N. Gao, Determination of kinetic parameters from calorimetric study of solid state reactions in 7150 Al-Zn-Mg alloy, Transactions of Nonferrous Metals Society of China, 21 (2011) 6, 1199–1209, doi:10.1016/S1003-6326(11)60843-1 15 T. Marlaud, A. Deschamps, F. Bley, W. Lefebvre, B. Baroux, Influence of alloy composition and heat treatment on precipitate composition in Al–Zn–Mg–Cu alloys, Acta Materialia, 58 (2010) 1, 248–260, doi:10.1016/j.actamat.2009.09.003 16 J. G. Tang, H. Chen, X. M. Zhang, S. D. Liu, W. J. Liu, H. Ouyang, H. P. Li, Influence of quench-induced precipitation on aging behavior of Al-Zn-Mg-Cu alloy, Transactions of Nonferrous Metals Society of China, 22 (2012) 6, 1255–1263, doi:10.1016/S1003-6326(11) 61313-7 17 R. W. Revie, H. H. Uhlig, Uhlig’s corrosion handbook: vol. 51, John Wiley & Sons, Hoboken 2011 18 G. Silva, B. Rivolta, R. Gerosa, U. Derudi, Study of the SCC Behavior of 7075 Aluminum Alloy After One-Step Aging at 163° C, Journal of Materials Engineering and Performance, 22 (2013) 1, 210–214, doi:10.1007/s11665-012-0221-4 M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 835 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 19 S. Maitra, G. English, Mechanism of localized corrosion of 7075 alloy plate, Metallurgical and Materials Transactions A, 12 (1981) 3, 535–541, doi:10.1007/BF02648553 20 H. Bakker, H. Bonzel, C. Bruff, M. Dayananda, W. Gust, J. Horvath, I. Kaur, G. Kidson, A. LeClaire, H. Mehrer, Diffusion in solid metals and alloys/Diffusion in festen metallen und legierungen: vol. 26, Springer, Berlin 1990 21 E. Hollingsworth, H. Hunsicker, Corrosion of aluminum and alumi- num alloys, ASM Handbook, 13 (1987), 583–609 22 J. T. Staley, S. Byrne, E. Colvin, K. Kinnear, Corrosion and stress- corrosion of 7xxx-w products, Materials Science Forum, 217 (1996), 1587–1592, doi:10.4028/MSF.217-222.1587 M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... 836 Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... 837–842 CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS DUBLEX Ni-P/Ni-B-W COATINGS KOROZIJSKA ODPORNOST PLATIRANIH IN NEELEKTRI^NO TOPOLOTNO OBDELANIH DUPLEKS Ni-P/Ni-B-W PREVLEK Behiye Yüksel1, Garip Erdogan2, Fatih Erdem Bastan2, Rasid Ahmed Yýldýz3 1Istanbul Aydin University, Mechanical Engineering Department, 34668 Istanbul, Turkey 2Sakarya University, Engineering Faculty, Department of Metallurgy and Materials Engineering, 54187 Sakarya, Turkey 3TU Istanbul, Mechanical Engineering Faculty, 34437 Istanbul, Turkey behiyeyuksel@aydin.edu.tr Prejem rokopisa – received: 2016-10-20; sprejem za objavo – accepted for publication: 2017-03-10 doi:10.17222/mit.2016.304 In this study, Ni-P/Ni-B-W dublex coatings were deposited on carbon steel substrates (AISI 1020) using the electroless plating process and their microstructure and corrosion properties were systematically evaluated based on different heat-treatment temperatures. Both, the surface morphology and cross-sectional morphology of the Ni-P/Ni-B-W coatings were studied using a scanning electron microscope (SEM), while X-ray diffraction (XRD) was applied for examining the structural modifications. The amorphous coating began to crystallize at a heat-treatment temperature of 350 °C. Potentiodynamic polarization measurements were carried out in an aqueous medium containing 3.5 % NaCl for evaluating the corrosion resistance of as-plated and heat-treated dublex coatings. The corrosion potentials of dublex coatings were observed to shift toward more positive values with increased heat-treatment temperatures. Depending on the heat-treatment temperature, it was identified that the crystallized dublex coatings generally had better corrosion resistance than the amorphous coating. Keywords: Ni-P/Ni-B-W coating, corrosion, heat treatment V {tudiji so bile dvojne Ni-P/Ni-B-W prevleke nane{ene na substrate iz ogljikovega jekla (AISI 1020) s postopkom neelek- tri~nega platiranja. Sistemati~no so bile ocenjene mikrostruktura ter korozijske lastnosti izdelanih prevlek, glede na razli~ne temperature toplotne obdelave. Morfologije povr{in in profili izdelanih Ni-P/Ni-B-W prevlek so bili pregledani z vrsti~nim elektronskim mikroskopom (SEM), medtem ko so bile z rentgensko difrakcijsko analizo (XRD) ugotovljene strukturne spremembe. Amorfna prevleka je za~ela kristalizirati pri temperaturi toplotne obdelave 350 °C. Korozijska odpornost platiranih in toplotno obdelanih dupleks prevlek je bila ocenjena s pomo~jo meritev potenciodinami~ne polarizacije v 3,5 % vodni razto- pini NaCl. Opazovali so korozijske potenciale dupleks prevlek, da bi dobili bolj pozitivne vrednosti z zvi{animi temperaturami topolotne obdelave. Ugotovljeno je bilo, da imajo kristalizirane dupleks prevleke na splo{no bolj{o korozijsko odpornost kot amorfne prevleke. Klju~ne besede: Ni-P/Ni-B-W dvojna prevleka, korozija, toplotna obdelava 1 INTRODUCTION Studied for the first time in 1946 by Brenner and Riddell, the electroless plating process has been used in many industrial applications because of its superior characteristics over electroplating, such as the ability to plate insulation materials and having a homogeneous coating-thickness distribution.1 Among electroless coat- ing types, electroless nickel coating is the most popular for having good hardness, wear and corrosion resistance properties.2 If the electroless Ni coating family is re- viewed, besides pure Ni coatings, Ni-P and Ni-B coat- ings deposited using reducing agents such as hypo- phosphite, borohydride or dimethylamine borane stand out. The major advantages of Ni-P coatings are their low cost, high corrosion resistance and easy process control.3 Nonetheless, a borohydride reduced nickel coating is harder and has higher wear resistance than tool steel and hard chrome coatings.4,5 On the other hand, some studies – although they are limited in number – have highlighted the addition of tungsten to this coating system for in- creasing its corrosion resistance, which is lower than that of a Ni-P coating.6–9 As a result of the co-deposition of this refractory material (which cannot be reduced in an aqueous solution as metallic tungsten) together with iron group metals, developing coatings with attractive corro- sion and tribological properties is possible.10–12 Recently, research for obtaining multilayer coatings to increase existing coating corrosion resistance to higher levels has also been conducted.13,14 The aim of the present study was to evaluate the effect of different heat-treatment temperatures on the microstructure and corrosion properties of electroless Ni-P/Ni-B-W coatings. To the best of our knowledge, no others studies are available on this topic. 2 EXPERIMENTAL PART For this experiment, 10 mm × 10 mm × 60 mm plain carbon steel (AISI 1020) plates were used as a substrate material for developing Ni-P/Ni-B-W dublex coatings. Prior to the plating process, the surfaces of all the Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 837 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 620.193:621.78:669.058 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)837(2017) 19 S. Maitra, G. English, Mechanism of localized corrosion of 7075 alloy plate, Metallurgical and Materials Transactions A, 12 (1981) 3, 535–541, doi:10.1007/BF02648553 20 H. Bakker, H. Bonzel, C. Bruff, M. Dayananda, W. Gust, J. Horvath, I. Kaur, G. Kidson, A. LeClaire, H. Mehrer, Diffusion in solid metals and alloys/Diffusion in festen metallen und legierungen: vol. 26, Springer, Berlin 1990 21 E. Hollingsworth, H. Hunsicker, Corrosion of aluminum and alumi- num alloys, ASM Handbook, 13 (1987), 583–609 22 J. T. Staley, S. Byrne, E. Colvin, K. Kinnear, Corrosion and stress- corrosion of 7xxx-w products, Materials Science Forum, 217 (1996), 1587–1592, doi:10.4028/MSF.217-222.1587 M. SHAKOURI et al.: DEVELOPMENT OF A HEAT TREATMENT FOR INCREASING THE MECHANICAL PROPERTIES ... 836 Materiali in tehnologije / Materials and technology 51 (2017) 5, 831–836 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... 837–842 CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS DUBLEX Ni-P/Ni-B-W COATINGS KOROZIJSKA ODPORNOST PLATIRANIH IN NEELEKTRI^NO TOPOLOTNO OBDELANIH DUPLEKS Ni-P/Ni-B-W PREVLEK Behiye Yüksel1, Garip Erdogan2, Fatih Erdem Bastan2, Rasid Ahmed Yýldýz3 1Istanbul Aydin University, Mechanical Engineering Department, 34668 Istanbul, Turkey 2Sakarya University, Engineering Faculty, Department of Metallurgy and Materials Engineering, 54187 Sakarya, Turkey 3TU Istanbul, Mechanical Engineering Faculty, 34437 Istanbul, Turkey behiyeyuksel@aydin.edu.tr Prejem rokopisa – received: 2016-10-20; sprejem za objavo – accepted for publication: 2017-03-10 doi:10.17222/mit.2016.304 In this study, Ni-P/Ni-B-W dublex coatings were deposited on carbon steel substrates (AISI 1020) using the electroless plating process and their microstructure and corrosion properties were systematically evaluated based on different heat-treatment temperatures. Both, the surface morphology and cross-sectional morphology of the Ni-P/Ni-B-W coatings were studied using a scanning electron microscope (SEM), while X-ray diffraction (XRD) was applied for examining the structural modifications. The amorphous coating began to crystallize at a heat-treatment temperature of 350 °C. Potentiodynamic polarization measurements were carried out in an aqueous medium containing 3.5 % NaCl for evaluating the corrosion resistance of as-plated and heat-treated dublex coatings. The corrosion potentials of dublex coatings were observed to shift toward more positive values with increased heat-treatment temperatures. Depending on the heat-treatment temperature, it was identified that the crystallized dublex coatings generally had better corrosion resistance than the amorphous coating. Keywords: Ni-P/Ni-B-W coating, corrosion, heat treatment V {tudiji so bile dvojne Ni-P/Ni-B-W prevleke nane{ene na substrate iz ogljikovega jekla (AISI 1020) s postopkom neelek- tri~nega platiranja. Sistemati~no so bile ocenjene mikrostruktura ter korozijske lastnosti izdelanih prevlek, glede na razli~ne temperature toplotne obdelave. Morfologije povr{in in profili izdelanih Ni-P/Ni-B-W prevlek so bili pregledani z vrsti~nim elektronskim mikroskopom (SEM), medtem ko so bile z rentgensko difrakcijsko analizo (XRD) ugotovljene strukturne spremembe. Amorfna prevleka je za~ela kristalizirati pri temperaturi toplotne obdelave 350 °C. Korozijska odpornost platiranih in toplotno obdelanih dupleks prevlek je bila ocenjena s pomo~jo meritev potenciodinami~ne polarizacije v 3,5 % vodni razto- pini NaCl. Opazovali so korozijske potenciale dupleks prevlek, da bi dobili bolj pozitivne vrednosti z zvi{animi temperaturami topolotne obdelave. Ugotovljeno je bilo, da imajo kristalizirane dupleks prevleke na splo{no bolj{o korozijsko odpornost kot amorfne prevleke. Klju~ne besede: Ni-P/Ni-B-W dvojna prevleka, korozija, toplotna obdelava 1 INTRODUCTION Studied for the first time in 1946 by Brenner and Riddell, the electroless plating process has been used in many industrial applications because of its superior characteristics over electroplating, such as the ability to plate insulation materials and having a homogeneous coating-thickness distribution.1 Among electroless coat- ing types, electroless nickel coating is the most popular for having good hardness, wear and corrosion resistance properties.2 If the electroless Ni coating family is re- viewed, besides pure Ni coatings, Ni-P and Ni-B coat- ings deposited using reducing agents such as hypo- phosphite, borohydride or dimethylamine borane stand out. The major advantages of Ni-P coatings are their low cost, high corrosion resistance and easy process control.3 Nonetheless, a borohydride reduced nickel coating is harder and has higher wear resistance than tool steel and hard chrome coatings.4,5 On the other hand, some studies – although they are limited in number – have highlighted the addition of tungsten to this coating system for in- creasing its corrosion resistance, which is lower than that of a Ni-P coating.6–9 As a result of the co-deposition of this refractory material (which cannot be reduced in an aqueous solution as metallic tungsten) together with iron group metals, developing coatings with attractive corro- sion and tribological properties is possible.10–12 Recently, research for obtaining multilayer coatings to increase existing coating corrosion resistance to higher levels has also been conducted.13,14 The aim of the present study was to evaluate the effect of different heat-treatment temperatures on the microstructure and corrosion properties of electroless Ni-P/Ni-B-W coatings. To the best of our knowledge, no others studies are available on this topic. 2 EXPERIMENTAL PART For this experiment, 10 mm × 10 mm × 60 mm plain carbon steel (AISI 1020) plates were used as a substrate material for developing Ni-P/Ni-B-W dublex coatings. Prior to the plating process, the surfaces of all the Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 837 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 620.193:621.78:669.058 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)837(2017) samples were mechanically cleaned (up to 1200 grade) and then soaked in trichloroethylene and cleaned with detergent in an ultrasonic bath at 70 °C. Finally, sample surfaces were activated in 30 % of volume fractions. HCl for two minutes and then rinsed in distilled water. The cleaned substrates were then soaked in Ni-P and Ni-B-W electroless plating baths for two hours, respectively. Reagent-grade chemicals and distilled water were used for the preparation of all the electrolytes and the pH was adjusted to 4.8±0.2 and 13.5±0.2 with H2SO4 or NaOH in Ni-P and Ni-B-W plating baths, respectively. The temperature of the baths was maintained at 90±2 °C for Ni-P and 88±2 °C for Ni-B-W. Deposition of Ni-P and Ni-B-W was carried out in an aqueous bath containing 15 g/L NiSO4.6H2O, 26 g/L Na2H2PO2·H2O, 13 g/L NaC2H3O2, 12 mL/L HF and 8 g/L NH4HF2 for the Ni-P coating, and 24 g/L NiCl2, 60 mL/L EDTA, 26.5 g/L KOH, 120 g/L NaBH4, 263 g/L NaOH, 2.6g/L PbWO4, 13 g/L EDTA and 40 g/L Na2WO4·2H2O for Ni-B-W coatings, respectively. The coated specimens were heat treated at 250 °C, 300 °C, 350 °C and 650 °C for 1 h each at a heating rate of 5 °C/min in an air-circulated furnace, followed by slow furnace cooling to room temperature. The composition of the dublex coatings was deter- mined by energy-dispersive X-ray analysis (EDX) using a Link Analytical QX-2000 attached to the SEM appa- ratus. The crystal structures of the dublex coatings were examined using a Philips PW 3710 grazing incidence x-ray diffractometer (Cu-K radiation). A Jeol JSM-7000F FE-SEM was used to characterize the microstructures and morphology of the coatings. The corrosion behavior was investigated using a PGZ 301 Dynamic Voltammetry and VoltaMaster4 software. Electrochemical experiments were carried out in a 3.5 % NaCl aqueous solution in a three-electrode cell at room temperature. In this cell, a platinum electrode was used as a counter electrode and a saturated calomel electrode was used as a reference electrode; the dublex coating was used as a working electrode and by masking with silicon, only a 1 cm2 surface area was left on the samples for exposure to the electrolyte. The dynamic potential scanning technique was used for obtaining the polarization curves of dublex coating samples. Prior to the potentiodynamic measurement, samples were held for approximately 15 min in the electrolyte and then electrode potential was raised from -600mV to +200 mV at a rate of 10 mV/min. 3 RESULTS The XRD patterns of the as-plated and heat-treated Ni-P/Ni-B-W coatings, and their chemical compositions are in Figure 1 and in Table 1. SEM micrographs of Ni-P/Ni-B-W coatings before and after the heat treat- ments are in Figure 2. Table 1: Chemical compositions of the Ni-P/Ni-B-W dublex coatings Type of coating Ni (w/%) P (w/%) B (w/%) W (w/%) Ni-P 87.27 12.73 - - Ni-B-W 87.45 - 7.86 4.69 B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... 838 Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Surface morphology of dublex coating a), b) as-plated Ni-P/Ni-B-W, c) heat treated at 300 °C, d) heat treated at 650 °C Figure 1: XRD spectra for a): as-plated Ni-P/Ni-B-W, b) 250 °C, c) 300 °C, d) 350 °C and e) 650 °C The cross-sections of the Ni-P/Ni-B-W dublex coat- ings are shown in Figure 3. Potentiodynamic curves of as-plated and heat-treated Ni-P/Ni-B-W coatings at different temperatures obtained in a 3.5 % NaCl aqueous medium are shown in Figure 4. The changes in the corrosion current density (Icorr) values with heat treatment are given in Table 2. Table 2: Corrosion resistance of as-plated and heat-treated Ni-P/Ni-B-W dublex coatings at different temperatures in 3.5 % NaCl solution NiP/NiBW Icorr (μA/cm2) Ecorr (mV) As-plated 1.55±0.59 -506±32 300 °C 1.70±0.41 -470±23 350 °C 0.71±0.20 -382 ±18 650 °C 0.57±0.09 -280±11 The surface morphologies of the as-plated sample and heat-treated sample at 650 °C following potentio- dynamic polarization measurements are shown in Fig- ure 5. The corrosion current density and corrosion potential values obtained by Tafel interpolation are shown in Table 3. Table 3: Corrosion resistance of substrate and Ni-B-W and Ni-P/Ni-B-W dublex coatings heat treated at 650 °C in 3.5 % NaCl solution Sample Icorr (μA/cm2) Ecorr (mV) Carbon steel substrate 4.59±1.01 -598±46 NiBW 3.40±0.35 -298±17 NiP/NiBW 0.57±0.09 -280±11 Figure 6 shows the polarization curves of the Ni-B-W and dublex Ni-P/Ni-B-W coatings heat treated at 650 °C and the steel substrate obtained in a 3.5 % NaCl aqueous medium. 4 DISCUSSION The Ni-P/Ni-B-W coating diffraction patterns up to 300 °C reveal a single broad peak indicating an amor- phous coating structure under these conditions (Fig- B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 839 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: Cross-section of Ni-P/Ni-B-W dublex coating: a) as-plated, b) heat treated at 250 °C c) heat treated at 350 °C, d) heat treated at Figure 4: Polarization curves in a 3.5 % NaCl aqueous medium of as-plated and heat-treated Ni-P/Ni-B-W coatings: a) as-plated NiP/NiBW, b) heat treated at 300 °C, c) heat treated at 350 °C d) heat treated at 650 °C Figure 5: Surface morphology of NiP/NiBW dublex coatings detected by SEM after potentiodynamic measurement: a) as-plated, b) heat treated at 650 °C samples were mechanically cleaned (up to 1200 grade) and then soaked in trichloroethylene and cleaned with detergent in an ultrasonic bath at 70 °C. Finally, sample surfaces were activated in 30 % of volume fractions. HCl for two minutes and then rinsed in distilled water. The cleaned substrates were then soaked in Ni-P and Ni-B-W electroless plating baths for two hours, respectively. Reagent-grade chemicals and distilled water were used for the preparation of all the electrolytes and the pH was adjusted to 4.8±0.2 and 13.5±0.2 with H2SO4 or NaOH in Ni-P and Ni-B-W plating baths, respectively. The temperature of the baths was maintained at 90±2 °C for Ni-P and 88±2 °C for Ni-B-W. Deposition of Ni-P and Ni-B-W was carried out in an aqueous bath containing 15 g/L NiSO4.6H2O, 26 g/L Na2H2PO2·H2O, 13 g/L NaC2H3O2, 12 mL/L HF and 8 g/L NH4HF2 for the Ni-P coating, and 24 g/L NiCl2, 60 mL/L EDTA, 26.5 g/L KOH, 120 g/L NaBH4, 263 g/L NaOH, 2.6g/L PbWO4, 13 g/L EDTA and 40 g/L Na2WO4·2H2O for Ni-B-W coatings, respectively. The coated specimens were heat treated at 250 °C, 300 °C, 350 °C and 650 °C for 1 h each at a heating rate of 5 °C/min in an air-circulated furnace, followed by slow furnace cooling to room temperature. The composition of the dublex coatings was deter- mined by energy-dispersive X-ray analysis (EDX) using a Link Analytical QX-2000 attached to the SEM appa- ratus. The crystal structures of the dublex coatings were examined using a Philips PW 3710 grazing incidence x-ray diffractometer (Cu-K radiation). A Jeol JSM-7000F FE-SEM was used to characterize the microstructures and morphology of the coatings. The corrosion behavior was investigated using a PGZ 301 Dynamic Voltammetry and VoltaMaster4 software. Electrochemical experiments were carried out in a 3.5 % NaCl aqueous solution in a three-electrode cell at room temperature. In this cell, a platinum electrode was used as a counter electrode and a saturated calomel electrode was used as a reference electrode; the dublex coating was used as a working electrode and by masking with silicon, only a 1 cm2 surface area was left on the samples for exposure to the electrolyte. The dynamic potential scanning technique was used for obtaining the polarization curves of dublex coating samples. Prior to the potentiodynamic measurement, samples were held for approximately 15 min in the electrolyte and then electrode potential was raised from -600mV to +200 mV at a rate of 10 mV/min. 3 RESULTS The XRD patterns of the as-plated and heat-treated Ni-P/Ni-B-W coatings, and their chemical compositions are in Figure 1 and in Table 1. SEM micrographs of Ni-P/Ni-B-W coatings before and after the heat treat- ments are in Figure 2. Table 1: Chemical compositions of the Ni-P/Ni-B-W dublex coatings Type of coating Ni (w/%) P (w/%) B (w/%) W (w/%) Ni-P 87.27 12.73 - - Ni-B-W 87.45 - 7.86 4.69 B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... 838 Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Surface morphology of dublex coating a), b) as-plated Ni-P/Ni-B-W, c) heat treated at 300 °C, d) heat treated at 650 °C Figure 1: XRD spectra for a): as-plated Ni-P/Ni-B-W, b) 250 °C, c) 300 °C, d) 350 °C and e) 650 °C The cross-sections of the Ni-P/Ni-B-W dublex coat- ings are shown in Figure 3. Potentiodynamic curves of as-plated and heat-treated Ni-P/Ni-B-W coatings at different temperatures obtained in a 3.5 % NaCl aqueous medium are shown in Figure 4. The changes in the corrosion current density (Icorr) values with heat treatment are given in Table 2. Table 2: Corrosion resistance of as-plated and heat-treated Ni-P/Ni-B-W dublex coatings at different temperatures in 3.5 % NaCl solution NiP/NiBW Icorr (μA/cm2) Ecorr (mV) As-plated 1.55±0.59 -506±32 300 °C 1.70±0.41 -470±23 350 °C 0.71±0.20 -382 ±18 650 °C 0.57±0.09 -280±11 The surface morphologies of the as-plated sample and heat-treated sample at 650 °C following potentio- dynamic polarization measurements are shown in Fig- ure 5. The corrosion current density and corrosion potential values obtained by Tafel interpolation are shown in Table 3. Table 3: Corrosion resistance of substrate and Ni-B-W and Ni-P/Ni-B-W dublex coatings heat treated at 650 °C in 3.5 % NaCl solution Sample Icorr (μA/cm2) Ecorr (mV) Carbon steel substrate 4.59±1.01 -598±46 NiBW 3.40±0.35 -298±17 NiP/NiBW 0.57±0.09 -280±11 Figure 6 shows the polarization curves of the Ni-B-W and dublex Ni-P/Ni-B-W coatings heat treated at 650 °C and the steel substrate obtained in a 3.5 % NaCl aqueous medium. 4 DISCUSSION The Ni-P/Ni-B-W coating diffraction patterns up to 300 °C reveal a single broad peak indicating an amor- phous coating structure under these conditions (Fig- B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 839 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: Cross-section of Ni-P/Ni-B-W dublex coating: a) as-plated, b) heat treated at 250 °C c) heat treated at 350 °C, d) heat treated at Figure 4: Polarization curves in a 3.5 % NaCl aqueous medium of as-plated and heat-treated Ni-P/Ni-B-W coatings: a) as-plated NiP/NiBW, b) heat treated at 300 °C, c) heat treated at 350 °C d) heat treated at 650 °C Figure 5: Surface morphology of NiP/NiBW dublex coatings detected by SEM after potentiodynamic measurement: a) as-plated, b) heat treated at 650 °C ure 1). For heat-treatment temperatures up to 300 °C, the dublex coating preserved its amorphous structure. However, after heat treatment at and above 300 °C, the diffraction patterns showed the formation of Ni3B inter- metallic crystals. For the higher heat-treatment tempe- ratures up to 650 °C, formation of the nickel borides progressively increased. Drovosekov et al., reported chemical bond formations between boron and tungsten in the Ni-B-W coating system. However, the major nickel borides formed during the heat treatments might have prevented the formation of the B-W bond.9,15 In other words, the quantity of the boron in the coating was high enough to prevent the formation of Ni-W com- pounds in the structure.16 SEM micrographs of the Ni-P/Ni-B-W coatings be- fore and after the heat treatments showed the formation of typical spherical nodular structures (Figure 2). Each nodule appears to be formed by the unification of many grains existing as colonies (Figure 2b), similar to the case suggested by S. Ruan et al.17 Some microcracks were noticeable on the surfaces of the as-plated Ni-P/Ni-B-W samples. The microcracks may be due to the internal stress caused by the relatively larger tungsten atoms in the coating structure.18,19 With the heat treatments, the structure progressively crystallized and, as consequence, a substantial reduction of the microcracks was noticed (Figure 2c and 2d). In a Ni-B-W coating system, W atoms may selectively exist at the boundaries of the nodules, from there, they may segregates to the coating surface, during the heat treat- ments.16,17 This process may decrease the internal stress inside the coating, resulting in a substantial decrease of the microcracks on the coating surface. The thickness of the Ni-P coating in the dublex coating system was measured as 2 μm, while the average thickness of the Ni-B-W (the other coating in the system) was measured as 11 μm (Figure 3). On the other hand, it was obvious that the nodular structure seen in the surface morphology of the Ni-B-W coating showed columnar growth starting from the Ni-P layer to the topmost surface of the coating (Figure 3a). Furthermore, it was observed that adherence of the Ni-P coating to the substrate was good and there was no pore formation between the substrate and Ni-P coating. Pores seen in the interface of Ni-P/Ni-B-W, in addition to pores partially observed inside the Ni-B-W coating had a columnar structure and they were considered to have been formed due to the capturing of released hydrogen – even if in small amounts – during the oxidation reaction of sodium borohydride used as a the boron source.5 The pores disappear with increasing heat-treatment temperatures. This may be a result of the diffusion of hydrogen atoms to the surface. They may promote better crystallization of the coating and ease the formation of the Ni3B and Ni2B phases (Figure 3b and 3c). After 650 °C heat treatment the major phase of the coating was noticed to be Ni3B. At that stage adherence between the coatings reached its best level, while the Ni-B-W coating showed columnar grains with a denser structure (Figure 3d). As Figure 4, a difference of approximately 220 mV was detected between corrosion potentials (Ecorr) of as-plated and 650 °C heat-treated dublex coatings. This is the shifting of Ecorr values of the coatings toward nobler values with increased heat treatments applied to the coatings. The anodic branch of the polarization curve of the coating before heat treatment showed the break- down potential at values around -310 mV (Figure 4a). But the current density increased at a constant rate, indicating that the unstable passive film formation on the coating has disappeared rapidly. For heat treatment at 300 °C, the breakdown potential was observed to shift toward -220 mV. At 300 °C, which is the onset tem- perature of crystallization, corrosion potential shifted towards a positive end approximately 40 mV, compared to the as-plated sample. At 350 °C, crystallization was nearly completed and the corrosion potential had shifted towards more positive values. However, the breakdown potential was observed (as with other samples) at around -180 mV. Nonetheless, it was detected that the current density of this sample became higher as the potential value increased compared with that for the to as-plated and heat-treated samples at 300 °C. For the heat treat- ment at 650 °C, a passive film formation on the coating surface was seen at around -190 mV, suggesting that corrosion rate was restrained at such potential range. However, as this protective film began to dissolve at around -90 mV, a steep increase in the current density was observed. The major reason for the difference between the corrosion resistance of as-plated and heat-treated Ni-P/Ni-B-W coatings exposed to chloride ions was undoubtedly associated with the transformation of the microstructure from an amorphous to a crystalline mor- phology. The main reason for having a higher Ecorr value the heat-treated coating at 300 °C (where an amorphous B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... 840 Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: Polarization curves: a) steel substrate, b) Ni-B-W coating heat treated at 650 °C, c) dublex Ni-P/Ni-B-W coating heat treated at 650 °C structure still prevailed) compared to the as-plated sample was the result of the borides formed during the onset of crystallization. The heat-treated sample at 350 °C (where the microstructure had completely been transformed to a crystalline structure) had a higher corrosion resistance compared to the sample heat-treated at 300 °C.20 On the other hand, it is a known fact that corrosion resistance generally increases with addition of tungsten to a electroless nickel-boron coating system, because of the tendency to form an oxide film on the surfaces.7,18 However, as seen in Figure 4, the corrosion resistance of the as-plated Ni-P/Ni-B-W sample having a nodular structure did not increase, with the addition of tungsten. This is because the tungsten atoms were stuck in the colony boundaries, not to be segregated to the surfaces during the corrosion process and therefore, could not be oxidized to promote the formation of a mixed oxide film. However, depending on the increased heat treatment temperatures segregation of W atoms toward surfaces at 650 °C was more likely to occur. The mixed oxide film a passive zone formed this way may be the main reason for the substantial decrease of the current density as observed here. It can be seen that uniform corrosion occurred on the surface of the as-plated sample in the amorphous structure that has the lowest corrosion resistance (Fig- ure 5a). Conversely, on the surface of the heat-treated sample at 650 °C complete crystallization were noted and the pits (indicated by arrows) in the grains within the colonies were observed (Figure 5b). To interpret the effect of the dublex coating on corrosion resistance, the potentiodynamic polarization measurements of the Ni-P/Ni-B-W and Ni-B-W coatings, as well as which were deposited under the same process conditions and heat-treated at 650 °C, and carbon steel substrate were studied (Table 3). As can be observed from Figure 6, while the Ni-B-W coating and the dublex coating almost had the same corrosion potential, more positive values at about 300 mV, compared with that of the substrate. Among the coating studied the dublex coating has the lowest corro- sion current density. The anodic branches of polarization curves of both the dublex and the Ni-B-W coatings have a passive zone, and both coatings had an Epit value of approximately -90 mV. Additionally, the current density of the Ni-B-W coating at a constant rate increases, depending on the increased potential value. On the other hand, the anodic branch of the dublex coating indicated that an entrance to a second passive zone at approxi- mately 20 mV. This passive zone may be related to the existence of Ni-P in the dublex coating. A similar inter- pretation was reported by Zhang et al. for the corrosion of Ni-B-W surfaces.14 In general, corrosion starts at the outermost surface of the substrate and proceed inward. However, the stable passive film formed on the Ni-P layer on the substrate may act as a barrier for corrosion propagation, causing the dublex coating to have better corrosion resistance than the single-layer Ni-B-W coating. 5 CONCLUSIONS The effect of heat treatments on the corrosion resis- tance of a Ni-P/Ni-B-W dublex coating deposited on carbon steel was studied. The Ni-P/Ni-B-W coatings have an amorphous structure for a heat treatments up to 300 °C. The coatings start to crystallize at about 350 °C and completed at about 650 °C with the major phases being the nickel borides. The heat treatment, besides causing nickel boride formation within Ni-B-W, also caused obtaining a denser Ni-P/Ni-B-W coating when increasing heat treatment temperature. The corrosion resistance of the dublex coating in- creases substantially with increasing the heat treatment temperature. The dublex coating and the single-layer Ni-B-W coating heat-treated at 650 °C indicated higher corrosion resistance as compared to that of steel sub- strate. When these two coatings were compared in terms of corrosion resistance, the dublex coating showed better performance. 6 REFERENCES 1 A. Brenner, G. Riddell, Nickel Plating on Steel by Chemical Reduction, J. Res. Nat. Bur. Std., 37 (1946), 31–35 2 J. Sudagar, J. Lian, W. Sha, Electroless nickel, alloy, composite and nano coatings-A critical review, J. Alloys Compd., 571(2013), 183–204, doi:10.1016/j.jallcom.2013.03.107 3 Y. F. Shen, W.Y. Xue, Z.Y. Liu, L. Zuo, Nanoscratching Deformation and Fracture Toughness of Electroless Ni-P Coatings, Surf. Coat. Technol., 205 (2010), 632–640, doi:10.1016/j.surfcoat.2010.07.066 4 K. M. Gorbunova, M.V. Ivanov, V. P. Moissev, Electroless Depo- sition of Nickel-Boron Alloys – Mechanýsm of Process, Structure, and Some Propertýes of Deposits, J. Electrochem. Soc., 120 (1973), 613–618, doi:10.1149/1.2403514 5 R. N. Duncan, T. L. Arney, Operatýon and Use of Sodium-Boro- hydride-Reduced Electroless Nickel, Plat. Surf. Finish., 71 (1984) 12, 49–54 6 R. A. C. Santana, S. Prasad, A. R. N. Campos, F. O. Arau´ jo, G. P. DA Silva, P. Lýma-Neto, Electrodeposition and Mechanical Pro- perties of Ni–W–B Composites from Tartrate Bath, Prot Met Phys Chem., 46 (2010) 1, 117–122, doi:10.1134/S207020511001017X 7 M. G. Hosseini, M. Abdolmaleki , H. Ebrahimzadeh, S. A. Seyed Sadjadi, Effect of 2-Butyne-1, 4-Diol on the Nanostructure and Corrosion Resistance Properties of Electrodeposited Ni-W-B Coatings, Int. J. Electrochem. Sci., 6 (2011) 4, 1189–1205 8 T. Osaka, N. Takano, T. Kurokawa, T. Kaneko, K. Ueno, Characteri- zation of chemically-deposited NiB and NiWB thin films as a capping layer for ULSI application, Surf. Coat. Technol., 169–170 (2003), 124–127, doi:10.1016/S0257-8972(03)00186-5 9 G. Graef, K. Anderson, J. Groza, A. Palazoglu, Phase evolution in electrodeposited Ni-W-B alloy, Mater. Sci. Eng., B41 (1996), 253–257, doi:10.1016/S0921-5107(96)01656-X 10 E. Lassner, W. D. Schubert, Tungsten: Properties, Chemistry, Tech- nology of the Element, Alloys, and Chemical Compounds, Kluwer Academic, New York, 1999 11 E. Beltowska-Lehman, Kinetics of Induced Electrodeposition of Alloys Containing Mo from Citrate Solutions, Phys. Stat. Sol., 5 (2008) 11, 3514–3520, doi:10.1002/pssc.200779404 B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 841 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS ure 1). For heat-treatment temperatures up to 300 °C, the dublex coating preserved its amorphous structure. However, after heat treatment at and above 300 °C, the diffraction patterns showed the formation of Ni3B inter- metallic crystals. For the higher heat-treatment tempe- ratures up to 650 °C, formation of the nickel borides progressively increased. Drovosekov et al., reported chemical bond formations between boron and tungsten in the Ni-B-W coating system. However, the major nickel borides formed during the heat treatments might have prevented the formation of the B-W bond.9,15 In other words, the quantity of the boron in the coating was high enough to prevent the formation of Ni-W com- pounds in the structure.16 SEM micrographs of the Ni-P/Ni-B-W coatings be- fore and after the heat treatments showed the formation of typical spherical nodular structures (Figure 2). Each nodule appears to be formed by the unification of many grains existing as colonies (Figure 2b), similar to the case suggested by S. Ruan et al.17 Some microcracks were noticeable on the surfaces of the as-plated Ni-P/Ni-B-W samples. The microcracks may be due to the internal stress caused by the relatively larger tungsten atoms in the coating structure.18,19 With the heat treatments, the structure progressively crystallized and, as consequence, a substantial reduction of the microcracks was noticed (Figure 2c and 2d). In a Ni-B-W coating system, W atoms may selectively exist at the boundaries of the nodules, from there, they may segregates to the coating surface, during the heat treat- ments.16,17 This process may decrease the internal stress inside the coating, resulting in a substantial decrease of the microcracks on the coating surface. The thickness of the Ni-P coating in the dublex coating system was measured as 2 μm, while the average thickness of the Ni-B-W (the other coating in the system) was measured as 11 μm (Figure 3). On the other hand, it was obvious that the nodular structure seen in the surface morphology of the Ni-B-W coating showed columnar growth starting from the Ni-P layer to the topmost surface of the coating (Figure 3a). Furthermore, it was observed that adherence of the Ni-P coating to the substrate was good and there was no pore formation between the substrate and Ni-P coating. Pores seen in the interface of Ni-P/Ni-B-W, in addition to pores partially observed inside the Ni-B-W coating had a columnar structure and they were considered to have been formed due to the capturing of released hydrogen – even if in small amounts – during the oxidation reaction of sodium borohydride used as a the boron source.5 The pores disappear with increasing heat-treatment temperatures. This may be a result of the diffusion of hydrogen atoms to the surface. They may promote better crystallization of the coating and ease the formation of the Ni3B and Ni2B phases (Figure 3b and 3c). After 650 °C heat treatment the major phase of the coating was noticed to be Ni3B. At that stage adherence between the coatings reached its best level, while the Ni-B-W coating showed columnar grains with a denser structure (Figure 3d). As Figure 4, a difference of approximately 220 mV was detected between corrosion potentials (Ecorr) of as-plated and 650 °C heat-treated dublex coatings. This is the shifting of Ecorr values of the coatings toward nobler values with increased heat treatments applied to the coatings. The anodic branch of the polarization curve of the coating before heat treatment showed the break- down potential at values around -310 mV (Figure 4a). But the current density increased at a constant rate, indicating that the unstable passive film formation on the coating has disappeared rapidly. For heat treatment at 300 °C, the breakdown potential was observed to shift toward -220 mV. At 300 °C, which is the onset tem- perature of crystallization, corrosion potential shifted towards a positive end approximately 40 mV, compared to the as-plated sample. At 350 °C, crystallization was nearly completed and the corrosion potential had shifted towards more positive values. However, the breakdown potential was observed (as with other samples) at around -180 mV. Nonetheless, it was detected that the current density of this sample became higher as the potential value increased compared with that for the to as-plated and heat-treated samples at 300 °C. For the heat treat- ment at 650 °C, a passive film formation on the coating surface was seen at around -190 mV, suggesting that corrosion rate was restrained at such potential range. However, as this protective film began to dissolve at around -90 mV, a steep increase in the current density was observed. The major reason for the difference between the corrosion resistance of as-plated and heat-treated Ni-P/Ni-B-W coatings exposed to chloride ions was undoubtedly associated with the transformation of the microstructure from an amorphous to a crystalline mor- phology. The main reason for having a higher Ecorr value the heat-treated coating at 300 °C (where an amorphous B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... 840 Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: Polarization curves: a) steel substrate, b) Ni-B-W coating heat treated at 650 °C, c) dublex Ni-P/Ni-B-W coating heat treated at 650 °C structure still prevailed) compared to the as-plated sample was the result of the borides formed during the onset of crystallization. The heat-treated sample at 350 °C (where the microstructure had completely been transformed to a crystalline structure) had a higher corrosion resistance compared to the sample heat-treated at 300 °C.20 On the other hand, it is a known fact that corrosion resistance generally increases with addition of tungsten to a electroless nickel-boron coating system, because of the tendency to form an oxide film on the surfaces.7,18 However, as seen in Figure 4, the corrosion resistance of the as-plated Ni-P/Ni-B-W sample having a nodular structure did not increase, with the addition of tungsten. This is because the tungsten atoms were stuck in the colony boundaries, not to be segregated to the surfaces during the corrosion process and therefore, could not be oxidized to promote the formation of a mixed oxide film. However, depending on the increased heat treatment temperatures segregation of W atoms toward surfaces at 650 °C was more likely to occur. The mixed oxide film a passive zone formed this way may be the main reason for the substantial decrease of the current density as observed here. It can be seen that uniform corrosion occurred on the surface of the as-plated sample in the amorphous structure that has the lowest corrosion resistance (Fig- ure 5a). Conversely, on the surface of the heat-treated sample at 650 °C complete crystallization were noted and the pits (indicated by arrows) in the grains within the colonies were observed (Figure 5b). To interpret the effect of the dublex coating on corrosion resistance, the potentiodynamic polarization measurements of the Ni-P/Ni-B-W and Ni-B-W coatings, as well as which were deposited under the same process conditions and heat-treated at 650 °C, and carbon steel substrate were studied (Table 3). As can be observed from Figure 6, while the Ni-B-W coating and the dublex coating almost had the same corrosion potential, more positive values at about 300 mV, compared with that of the substrate. Among the coating studied the dublex coating has the lowest corro- sion current density. The anodic branches of polarization curves of both the dublex and the Ni-B-W coatings have a passive zone, and both coatings had an Epit value of approximately -90 mV. Additionally, the current density of the Ni-B-W coating at a constant rate increases, depending on the increased potential value. On the other hand, the anodic branch of the dublex coating indicated that an entrance to a second passive zone at approxi- mately 20 mV. This passive zone may be related to the existence of Ni-P in the dublex coating. A similar inter- pretation was reported by Zhang et al. for the corrosion of Ni-B-W surfaces.14 In general, corrosion starts at the outermost surface of the substrate and proceed inward. However, the stable passive film formed on the Ni-P layer on the substrate may act as a barrier for corrosion propagation, causing the dublex coating to have better corrosion resistance than the single-layer Ni-B-W coating. 5 CONCLUSIONS The effect of heat treatments on the corrosion resis- tance of a Ni-P/Ni-B-W dublex coating deposited on carbon steel was studied. The Ni-P/Ni-B-W coatings have an amorphous structure for a heat treatments up to 300 °C. The coatings start to crystallize at about 350 °C and completed at about 650 °C with the major phases being the nickel borides. The heat treatment, besides causing nickel boride formation within Ni-B-W, also caused obtaining a denser Ni-P/Ni-B-W coating when increasing heat treatment temperature. The corrosion resistance of the dublex coating in- creases substantially with increasing the heat treatment temperature. The dublex coating and the single-layer Ni-B-W coating heat-treated at 650 °C indicated higher corrosion resistance as compared to that of steel sub- strate. When these two coatings were compared in terms of corrosion resistance, the dublex coating showed better performance. 6 REFERENCES 1 A. Brenner, G. Riddell, Nickel Plating on Steel by Chemical Reduction, J. Res. Nat. Bur. Std., 37 (1946), 31–35 2 J. Sudagar, J. Lian, W. Sha, Electroless nickel, alloy, composite and nano coatings-A critical review, J. Alloys Compd., 571(2013), 183–204, doi:10.1016/j.jallcom.2013.03.107 3 Y. F. Shen, W.Y. Xue, Z.Y. Liu, L. Zuo, Nanoscratching Deformation and Fracture Toughness of Electroless Ni-P Coatings, Surf. Coat. Technol., 205 (2010), 632–640, doi:10.1016/j.surfcoat.2010.07.066 4 K. M. Gorbunova, M.V. Ivanov, V. P. Moissev, Electroless Depo- sition of Nickel-Boron Alloys – Mechanýsm of Process, Structure, and Some Propertýes of Deposits, J. Electrochem. Soc., 120 (1973), 613–618, doi:10.1149/1.2403514 5 R. N. Duncan, T. L. Arney, Operatýon and Use of Sodium-Boro- hydride-Reduced Electroless Nickel, Plat. Surf. Finish., 71 (1984) 12, 49–54 6 R. A. C. Santana, S. Prasad, A. R. N. Campos, F. O. Arau´ jo, G. P. DA Silva, P. Lýma-Neto, Electrodeposition and Mechanical Pro- perties of Ni–W–B Composites from Tartrate Bath, Prot Met Phys Chem., 46 (2010) 1, 117–122, doi:10.1134/S207020511001017X 7 M. G. Hosseini, M. Abdolmaleki , H. Ebrahimzadeh, S. A. Seyed Sadjadi, Effect of 2-Butyne-1, 4-Diol on the Nanostructure and Corrosion Resistance Properties of Electrodeposited Ni-W-B Coatings, Int. J. Electrochem. Sci., 6 (2011) 4, 1189–1205 8 T. Osaka, N. Takano, T. Kurokawa, T. Kaneko, K. Ueno, Characteri- zation of chemically-deposited NiB and NiWB thin films as a capping layer for ULSI application, Surf. Coat. Technol., 169–170 (2003), 124–127, doi:10.1016/S0257-8972(03)00186-5 9 G. Graef, K. Anderson, J. Groza, A. Palazoglu, Phase evolution in electrodeposited Ni-W-B alloy, Mater. Sci. Eng., B41 (1996), 253–257, doi:10.1016/S0921-5107(96)01656-X 10 E. Lassner, W. D. Schubert, Tungsten: Properties, Chemistry, Tech- nology of the Element, Alloys, and Chemical Compounds, Kluwer Academic, New York, 1999 11 E. Beltowska-Lehman, Kinetics of Induced Electrodeposition of Alloys Containing Mo from Citrate Solutions, Phys. Stat. Sol., 5 (2008) 11, 3514–3520, doi:10.1002/pssc.200779404 B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 841 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 12 A. I. Aydeniz, A. Göksenli, G. Dil, F. Muhaffel, C. Calli, B. Yüksel, Electroless Ni-B-W Coatings for Improving Hardness, Wear And Corrosion Resistance, MTAEC9, 47 (2013) 6, 803–806 13 T. S. N. Sankara Narayanan, K. Krishnaveni, S. K. Seshadri, Elec- troless Ni–P/Ni–B Dublex Coatings: Preparation and Evaluation of Microhardness, Wear and Corrosion Resistance, Mater. Chem. Phys., 82 (2003), 771–779, doi:10.1016/S0254-0584(03)00390-0 14 W. X. Zhang, Z. H. Jiang, G. Y. Li, Q. Jiang, J. S. Lian, Electroless NiP/NiB Dublex Coatings For Improving The Hardness and The Corrosion Resistance of AZ91 D Magnesium Alloy, Appl. Surf. Sci., 254 (2008), 4949–4955, doi:10.1016/j.apsusc.2008.01.144 15 A. B. Drovosekov, M.V. Ivanov, V. M. Krutskikh, E. N. Lubnin, Y. M. Polukarov, Chemically Deposited Ni-W-B Coatings: Composi- tion, Structure and Properties, Prot. Met., 41(2005) 1, 55–62, doi:10.1007/s11124-005-0008-1 16 F. Z. Yang, Z. H. Ma, L. Huang, S. K. Xu, S. M. Zhou, Electro- deposition and Properties of an Amorphous Ni-W-B Alloy before and after Heat Treatment, Chin. J. Chem., 24 (2006), 114–118, doi:10.1002/cjoc.200690004 17 S. Ruan, C. A. Schu, Mesoscale Structure and Segregation in Electrodeposited Nanocrystalline Alloys, Scripta Mater., 59 (2008) 11, 1218–1221, doi:10.1016/j.scriptamat.2008.08.010 18 R. A.C. Santana, S. Prasad, A. R. N. Campos, F. O. Arau´ jo, G.P. Sýlva, P. Lima-Neto, Electrodeposition and Corrosion Behaviour of a Ni–W–B Amorphous Alloy, J. Appl. Electrochem., 36 (2006), 105–113, doi:10.1007/s10800-005-9046-2 19 R. A. Yildiz, A. Göksenli, B. Yüksel, F. Muhaffel, A. Aydeniz, Effect of Annealing Temperature on The Corrosion Resistance of Elec- troless Produced Ni-B-W Coatings, Adv Mat Res., 652 (2013), 263–268, doi:10.4028/www.scientific.net/AMR.651.263 20 Z. Z. Hamid, H. B. Hassan, A. M. Attyia, Influence of Deposition Temperature and Heat Treatment on The Performance of Electroless Ni–B Films, Surf. Coat. Technol., 205 (2010) 7, 2348–2354, doi:10.1016/j.surfcoat.2010.09.025 B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... 842 Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS J. ZHE et al.: SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES ... 843–847 SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES AND AN INSTANTANEOUS-STRESS-CHANGE TESTING KRATKOTRAJNO LEZENJE TOPLOTNO ODPORNEGA JEKLA P91 PRI NIZKIH NAPETOSTIH IN NENADNI MENJAVI NAPETOSTI OBREMENJEVANJA Jiang Zhe, Shen Junjie, Zhao Pengshuo Tianjin University of Technology, Tianjin Key Laboratory for Advanced Mechatronic System Design and Intelligent Control, National Demonstration Center for Experimental Mechanical and Electrical Engineering Education, 391 Binshui Road, Xiqing District, Tianjin, China sjj1982428@sina.com, j211209977@live.com Prejem rokopisa – received: 2016-1'-20; sprejem za objavo – accepted for publication: 2017-03-21 doi:10.17222/mit.2016.305 For the short-term creep behavior to be evaluated and the creep mechanism of P91 heat-resistant steels at low stresses and high temperatures to be clarified, stress-change testing was conducted with a “helicoidal-spring creep test” demonstrating a high strain resolution. The creep deformation consists of the primary creep stage, whereas no secondary creep stage was observed. Blackburn’s law was suggested to be the best choice for a short-term-creep-behavior description because it provides a good representation of an experimental creep curve. An anelastic backflow at a low stress was confirmed, following a high reduction in the stress. The absolute value of the instantaneous strain for a load increase was equal to the value for a load decrease and the creep of the P91 steels at low stresses might have been controlled by the viscous glide of dislocations. Keywords: P91 heat-resistant steel, creep, anelastic, stress change Da bi bilo mo~ oceniti kratkotrajno lezenje in njegov mehanizem toplotno obstojnega jekla P91 pri nizkih napetostih in visokih temperaturah, so bili izvedeni preizkusi spremembe napetosti z uporabo t.i. testa spiralne vzmeti, ki omogo~a veliko lo~ljivost deformacije. Deformacija lezenja sestoji iz primarne in sekundarne faze lezenja, vendar sekundarne faze niso analizirali oz. opazovali. Blackburnov zakon je najbolj uporaben zakon za opis obna{anja jekla med kratkotrajnim lezenjem, ker se z njim najbolj pribli`amo rezultatom, dobljenim z eksperimenti. Neelasti~ni povratni tok pri nizki napetosti je bil potrjen z ve~jim zmanj{anjem napetosti. Zaradi absolutne vrednosti trenutne deformacije za dano obremenitev je bil narastek deformacije enak zmanj{anju obremenitve, medtem ko je lezenje pri nizkih napetostih jekla P91 kontrolirano z viskoznim drsenjem dislokacij. Klju~ne besede: P91 jeklo za delo v vro~em, lezenje, neelasti~nost, sprememba napetosti 1 INTRODUCTION The P91 ferritic heat-resistant steel is utilized for the material production of thermal-power-plant components, which are exposed to elevated temperatures for extensive periods. This requires that the P91 structural materials resist creep deformation at both high temperatures and low stresses. The creep strength obtained with the extra- polation method based on the data of the creep deforma- tion at a high stress is higher than the true creep strength under these conditions (at low stresses).1,2 The creep- strength degradation is not only based on the microstruc- tural degradation3–6, but it is also related to the change in the creep mechanism.7–9 The creep-mechanism clarifica- tion has to depend on the instantaneous creep behavior because the long-term creep is associated with the micro- structural degradation. The stress-change test constitutes an effective method for a creep-deformation-mechanism clarification based on the instantaneous-strain and creep-rate-variation evaluation at the stress changes during the creep testing.10–12 At a high strain rate (a high stress) and a high tem- perature, a sudden-stress-change experiment has been widely utilized for pure metals and solution-strengthened alloys.10–12 In contrast, at a low strain rate 0, especially the ultra-low 0, lower than 10–10 s–1, it is usually impossi- ble for a conventional tension-creep technique to dist- inguish the instantaneous plastic strain from the total strain under a sudden stress change, due to the unsatis- factory strain resolution, existing between 10–6 and 10–5. As an additional creep technique, the helicoidal-spring creep test based on torsion deformation provides a significantly high strain resolution, even up to 10–9.13 Due to this high strain resolution, it is possible for very low instantaneous strains to be measured during a sudden stress change.14 In the present study, the instantaneous creep of the P91 ferritic heat-resistant steels was studied during sudden-stress-change experiments using the helicoidal- spring-specimen technique for the creep mechanism under both a low stress and a high temperature. Materiali in tehnologije / Materials and technology 51 (2017) 5, 843–847 843 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 621.9.011:620.172.24:691.714 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)843(2017) 12 A. I. Aydeniz, A. Göksenli, G. Dil, F. Muhaffel, C. Calli, B. Yüksel, Electroless Ni-B-W Coatings for Improving Hardness, Wear And Corrosion Resistance, MTAEC9, 47 (2013) 6, 803–806 13 T. S. N. Sankara Narayanan, K. Krishnaveni, S. K. Seshadri, Elec- troless Ni–P/Ni–B Dublex Coatings: Preparation and Evaluation of Microhardness, Wear and Corrosion Resistance, Mater. Chem. Phys., 82 (2003), 771–779, doi:10.1016/S0254-0584(03)00390-0 14 W. X. Zhang, Z. H. Jiang, G. Y. Li, Q. Jiang, J. S. Lian, Electroless NiP/NiB Dublex Coatings For Improving The Hardness and The Corrosion Resistance of AZ91 D Magnesium Alloy, Appl. Surf. Sci., 254 (2008), 4949–4955, doi:10.1016/j.apsusc.2008.01.144 15 A. B. Drovosekov, M.V. Ivanov, V. M. Krutskikh, E. N. Lubnin, Y. M. Polukarov, Chemically Deposited Ni-W-B Coatings: Composi- tion, Structure and Properties, Prot. Met., 41(2005) 1, 55–62, doi:10.1007/s11124-005-0008-1 16 F. Z. Yang, Z. H. Ma, L. Huang, S. K. Xu, S. M. Zhou, Electro- deposition and Properties of an Amorphous Ni-W-B Alloy before and after Heat Treatment, Chin. J. Chem., 24 (2006), 114–118, doi:10.1002/cjoc.200690004 17 S. Ruan, C. A. Schu, Mesoscale Structure and Segregation in Electrodeposited Nanocrystalline Alloys, Scripta Mater., 59 (2008) 11, 1218–1221, doi:10.1016/j.scriptamat.2008.08.010 18 R. A.C. Santana, S. Prasad, A. R. N. Campos, F. O. Arau´ jo, G.P. Sýlva, P. Lima-Neto, Electrodeposition and Corrosion Behaviour of a Ni–W–B Amorphous Alloy, J. Appl. Electrochem., 36 (2006), 105–113, doi:10.1007/s10800-005-9046-2 19 R. A. Yildiz, A. Göksenli, B. Yüksel, F. Muhaffel, A. Aydeniz, Effect of Annealing Temperature on The Corrosion Resistance of Elec- troless Produced Ni-B-W Coatings, Adv Mat Res., 652 (2013), 263–268, doi:10.4028/www.scientific.net/AMR.651.263 20 Z. Z. Hamid, H. B. Hassan, A. M. Attyia, Influence of Deposition Temperature and Heat Treatment on The Performance of Electroless Ni–B Films, Surf. Coat. Technol., 205 (2010) 7, 2348–2354, doi:10.1016/j.surfcoat.2010.09.025 B. YÜKSEL et al.: CORROSION RESISTANCE OF AS-PLATED AND HEAT-TREATED ELECTROLESS ... 842 Materiali in tehnologije / Materials and technology 51 (2017) 5, 837–842 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS J. ZHE et al.: SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES ... 843–847 SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES AND AN INSTANTANEOUS-STRESS-CHANGE TESTING KRATKOTRAJNO LEZENJE TOPLOTNO ODPORNEGA JEKLA P91 PRI NIZKIH NAPETOSTIH IN NENADNI MENJAVI NAPETOSTI OBREMENJEVANJA Jiang Zhe, Shen Junjie, Zhao Pengshuo Tianjin University of Technology, Tianjin Key Laboratory for Advanced Mechatronic System Design and Intelligent Control, National Demonstration Center for Experimental Mechanical and Electrical Engineering Education, 391 Binshui Road, Xiqing District, Tianjin, China sjj1982428@sina.com, j211209977@live.com Prejem rokopisa – received: 2016-1'-20; sprejem za objavo – accepted for publication: 2017-03-21 doi:10.17222/mit.2016.305 For the short-term creep behavior to be evaluated and the creep mechanism of P91 heat-resistant steels at low stresses and high temperatures to be clarified, stress-change testing was conducted with a “helicoidal-spring creep test” demonstrating a high strain resolution. The creep deformation consists of the primary creep stage, whereas no secondary creep stage was observed. Blackburn’s law was suggested to be the best choice for a short-term-creep-behavior description because it provides a good representation of an experimental creep curve. An anelastic backflow at a low stress was confirmed, following a high reduction in the stress. The absolute value of the instantaneous strain for a load increase was equal to the value for a load decrease and the creep of the P91 steels at low stresses might have been controlled by the viscous glide of dislocations. Keywords: P91 heat-resistant steel, creep, anelastic, stress change Da bi bilo mo~ oceniti kratkotrajno lezenje in njegov mehanizem toplotno obstojnega jekla P91 pri nizkih napetostih in visokih temperaturah, so bili izvedeni preizkusi spremembe napetosti z uporabo t.i. testa spiralne vzmeti, ki omogo~a veliko lo~ljivost deformacije. Deformacija lezenja sestoji iz primarne in sekundarne faze lezenja, vendar sekundarne faze niso analizirali oz. opazovali. Blackburnov zakon je najbolj uporaben zakon za opis obna{anja jekla med kratkotrajnim lezenjem, ker se z njim najbolj pribli`amo rezultatom, dobljenim z eksperimenti. Neelasti~ni povratni tok pri nizki napetosti je bil potrjen z ve~jim zmanj{anjem napetosti. Zaradi absolutne vrednosti trenutne deformacije za dano obremenitev je bil narastek deformacije enak zmanj{anju obremenitve, medtem ko je lezenje pri nizkih napetostih jekla P91 kontrolirano z viskoznim drsenjem dislokacij. Klju~ne besede: P91 jeklo za delo v vro~em, lezenje, neelasti~nost, sprememba napetosti 1 INTRODUCTION The P91 ferritic heat-resistant steel is utilized for the material production of thermal-power-plant components, which are exposed to elevated temperatures for extensive periods. This requires that the P91 structural materials resist creep deformation at both high temperatures and low stresses. The creep strength obtained with the extra- polation method based on the data of the creep deforma- tion at a high stress is higher than the true creep strength under these conditions (at low stresses).1,2 The creep- strength degradation is not only based on the microstruc- tural degradation3–6, but it is also related to the change in the creep mechanism.7–9 The creep-mechanism clarifica- tion has to depend on the instantaneous creep behavior because the long-term creep is associated with the micro- structural degradation. The stress-change test constitutes an effective method for a creep-deformation-mechanism clarification based on the instantaneous-strain and creep-rate-variation evaluation at the stress changes during the creep testing.10–12 At a high strain rate (a high stress) and a high tem- perature, a sudden-stress-change experiment has been widely utilized for pure metals and solution-strengthened alloys.10–12 In contrast, at a low strain rate 0, especially the ultra-low 0, lower than 10–10 s–1, it is usually impossi- ble for a conventional tension-creep technique to dist- inguish the instantaneous plastic strain from the total strain under a sudden stress change, due to the unsatis- factory strain resolution, existing between 10–6 and 10–5. As an additional creep technique, the helicoidal-spring creep test based on torsion deformation provides a significantly high strain resolution, even up to 10–9.13 Due to this high strain resolution, it is possible for very low instantaneous strains to be measured during a sudden stress change.14 In the present study, the instantaneous creep of the P91 ferritic heat-resistant steels was studied during sudden-stress-change experiments using the helicoidal- spring-specimen technique for the creep mechanism under both a low stress and a high temperature. Materiali in tehnologije / Materials and technology 51 (2017) 5, 843–847 843 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 621.9.011:620.172.24:691.714 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)843(2017) 2 EXPERIMENTAL PART 2.1 Materials The P91 heat-resistant steel was utilized in this expe- riment. The chemical components are presented in Table 1. 2.2 Specimen processing The steel was processed into helicoidal-spring speci- mens using the method of high-speed wire-electrode cutting machining; the outer and inner diameters of the specimens were 12 mm and 8 mm, respectively. The cross-section of the helicoidal-spring specimens was a rectangle. The side length of the rectangle was 2 mm, as presented in Figure 1. The helicoidal-spring specimens were annealed at 1313 K for 60 min and consequently tempered at 1033 K for 60 min. 2.3 Test methods Figure 2 presents the experimental apparatus, which includes an electric furnace with three adjacent heaters, a high-precision optical micrometer, a load, a weight and a system for receiving and processing information. By measuring the helicoidal-spring-specimen pitch change, the creep curves could be obtained. The test-apparatus details were previously reported.14 During the testing, the number of weights was in- creased or decreased for the instantaneous stress change to be observed. The weight was suspended under a load with a flammable thread. In order for a load to be increased, the load was hooked quickly. For a load to be decreased, the flammable thread was burned. The test was performed at a temperature of 923 K and a slight oscillation. The specimen was loaded sub- sequently for five days when the strain rate was down to 10–9–10–8 s–1. Also, the stress consequently increased or decreased at various applied levels and the instantaneous strain was measured. The following equations15,16 were utilized for calculating the mean surface shear stress, , and the surface shear strain, , with the assumption of the pure torsion of the helicoidal-spring specimen:  = PD a b2 2 (1)  = 2 2 a Dπ Δ (2) where P is the average load, D is the coil diameter (12 mm) and  is the displacement of the mean coil-pitch spacing. In this study, the torsion was the dominant component of deformation, because D was quite higher than d (D/d > 12)17 and the value of was between 2 mm and 4 mm.18 Since the stress and strain in the helicoidal spring had essentially shear components, the former could be transformed into the equivalent tensile quantities with the von Mises equations for the tensile stress  = 3 and the tensile strain  = / 3. 3 RESULTS AND DISCUSSION Figure 3 demonstrates the creep curves obtained at 923 K and at various stresses: (34.85, 27.75 and 19.35) MPa. J. ZHE et al.: SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES ... 844 Materiali in tehnologije / Materials and technology 51 (2017) 5, 843–847 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Table 1: Chemical components of P91 Element Cr Mo Si V C Nb Cu N P Ni Ti Al S Content (%) 9.50 0.91 0.60 0.20 0.10 0.10 0.07 0.04 0.02 0.02 0.01 0.01 0.01 Figure 2: Experimental apparatus including electric furnace, high- precision optical micrometer, load, weight and information-receiving and processing systemFigure 1: Helicoidal-spring specimen for the stress-change testing The creep rate which corresponded to the slope of the curves decreased as the duration increased. Figure 4 presents the strain rate on a logarithmic plot versus the strain at 923 K and various stresses. The creep deformation of the P91 steels consisted of the primary creep stages where the creep rate decreased along with the strain and no apparent secondary (steady-state) creep stage was observed where the creep rate would not change along with the strain. The constitutive creep equations expressing the pri- mary stages should be utilized in the experimental creep-curve analysis. The following constitutive creep equations were utilized, being widely accepted as the basic creep equations.19,20 Power law:  = 0 + at b (3) Exponential law:  = 0 + a[1-exp(-bt)] (4) Logarithmic law:  = 0 + a ln(1+bt) (5) Blackburn’s law:  = 0 + ac[1 - exp(-bt)]+c[1 - exp(-dt)] (6) where  is the strain, 0 is the instantaneous strain on dead weight, t is the elapsed time and a, b, c and d are the parameters characterizing the primary creep region. The a, b, c and d values were called the "scaling fac- tors".21 Figure 5 presents representative results of a re- gression analysis, for the creep at 923 K and 27.75 MPa. The exponential-law and power-law equations did not reproduce the experimental data. Compared to the loga- rithmic-law equation, the Blackburn equation provided a better representation of the experimental creep curve. Therefore, the short-term creep for the P91 heat-resistant steels at low stresses could be described with the Blackburn equation. In order for the creep mechanism of the P91 steels at low stresses to be studied, the instantaneous creep behavior was investigated. The specimens were loaded for 4.3 × 105 s until the strain reached 6 × 10–3. In this case, the dislocations could be expected to move. The stress increased or decreased at various levels that were consequently applied and the instantaneous strain was measured. J. ZHE et al.: SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 843–847 845 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: Strain/time creep curves for P91 at 923 K and various stresses Figure 4: Strain rate/strain creep curves at 923 K and various stresses Figure 6: Example of instantaneous elongation and contraction upon low stress changes at a high temperature Figure 5: Comparison of experimental data and predicted curves of strain versus time at 923 K and 27.75 MPa 2 EXPERIMENTAL PART 2.1 Materials The P91 heat-resistant steel was utilized in this expe- riment. The chemical components are presented in Table 1. 2.2 Specimen processing The steel was processed into helicoidal-spring speci- mens using the method of high-speed wire-electrode cutting machining; the outer and inner diameters of the specimens were 12 mm and 8 mm, respectively. The cross-section of the helicoidal-spring specimens was a rectangle. The side length of the rectangle was 2 mm, as presented in Figure 1. The helicoidal-spring specimens were annealed at 1313 K for 60 min and consequently tempered at 1033 K for 60 min. 2.3 Test methods Figure 2 presents the experimental apparatus, which includes an electric furnace with three adjacent heaters, a high-precision optical micrometer, a load, a weight and a system for receiving and processing information. By measuring the helicoidal-spring-specimen pitch change, the creep curves could be obtained. The test-apparatus details were previously reported.14 During the testing, the number of weights was in- creased or decreased for the instantaneous stress change to be observed. The weight was suspended under a load with a flammable thread. In order for a load to be increased, the load was hooked quickly. For a load to be decreased, the flammable thread was burned. The test was performed at a temperature of 923 K and a slight oscillation. The specimen was loaded sub- sequently for five days when the strain rate was down to 10–9–10–8 s–1. Also, the stress consequently increased or decreased at various applied levels and the instantaneous strain was measured. The following equations15,16 were utilized for calculating the mean surface shear stress, , and the surface shear strain, , with the assumption of the pure torsion of the helicoidal-spring specimen:  = PD a b2 2 (1)  = 2 2 a Dπ Δ (2) where P is the average load, D is the coil diameter (12 mm) and  is the displacement of the mean coil-pitch spacing. In this study, the torsion was the dominant component of deformation, because D was quite higher than d (D/d > 12)17 and the value of was between 2 mm and 4 mm.18 Since the stress and strain in the helicoidal spring had essentially shear components, the former could be transformed into the equivalent tensile quantities with the von Mises equations for the tensile stress  = 3 and the tensile strain  = / 3. 3 RESULTS AND DISCUSSION Figure 3 demonstrates the creep curves obtained at 923 K and at various stresses: (34.85, 27.75 and 19.35) MPa. J. ZHE et al.: SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES ... 844 Materiali in tehnologije / Materials and technology 51 (2017) 5, 843–847 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Table 1: Chemical components of P91 Element Cr Mo Si V C Nb Cu N P Ni Ti Al S Content (%) 9.50 0.91 0.60 0.20 0.10 0.10 0.07 0.04 0.02 0.02 0.01 0.01 0.01 Figure 2: Experimental apparatus including electric furnace, high- precision optical micrometer, load, weight and information-receiving and processing systemFigure 1: Helicoidal-spring specimen for the stress-change testing The creep rate which corresponded to the slope of the curves decreased as the duration increased. Figure 4 presents the strain rate on a logarithmic plot versus the strain at 923 K and various stresses. The creep deformation of the P91 steels consisted of the primary creep stages where the creep rate decreased along with the strain and no apparent secondary (steady-state) creep stage was observed where the creep rate would not change along with the strain. The constitutive creep equations expressing the pri- mary stages should be utilized in the experimental creep-curve analysis. The following constitutive creep equations were utilized, being widely accepted as the basic creep equations.19,20 Power law:  = 0 + at b (3) Exponential law:  = 0 + a[1-exp(-bt)] (4) Logarithmic law:  = 0 + a ln(1+bt) (5) Blackburn’s law:  = 0 + ac[1 - exp(-bt)]+c[1 - exp(-dt)] (6) where  is the strain, 0 is the instantaneous strain on dead weight, t is the elapsed time and a, b, c and d are the parameters characterizing the primary creep region. The a, b, c and d values were called the "scaling fac- tors".21 Figure 5 presents representative results of a re- gression analysis, for the creep at 923 K and 27.75 MPa. The exponential-law and power-law equations did not reproduce the experimental data. Compared to the loga- rithmic-law equation, the Blackburn equation provided a better representation of the experimental creep curve. Therefore, the short-term creep for the P91 heat-resistant steels at low stresses could be described with the Blackburn equation. In order for the creep mechanism of the P91 steels at low stresses to be studied, the instantaneous creep behavior was investigated. The specimens were loaded for 4.3 × 105 s until the strain reached 6 × 10–3. In this case, the dislocations could be expected to move. The stress increased or decreased at various levels that were consequently applied and the instantaneous strain was measured. J. ZHE et al.: SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 843–847 845 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: Strain/time creep curves for P91 at 923 K and various stresses Figure 4: Strain rate/strain creep curves at 923 K and various stresses Figure 6: Example of instantaneous elongation and contraction upon low stress changes at a high temperature Figure 5: Comparison of experimental data and predicted curves of strain versus time at 923 K and 27.75 MPa Figure 7 presents instantaneous-elongation and con- traction examples upon low changes in the stress of the P91 steels. In the figure, the – was the instantaneous strain under a stress decrease and the + was the in- stantaneous strain under a stress increase. The relationship between the stress increment || and the instantaneous strain |E| is presented in Fig- ure 7. The creep demonstrates a viscous behavior be- cause the absolute values of the instantaneous strain for the load increase were equal to the values for the load decrease. Two types of creep at low stress exist. One is creep controlled by the diffusion including the lattice diffusion creep (Nabarro-Herring type)22 at a high-T and grain boundary diffusion creep (Coble type)22 at an inter- mediate-T. The other is creep associated with dislocation movement including free flight motion (climb con- trolled) and viscous motion (glide controlled)23. When creep is controlled by the diffusion, the creep should be a non-viscous behavior. In this case, the absolute values of the instantaneous strain for a small-load increase should be apparently larger than the values for a small-load decrease. If creep is controlled by free flight motion of dislocation, plastic strain can occur instantaneously when the stress is increased by a small amount. There- fore, creep shows non-viscous behavior. When creep is controlled by the viscous glide of dislocations, instant- aneous plastic strain does not occur even if the applied stress is increase suddenly. Thus, creep shows viscous behavior. In this study, the creep demonstrates a viscous behavior, because non instantaneous strain is observed during the stress increase. It means creep may be con- trolled by the viscous glide of dislocations. Figure 8 gives an example of the anelastic backflow, observed for a high reduction in the stress during the transient-creep stage. Specifically, the helicoidal spring specimen was loaded at 27.75 MPa and 923 K for 120 h and consequently unloaded all the stress for 120 h. Two different parts of the creep strain existed: the anelastic strain was reversible, which is presented with solid double arrows, whereas the plastic strain was irreversible and it is presented with dashed double arrows. Therefore, the instantaneous strain that occurred under a sudden stress change included both elastic and anelastic compo- nents. 4 CONCLUSIONS The short-term creep behavior in the P91 heat-resis- tant steels was investigated using instantaneous-stress- change tests. The results were as follows: 1) The creep deformation consisted of the primary creep stages, but no secondary creep stage was observed. 2) The Blackburn equation could be utilized for the creep-curve description because it provided an im- proved representation of the experimental creep curve. 3) The viscous glide of dislocations might have been the dominant creep mechanism because the creep dis- played a viscous behavior. 4) The short-term creep at a low stress displayed an anelastic behavior. Acknowledgment This project was supported by the National Natural Science Foundation of China (Grant No. 51605330). 5 REFERENCES 1 K. Kimura, Y. Takahashi, ASME 2012 Pressure Vessels & Piping Division Conference, American Society of Mechanical Engineers, 18 (2012) 2, 177, doi:10.1016/0308-0161(93)90110-f 2 K. Maruyama, J. S. Lee, Creep & Fracture in High Temperature Components: Design & Life Assessment Issues, DEStech 3 H. G. Armaki, R. Chen, K. Maruyama, M. Igaras, Premature creep failure in strength enhanced high Cr ferritic steels caused by static recovery of tempered martensite lath structures, Materials Science and Engineering A, 527 (2010) 24–25, 6581–6588, doi:10.1016/ j. Msea. 2010. 07. 037 J. ZHE et al.: SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES ... 846 Materiali in tehnologije / Materials and technology 51 (2017) 5, 843–847 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 7: Instantaneous strain at a high temperature (T = 923 K) ver- sus the stress change during P91 steady-state creep Figure 8: Elongation versus time during a high stress reduction with- out an elastic strain 4 O. Prat, J. Garcia, D. Rojas, G. Sauthoff, G. Inden, The role of Laves phase on microstructure evolution and creep strength of novel 9%Cr heat resistant steels, Intermetallics, 32 (2013) 32, 362–372, doi:10.1016/j.internet.2012.08.016 5 L. Kyu-Ho, H. Sung-Min, S. Jae-Hyeok, S. Jin-Yoo, H. Joo-Youl, J. Woo-Sang, Effect of Nb addition on Z-phase formation and creep strength in high-Cr martensitic heat-resistant steels, Materials Characterization, 102 (2015) 2, 79–84, doi:10.1016/j.Matchar. 2014. 12. 028 6 X. Dong, Q. Z. Gao, C. Li, 9Cr-1.7W-0.4Mo-Co Ferritic Heat-Resis- tant Steel Isothermal Aging Microstructural Evolution, Publications, Lancaster, 85 (2008) 1–2, 372–379, doi:10.1016/j.ijpvp. 2007. 06. 006 7 Transactions of Materials and Heat Treatment, 36 (2015) 3, 81–86, doi:10. 1016/j.Msea.2013.09.033 8 K. Kimura, H. Kushima, K. Sawada, Long-term creep deformation property of modified 9Cr–1Mo steel, Materials Science and Engi- neering A, 511 (2009) 18, 58–63, doi:10.1016/j.msea.2008.04.095 9 K. Kimura, Y. Toda, H. Kushima, K. Sawada, Creep strength of high chromium steel with ferrite matrix, International Journal of Pressure Vessels and Piping, 87 (2010) 6, 282–288, doi:10.1016/j.ij2010.03. 016 10 E. M. Haney, F. Dalle, M. Sauzay, L. Vincent, I. Tournié, L. Allais, B. Fournier, Macroscopic results of long-term creep on a modified 9Cr–1Mo steel (T91), Materials Science and Engineering A, 510 (2009) 18, 99–103, doi: 10.1016/ j.msea. 2008. 04.099 11 K. Maruyama, S. Karashima, Theoretical Consideration of Measure- ment of Work-Hardening and Recovery Rates during High Tem- perature Creep, Trans. Japan Inst. Metals, 16 (1975) 11, 671–678, doi:10.2320/matertrans1960.16.671 12 K. Abe, H. Yoshinaga, S. Morozumi, A Method of Discerning Fric- tional Stress and Internal Stress by the Stress Relaxation Test, Journal of the Japan Institute of Metals and Materials, 18 (1997) 6, 479–487, doi:10.2320/matertrans1960.18.479 13 H. Oikawa, H. Sugawara, Instantaneous plastic strain associated with stress increments during the steady state creep of Al and Al - 5.5 At. Pct. Mg alloy, Scripta Metall., 12 (1978) 1, 85–89, doi:10.1016/ 0036-9748(78)90234-x 14 A. Magnin, Measurement of Very Low Creep Strains, Journal of Testing and Evaluation, 37 (2009) 1, 53–58, doi:10.1122/1.5550242 15 J. J. Shen, K. Ikeda, S. Hata , H. Nakashima, Instantaneous creep in face-centered cubic metals at ultra-low strain rates by a high- resolution strain measurement, Journal of Wuhan University of Technology (Materials Science Edition), 28 (2013) 6, 1096–1100, doi:10.1007/s11595-013-0826-y 16 S. Timoshenko, D. H. Yong, Elements of Strength of Materials, 5th ed., New York, Van Nostrand, 1968, 77–80 17 A. M. Wahi, Mechanical Springs, 2nd ed., New York, McGraw Hill, 1963, 229–230 18 M. R. Mitchel, R. E. Link, P. Marecek, Measurement of Very Low Creep Strain, A Review, J. Test. Eval., 37 (2009) 1, 53–58, doi:10. 15 20/jte101475 19 P. S. Zhang, J. J. Shen, H. Zhang, Short-Term Creep Behavior in P91 Heat-Resistant Steel at Low Stress, J. Material Science Forum, 850 (2016) 1, 922–926, doi:10.4028/www. scientific. net/ MSF.850.922 20 S. Holdsworth, Developments in the assessment of creep strain and ductility data, Materials at High Temperatures, 21 (2004) 1, 25–32, doi:10.3184/096034004782750041 21 R. P. Reed, N. J. Simon , R. P. Walsh, Creep of copper: 4–300 K, Materials Science and Engineering A, 147 (1991) 1, 23–32, doi:10.1016/0921-5093(91)90801-s 22 W. D. Nix, Effects of Grain Shape on Nabarro-Herring and Coble Creep Processes, Metals Forum, 4 (1981) 1, 38–43 23 H. Hiroyuki, N. Satoshi, K. Junichi, K. Akihiro, Creep deformation characterization of heat resistant steel by stress change test, International Journal of Pressure Vessels and Piping, 24 (2009) 9, 556–562, doi:10.1016/j.ijpvp.2008.06.003 24 K. Maruyama, S. Karashima, Theorethical Consideration of Measurement of Wor-Hardening and Recovery Rates during high temperature Creep, Trans. Japan Inst. Metals, 16 (1975) 11, 671–678, doi:10.2320/matertrans1960.16.671 J. ZHE et al.: SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 843–847 847 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 7 presents instantaneous-elongation and con- traction examples upon low changes in the stress of the P91 steels. In the figure, the – was the instantaneous strain under a stress decrease and the + was the in- stantaneous strain under a stress increase. The relationship between the stress increment || and the instantaneous strain |E| is presented in Fig- ure 7. The creep demonstrates a viscous behavior be- cause the absolute values of the instantaneous strain for the load increase were equal to the values for the load decrease. Two types of creep at low stress exist. One is creep controlled by the diffusion including the lattice diffusion creep (Nabarro-Herring type)22 at a high-T and grain boundary diffusion creep (Coble type)22 at an inter- mediate-T. The other is creep associated with dislocation movement including free flight motion (climb con- trolled) and viscous motion (glide controlled)23. When creep is controlled by the diffusion, the creep should be a non-viscous behavior. In this case, the absolute values of the instantaneous strain for a small-load increase should be apparently larger than the values for a small-load decrease. If creep is controlled by free flight motion of dislocation, plastic strain can occur instantaneously when the stress is increased by a small amount. There- fore, creep shows non-viscous behavior. When creep is controlled by the viscous glide of dislocations, instant- aneous plastic strain does not occur even if the applied stress is increase suddenly. Thus, creep shows viscous behavior. In this study, the creep demonstrates a viscous behavior, because non instantaneous strain is observed during the stress increase. It means creep may be con- trolled by the viscous glide of dislocations. Figure 8 gives an example of the anelastic backflow, observed for a high reduction in the stress during the transient-creep stage. Specifically, the helicoidal spring specimen was loaded at 27.75 MPa and 923 K for 120 h and consequently unloaded all the stress for 120 h. Two different parts of the creep strain existed: the anelastic strain was reversible, which is presented with solid double arrows, whereas the plastic strain was irreversible and it is presented with dashed double arrows. Therefore, the instantaneous strain that occurred under a sudden stress change included both elastic and anelastic compo- nents. 4 CONCLUSIONS The short-term creep behavior in the P91 heat-resis- tant steels was investigated using instantaneous-stress- change tests. The results were as follows: 1) The creep deformation consisted of the primary creep stages, but no secondary creep stage was observed. 2) The Blackburn equation could be utilized for the creep-curve description because it provided an im- proved representation of the experimental creep curve. 3) The viscous glide of dislocations might have been the dominant creep mechanism because the creep dis- played a viscous behavior. 4) The short-term creep at a low stress displayed an anelastic behavior. Acknowledgment This project was supported by the National Natural Science Foundation of China (Grant No. 51605330). 5 REFERENCES 1 K. Kimura, Y. Takahashi, ASME 2012 Pressure Vessels & Piping Division Conference, American Society of Mechanical Engineers, 18 (2012) 2, 177, doi:10.1016/0308-0161(93)90110-f 2 K. Maruyama, J. S. Lee, Creep & Fracture in High Temperature Components: Design & Life Assessment Issues, DEStech 3 H. G. Armaki, R. Chen, K. Maruyama, M. Igaras, Premature creep failure in strength enhanced high Cr ferritic steels caused by static recovery of tempered martensite lath structures, Materials Science and Engineering A, 527 (2010) 24–25, 6581–6588, doi:10.1016/ j. Msea. 2010. 07. 037 J. ZHE et al.: SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES ... 846 Materiali in tehnologije / Materials and technology 51 (2017) 5, 843–847 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 7: Instantaneous strain at a high temperature (T = 923 K) ver- sus the stress change during P91 steady-state creep Figure 8: Elongation versus time during a high stress reduction with- out an elastic strain 4 O. Prat, J. Garcia, D. Rojas, G. Sauthoff, G. Inden, The role of Laves phase on microstructure evolution and creep strength of novel 9%Cr heat resistant steels, Intermetallics, 32 (2013) 32, 362–372, doi:10.1016/j.internet.2012.08.016 5 L. Kyu-Ho, H. Sung-Min, S. Jae-Hyeok, S. Jin-Yoo, H. Joo-Youl, J. Woo-Sang, Effect of Nb addition on Z-phase formation and creep strength in high-Cr martensitic heat-resistant steels, Materials Characterization, 102 (2015) 2, 79–84, doi:10.1016/j.Matchar. 2014. 12. 028 6 X. Dong, Q. Z. Gao, C. Li, 9Cr-1.7W-0.4Mo-Co Ferritic Heat-Resis- tant Steel Isothermal Aging Microstructural Evolution, Publications, Lancaster, 85 (2008) 1–2, 372–379, doi:10.1016/j.ijpvp. 2007. 06. 006 7 Transactions of Materials and Heat Treatment, 36 (2015) 3, 81–86, doi:10. 1016/j.Msea.2013.09.033 8 K. Kimura, H. Kushima, K. Sawada, Long-term creep deformation property of modified 9Cr–1Mo steel, Materials Science and Engi- neering A, 511 (2009) 18, 58–63, doi:10.1016/j.msea.2008.04.095 9 K. Kimura, Y. Toda, H. Kushima, K. Sawada, Creep strength of high chromium steel with ferrite matrix, International Journal of Pressure Vessels and Piping, 87 (2010) 6, 282–288, doi:10.1016/j.ij2010.03. 016 10 E. M. Haney, F. Dalle, M. Sauzay, L. Vincent, I. Tournié, L. Allais, B. Fournier, Macroscopic results of long-term creep on a modified 9Cr–1Mo steel (T91), Materials Science and Engineering A, 510 (2009) 18, 99–103, doi: 10.1016/ j.msea. 2008. 04.099 11 K. Maruyama, S. Karashima, Theoretical Consideration of Measure- ment of Work-Hardening and Recovery Rates during High Tem- perature Creep, Trans. Japan Inst. Metals, 16 (1975) 11, 671–678, doi:10.2320/matertrans1960.16.671 12 K. Abe, H. Yoshinaga, S. Morozumi, A Method of Discerning Fric- tional Stress and Internal Stress by the Stress Relaxation Test, Journal of the Japan Institute of Metals and Materials, 18 (1997) 6, 479–487, doi:10.2320/matertrans1960.18.479 13 H. Oikawa, H. Sugawara, Instantaneous plastic strain associated with stress increments during the steady state creep of Al and Al - 5.5 At. Pct. Mg alloy, Scripta Metall., 12 (1978) 1, 85–89, doi:10.1016/ 0036-9748(78)90234-x 14 A. Magnin, Measurement of Very Low Creep Strains, Journal of Testing and Evaluation, 37 (2009) 1, 53–58, doi:10.1122/1.5550242 15 J. J. Shen, K. Ikeda, S. Hata , H. Nakashima, Instantaneous creep in face-centered cubic metals at ultra-low strain rates by a high- resolution strain measurement, Journal of Wuhan University of Technology (Materials Science Edition), 28 (2013) 6, 1096–1100, doi:10.1007/s11595-013-0826-y 16 S. Timoshenko, D. H. Yong, Elements of Strength of Materials, 5th ed., New York, Van Nostrand, 1968, 77–80 17 A. M. Wahi, Mechanical Springs, 2nd ed., New York, McGraw Hill, 1963, 229–230 18 M. R. Mitchel, R. E. Link, P. Marecek, Measurement of Very Low Creep Strain, A Review, J. Test. Eval., 37 (2009) 1, 53–58, doi:10. 15 20/jte101475 19 P. S. Zhang, J. J. Shen, H. Zhang, Short-Term Creep Behavior in P91 Heat-Resistant Steel at Low Stress, J. Material Science Forum, 850 (2016) 1, 922–926, doi:10.4028/www. scientific. net/ MSF.850.922 20 S. Holdsworth, Developments in the assessment of creep strain and ductility data, Materials at High Temperatures, 21 (2004) 1, 25–32, doi:10.3184/096034004782750041 21 R. P. Reed, N. J. Simon , R. P. Walsh, Creep of copper: 4–300 K, Materials Science and Engineering A, 147 (1991) 1, 23–32, doi:10.1016/0921-5093(91)90801-s 22 W. D. Nix, Effects of Grain Shape on Nabarro-Herring and Coble Creep Processes, Metals Forum, 4 (1981) 1, 38–43 23 H. Hiroyuki, N. Satoshi, K. Junichi, K. Akihiro, Creep deformation characterization of heat resistant steel by stress change test, International Journal of Pressure Vessels and Piping, 24 (2009) 9, 556–562, doi:10.1016/j.ijpvp.2008.06.003 24 K. Maruyama, S. Karashima, Theorethical Consideration of Measurement of Wor-Hardening and Recovery Rates during high temperature Creep, Trans. Japan Inst. Metals, 16 (1975) 11, 671–678, doi:10.2320/matertrans1960.16.671 J. ZHE et al.: SHORT-TERM CREEP OF P91 HEAT-RESISTANT STEELS AT LOW STRESSES ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 843–847 847 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS J. PALÁN et al.: EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ... 849–853 EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ON THE STRUCTURAL AND MECHANICAL PROPERTIES OF COMMERCIALLY PURE TITANIUM U^INEK PLASTI^NOSTI IN DEFORMACIJE PRI PODHLAJEVANJU NA STRUKTURNE IN MEHANSKE LASTNOSTI ^ISTEGA KOMERCIALNEGA TITANA Jan Palán1, Pavol [utta2, Tomá{ Kubina1, Mária Dománková3 1COMTES FHT, Prumyslova 995, 334 41 Dobrany, Czech Republic 2New Technologies – Research Centre Westbohemian Region, University of West Bohemia, 306 14 Plzen, Czech Republic 3Institute of Materials Science, Faculty of Materials Science and Technology Trnava, STU Bratislava, Bottova 25, 917 24 Trnava, Slovakia jpalan@comtesfht.cz Prejem rokopisa – received: 2016-10-25; sprejem za objavo – accepted for publication: 2017-01-23 doi:10.17222/mit.2016.307 SPD (severe plastic deformation) processing of materials provides a great potential associated with the enhancement of their properties by refining the initial grain structure. The present experiments involved mechanical working of commercial-purity titanium (Ti Grade 2) with the CONFORM SPD technique, which is one of the SPD methods, and with rotary swaging. The objective was to process the material at as low temperatures as possible in order to avoid softening processes and, therefore, to achieve the maximum strengthening through a microstructure refinement. Three passes through a CONFORM SPD machine were completed and the resulting ultimate strength was 673 MPa. The average grain size was 330 nm. The greatest improvement of the mechanical properties was achieved in the first pass. In the subsequent passes, the contributions were minor. The pro- cessing in the CONFORM SPD machine did not impair the ductility of the material. Subsequently, the wires were rotary swaged. The ultimate strength achieved was 1070 MPa. The response of the properties to this forming method was markedly different. The reason is that rotary swaging does not belong to SPD techniques. It causes rapid work hardening and reduces the ductility of the material. The workpiece was subsequently investigated with the aid of several techniques. Light and trans- mission electron microscopy and X-ray diffraction were employed for evaluating the grain size, distribution and orientation. Keywords: equal-channel angular pressing, CONFORM SPD technique, rotary swaging, titanium, extrusion Velika plasti~na deformacija (angl. SPD) pri obdelavi materialov omogo~a veliko mo`nosti, povezane z izbolj{anjem njihovih lastnosti z izbolj{avo izvorne strukture zrn. Pri~ujo~i preizkus je vklju~eval mehansko obdelavo komercialno dostopnega ~istega titana (Ti Grade 2) s CONFORM SPD-tehniko, ki je ena od SPD-metod, in s kovanjem. Namen je bil, da bi se material obdelalo pri ~im ni`jih temperaturah, kot je mogo~e, da bi se izognili procesom meh~anja in, da bi dosegli maksimalno krepitev s pre- ~i{~enjem strukture. Narejeni so bili trije nizi v CONFORM SPD-stroju in kon~na mo~ je dosegla 673 MPa. Povpre~na velikost zrn je bila 330 nm. Najve~je izbolj{ave mehanskih lastnosti so bile dose`ene v prvem nizu. V slede~ih nizih so bile le-te manj{e. Obdelava v CONFORM SPD-stroju ni {kodovala duktilnosti materiala. Posledi~no so bile `ice zvite. Kon~na mo~ je bila 1070 MPa. Odziv lastnosti na to metodo oblikovanja je bil izrazito druga~en. Razlog je v tem, da rotacijsko zvijanje ne sodi v tehnike SPD. Rotacijsko zvijanje povzro~i hitro utrjevanje in zmanj{uje duktilnost materiala. Vzorec je bil nato raziskan s pomo~jo ve~ tehnik. Za ovrednotenje velikosti, razporeditve in orientacije zrn, sta bili uporabljeni transmisijska elektronska mikroskopija in rentgenska difrakcija. Klju~ne besede: enakomerno kotno stiskanje, CONFORM SPD-tehnika, rotacijsko zvijanje, titan, iztiskanje 1 INTRODUCTION Severe plastic deformation (SPD) is a generic term for a group of methods which cause grain refinement in a material and produce equiaxed grains.1 The occurrence of an ultrafine structure or nanostructure is conditional on high hydrostatic pressure (P >1 GPa) and large shear deformation applied at relatively low temperatures. The temperatures of SPD processes should meet the condition of T(SPD) < 0.4 T(melting).1 Another aspect is the strain magnitude, defined as e(true strain) > 6–8. When the above conditions are met, the forming process leads to a high density of lattice defects, predominantly dislocations, and to the formation of subgrains, which reduces the stored energy.2 The relationship between the mechanical properties and the grain size is described with the Hall-Petch equation3,4. In the present study, commercially pure (CP) titanium Grade 2 was worked by means of the CONFORM SPD (CONSPD) method. This method uses continuous angular pressing through a die chamber of a modified design to produce ultrafine structures and nanostructures of the materials.5–7 The design modification reflects the principles of the ECAP method (equal-channel angular pressing), as described in studies.5–7 Continuous operation is the key advantage of this forming method. The processing of CP titanium Grade 2 with the CONFORM SDP machine was des- cribed in multiple papers.5,7 The main outcomes were the improved mechanical properties (ultimate tensile strength (UTS) = 698 MPa, 0.2 % offset yield stress (0.2 OYS) = 637 MPa), no significant decrease in the Materiali in tehnologije / Materials and technology 51 (2017) 5, 849–853 849 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 67.017:620.1:536.421.48 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)849(2017) J. PALÁN et al.: EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ... 849–853 EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ON THE STRUCTURAL AND MECHANICAL PROPERTIES OF COMMERCIALLY PURE TITANIUM U^INEK PLASTI^NOSTI IN DEFORMACIJE PRI PODHLAJEVANJU NA STRUKTURNE IN MEHANSKE LASTNOSTI ^ISTEGA KOMERCIALNEGA TITANA Jan Palán1, Pavol [utta2, Tomá{ Kubina1, Mária Dománková3 1COMTES FHT, Prumyslova 995, 334 41 Dobrany, Czech Republic 2New Technologies – Research Centre Westbohemian Region, University of West Bohemia, 306 14 Plzen, Czech Republic 3Institute of Materials Science, Faculty of Materials Science and Technology Trnava, STU Bratislava, Bottova 25, 917 24 Trnava, Slovakia jpalan@comtesfht.cz Prejem rokopisa – received: 2016-10-25; sprejem za objavo – accepted for publication: 2017-01-23 doi:10.17222/mit.2016.307 SPD (severe plastic deformation) processing of materials provides a great potential associated with the enhancement of their properties by refining the initial grain structure. The present experiments involved mechanical working of commercial-purity titanium (Ti Grade 2) with the CONFORM SPD technique, which is one of the SPD methods, and with rotary swaging. The objective was to process the material at as low temperatures as possible in order to avoid softening processes and, therefore, to achieve the maximum strengthening through a microstructure refinement. Three passes through a CONFORM SPD machine were completed and the resulting ultimate strength was 673 MPa. The average grain size was 330 nm. The greatest improvement of the mechanical properties was achieved in the first pass. In the subsequent passes, the contributions were minor. The pro- cessing in the CONFORM SPD machine did not impair the ductility of the material. Subsequently, the wires were rotary swaged. The ultimate strength achieved was 1070 MPa. The response of the properties to this forming method was markedly different. The reason is that rotary swaging does not belong to SPD techniques. It causes rapid work hardening and reduces the ductility of the material. The workpiece was subsequently investigated with the aid of several techniques. Light and trans- mission electron microscopy and X-ray diffraction were employed for evaluating the grain size, distribution and orientation. Keywords: equal-channel angular pressing, CONFORM SPD technique, rotary swaging, titanium, extrusion Velika plasti~na deformacija (angl. SPD) pri obdelavi materialov omogo~a veliko mo`nosti, povezane z izbolj{anjem njihovih lastnosti z izbolj{avo izvorne strukture zrn. Pri~ujo~i preizkus je vklju~eval mehansko obdelavo komercialno dostopnega ~istega titana (Ti Grade 2) s CONFORM SPD-tehniko, ki je ena od SPD-metod, in s kovanjem. Namen je bil, da bi se material obdelalo pri ~im ni`jih temperaturah, kot je mogo~e, da bi se izognili procesom meh~anja in, da bi dosegli maksimalno krepitev s pre- ~i{~enjem strukture. Narejeni so bili trije nizi v CONFORM SPD-stroju in kon~na mo~ je dosegla 673 MPa. Povpre~na velikost zrn je bila 330 nm. Najve~je izbolj{ave mehanskih lastnosti so bile dose`ene v prvem nizu. V slede~ih nizih so bile le-te manj{e. Obdelava v CONFORM SPD-stroju ni {kodovala duktilnosti materiala. Posledi~no so bile `ice zvite. Kon~na mo~ je bila 1070 MPa. Odziv lastnosti na to metodo oblikovanja je bil izrazito druga~en. Razlog je v tem, da rotacijsko zvijanje ne sodi v tehnike SPD. Rotacijsko zvijanje povzro~i hitro utrjevanje in zmanj{uje duktilnost materiala. Vzorec je bil nato raziskan s pomo~jo ve~ tehnik. Za ovrednotenje velikosti, razporeditve in orientacije zrn, sta bili uporabljeni transmisijska elektronska mikroskopija in rentgenska difrakcija. Klju~ne besede: enakomerno kotno stiskanje, CONFORM SPD-tehnika, rotacijsko zvijanje, titan, iztiskanje 1 INTRODUCTION Severe plastic deformation (SPD) is a generic term for a group of methods which cause grain refinement in a material and produce equiaxed grains.1 The occurrence of an ultrafine structure or nanostructure is conditional on high hydrostatic pressure (P >1 GPa) and large shear deformation applied at relatively low temperatures. The temperatures of SPD processes should meet the condition of T(SPD) < 0.4 T(melting).1 Another aspect is the strain magnitude, defined as e(true strain) > 6–8. When the above conditions are met, the forming process leads to a high density of lattice defects, predominantly dislocations, and to the formation of subgrains, which reduces the stored energy.2 The relationship between the mechanical properties and the grain size is described with the Hall-Petch equation3,4. In the present study, commercially pure (CP) titanium Grade 2 was worked by means of the CONFORM SPD (CONSPD) method. This method uses continuous angular pressing through a die chamber of a modified design to produce ultrafine structures and nanostructures of the materials.5–7 The design modification reflects the principles of the ECAP method (equal-channel angular pressing), as described in studies.5–7 Continuous operation is the key advantage of this forming method. The processing of CP titanium Grade 2 with the CONFORM SDP machine was des- cribed in multiple papers.5,7 The main outcomes were the improved mechanical properties (ultimate tensile strength (UTS) = 698 MPa, 0.2 % offset yield stress (0.2 OYS) = 637 MPa), no significant decrease in the Materiali in tehnologije / Materials and technology 51 (2017) 5, 849–853 849 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 67.017:620.1:536.421.48 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)849(2017) elongation, and an equiaxed ultrafine-grained (UFG) microstructure.7 The present study proposes a process route which involves CONFORM SPD forming and rotary swaging (RS) in order to enhance the mechanical properties of a workpiece.8–10 The effects of severe plastic deformation and work hardening are thus combined. The proposed technology allows the mechanical properties of pure titanium to be improved dramatically. The ultrafine- grained CP Ti Grade 2 has a great potential because its mechanical properties are comparable to the Ti6Al4V alloy, which is mostly used in medicine. Recently, it was suggested that Al and V might be toxic for the human body.11,12 2 EXPERIMENTAL PART The material under investigation was CP titanium Grade 2 with the chemical composition given in Table 1. The composition was measured by means of the Bruker Q4 Tasman optical emission spectrometer and the Bruker G8 Galileo gas analyser. The diameter of Ti rods was 10 mm. Table 1: Chemical composition of feedstock in weight percent Fe O C H N Ti 0.046 0.12 0.023 0.0026 0.0076 99.822 The processing was carried out in a CONFORM SPD machine, type 315i with a modified die chamber (Figure 1), and in a HMP R4-4 rotary-swaging machine. During the CONFORM SPD process, the temperature was 220 °C, the wheel speed was 0.5 min–1, and the angle of the die chamber was 90°. Three passes through the CONFORM SPD machine were completed. The pro- duct’s cross-section was identical to that of the feed- stock. Rotary swaging, the subsequent process, was carried out at ambient temperature. In this operation, the cross-section area was reduced by 20 % in each pass. The total area reduction was 90 %. For the purpose of observation with a transmission electron microscope (TEM), thin foils were prepared with the final electrolytic thinning in a Tenupol 5 device, using a solution of 300 mL CH3OH + 175 mL 2-butanol + 30 mL HClO4 at –10 °C and a voltage of 40 V. The TEM analysis was performed with a JEOL 200CX instrument with an acceleration voltage of 200 kV. Selec- tive electron diffraction was used for the determination of the phases. The grain size was measured using the linear intercept method. The preferred orientation of crystallites (texture) was analysed with an automatic powder diffractometer X’Pert-Pro equipped with an ultra-fast semiconductor position-sensitive detector Pixcel. Cu-K1 radiation ( = 0.154056 nm) was used. The texture was characte- rized from the radial X-ray diffraction patterns using the Harris (1952) texture index, Equation (1): T n I /R I /R i i i j j j n= ⋅ = ∑ 1 (1) where n is the number of the reflections investigated, Ii is the observed intensity and Ri is the corresponding intensity for the sample with randomly oriented crystallites. The Ri values were obtained from the ICDD PDF standard diffraction data file (reference code 00-005-0682). The texture was analysed in the direction of the material flow, both after CONFORM SPD (longitudinal direction), and after CONFORM SPD + rotary swaging. 3 MICROSTRUCTURE The mean grain size in the initial condition of the feedstock was 5390 nm. After the first pass, the mean grain size was daverage ~ 320±35 nm in the transverse direction. After the first pass, the microstructure was equiaxed with a non-uniform dislocation density (Figure 2a). Some locations exhibited a preferential orientation (Figure 2b). After the second pass (Figure 2c), the mean grain size was daverage ~ 250±25 nm in transverse direction. The dislocation density was still non-uniform after the second pass. After the third pass, the mean grain size was daverage ~ 330±30 nm (Figure 3d): the grain was larger than after the second pass. This increase can be attributed, in part, to the non-uniform deformation and to the high surface activity of the UFG microstructure. High surface activity and dislocation density (strain mag- nitude) lower the temperatures of softening processes. A certain grain growth can be expected to occur during the forming due to deformation heat and the heat retained in the die chamber (220 °C). The grain growth upon multi- ple passes through the CONFORM SPD machine was already reported in 9. Figure 3a is a micrograph of the structure in the transverse direction after three passes through the CON- FORM SPD machine and after a 35 % area reduction J. PALÁN et al.: EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ... 850 Materiali in tehnologije / Materials and technology 51 (2017) 5, 849–853 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 1: Schematic illustration of the CONFORM SPD process with rotary swaging (RS). The mean grain size in the transverse direction is daverage ~ 300±150 nm. The large standard deviation is due to a large variation in the grain size. Figure 3b is a micrograph taken in the longitudinal direction after CONFORM SPD and RS. The grains are elongated and aligned in the direction of forming due to the nature of the deformation introduced by rotary swaging. Unlike in CONFORM SPD, the grains become elongated instead of being refined or converted into new polyhedral grains. After the additional reduction due to rotary swaging (3 passes through CONFORM SPD + the total area reduction of 50 % by RS), it became im- possible to determine the grain size due to the high dislocation density, as seen in Figure 3c. The summary of the average grain sizes is given in Table 2. Table 2: Average grain size in the transverse and longitudinal sections obtained from TEM images Condition Grain size (nm) Transverse Longitudinal as received 5390±20 CON SPD – 1 pass 320±35 340±30 CONSPD – 2 passes 250±25 310±30 CONSPD – 3 passes 330±30 420±30 CON SPD – 3 passes + RS (35 %) 300±150 220±50 CONSPD – 3 passes + RS (50 %) Not possible to evaluate due to high dislocation density J. PALÁN et al.: EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 849–853 851 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: TEM images: a) CON SPD – 3 passes + RS (area reduction by 35 %) in the transverse direction, b) CON SPD – 3 passes + RS (35 %) in the longitudinal direction, c) CON SPD – 3 passes + RS (50 %) in the transverse direction Figure 2: TEM images: a) CON SPD – 1 pass (transverse), b) CON SPD – 1 pass (longitudinal), c) CON SPD – 2 passes (transverse), d) CON SPD – 3 passes (transverse) elongation, and an equiaxed ultrafine-grained (UFG) microstructure.7 The present study proposes a process route which involves CONFORM SPD forming and rotary swaging (RS) in order to enhance the mechanical properties of a workpiece.8–10 The effects of severe plastic deformation and work hardening are thus combined. The proposed technology allows the mechanical properties of pure titanium to be improved dramatically. The ultrafine- grained CP Ti Grade 2 has a great potential because its mechanical properties are comparable to the Ti6Al4V alloy, which is mostly used in medicine. Recently, it was suggested that Al and V might be toxic for the human body.11,12 2 EXPERIMENTAL PART The material under investigation was CP titanium Grade 2 with the chemical composition given in Table 1. The composition was measured by means of the Bruker Q4 Tasman optical emission spectrometer and the Bruker G8 Galileo gas analyser. The diameter of Ti rods was 10 mm. Table 1: Chemical composition of feedstock in weight percent Fe O C H N Ti 0.046 0.12 0.023 0.0026 0.0076 99.822 The processing was carried out in a CONFORM SPD machine, type 315i with a modified die chamber (Figure 1), and in a HMP R4-4 rotary-swaging machine. During the CONFORM SPD process, the temperature was 220 °C, the wheel speed was 0.5 min–1, and the angle of the die chamber was 90°. Three passes through the CONFORM SPD machine were completed. The pro- duct’s cross-section was identical to that of the feed- stock. Rotary swaging, the subsequent process, was carried out at ambient temperature. In this operation, the cross-section area was reduced by 20 % in each pass. The total area reduction was 90 %. For the purpose of observation with a transmission electron microscope (TEM), thin foils were prepared with the final electrolytic thinning in a Tenupol 5 device, using a solution of 300 mL CH3OH + 175 mL 2-butanol + 30 mL HClO4 at –10 °C and a voltage of 40 V. The TEM analysis was performed with a JEOL 200CX instrument with an acceleration voltage of 200 kV. Selec- tive electron diffraction was used for the determination of the phases. The grain size was measured using the linear intercept method. The preferred orientation of crystallites (texture) was analysed with an automatic powder diffractometer X’Pert-Pro equipped with an ultra-fast semiconductor position-sensitive detector Pixcel. Cu-K1 radiation ( = 0.154056 nm) was used. The texture was characte- rized from the radial X-ray diffraction patterns using the Harris (1952) texture index, Equation (1): T n I /R I /R i i i j j j n= ⋅ = ∑ 1 (1) where n is the number of the reflections investigated, Ii is the observed intensity and Ri is the corresponding intensity for the sample with randomly oriented crystallites. The Ri values were obtained from the ICDD PDF standard diffraction data file (reference code 00-005-0682). The texture was analysed in the direction of the material flow, both after CONFORM SPD (longitudinal direction), and after CONFORM SPD + rotary swaging. 3 MICROSTRUCTURE The mean grain size in the initial condition of the feedstock was 5390 nm. After the first pass, the mean grain size was daverage ~ 320±35 nm in the transverse direction. After the first pass, the microstructure was equiaxed with a non-uniform dislocation density (Figure 2a). Some locations exhibited a preferential orientation (Figure 2b). After the second pass (Figure 2c), the mean grain size was daverage ~ 250±25 nm in transverse direction. The dislocation density was still non-uniform after the second pass. After the third pass, the mean grain size was daverage ~ 330±30 nm (Figure 3d): the grain was larger than after the second pass. This increase can be attributed, in part, to the non-uniform deformation and to the high surface activity of the UFG microstructure. High surface activity and dislocation density (strain mag- nitude) lower the temperatures of softening processes. A certain grain growth can be expected to occur during the forming due to deformation heat and the heat retained in the die chamber (220 °C). The grain growth upon multi- ple passes through the CONFORM SPD machine was already reported in 9. Figure 3a is a micrograph of the structure in the transverse direction after three passes through the CON- FORM SPD machine and after a 35 % area reduction J. PALÁN et al.: EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ... 850 Materiali in tehnologije / Materials and technology 51 (2017) 5, 849–853 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 1: Schematic illustration of the CONFORM SPD process with rotary swaging (RS). The mean grain size in the transverse direction is daverage ~ 300±150 nm. The large standard deviation is due to a large variation in the grain size. Figure 3b is a micrograph taken in the longitudinal direction after CONFORM SPD and RS. The grains are elongated and aligned in the direction of forming due to the nature of the deformation introduced by rotary swaging. Unlike in CONFORM SPD, the grains become elongated instead of being refined or converted into new polyhedral grains. After the additional reduction due to rotary swaging (3 passes through CONFORM SPD + the total area reduction of 50 % by RS), it became im- possible to determine the grain size due to the high dislocation density, as seen in Figure 3c. The summary of the average grain sizes is given in Table 2. Table 2: Average grain size in the transverse and longitudinal sections obtained from TEM images Condition Grain size (nm) Transverse Longitudinal as received 5390±20 CON SPD – 1 pass 320±35 340±30 CONSPD – 2 passes 250±25 310±30 CONSPD – 3 passes 330±30 420±30 CON SPD – 3 passes + RS (35 %) 300±150 220±50 CONSPD – 3 passes + RS (50 %) Not possible to evaluate due to high dislocation density J. PALÁN et al.: EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 849–853 851 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: TEM images: a) CON SPD – 3 passes + RS (area reduction by 35 %) in the transverse direction, b) CON SPD – 3 passes + RS (35 %) in the longitudinal direction, c) CON SPD – 3 passes + RS (50 %) in the transverse direction Figure 2: TEM images: a) CON SPD – 1 pass (transverse), b) CON SPD – 1 pass (longitudinal), c) CON SPD – 2 passes (transverse), d) CON SPD – 3 passes (transverse) 4 TEXTURE EVALUATION The preferred orientation of crystallites expressed with the texture index Ti (Equation 1, Table 3) is, in all the cases, in the [001] direction perpendicular to the sample surface analysed. The highest value was found for CONSPD – 3 passes + RS (75 %) titanium sample, i.e., the sample that was processed using the CONFORM SPD device and rotary swaging (basal texture, Figure 5). An increased value of the texture index can be seen in the initial state, due to the previous wire-drawing process (Figure 4). The texture index is lower for the samples that were processed with the CONFORM SPD device. Texture indices for the [101] direction are listed for the sake of comparison. Nevertheless, all their values are lower than 1. It means that no preferred orientation of crystallites in the [101] direction is present, as opposed to the sample surface. Table 3: Texture evaluation using X-ray diffraction Plane (h k l) Texture index as received CONSPD – 1pass CONSPD – 2 passes Ti (002) 3.13 2.49 2.67 Ti (101) 0.66 0.69 0.86 CONSPD – 3 passes CONSPD – 3 passes + RS (75 %) Standard Ti (002) 2.37 3.48 1 Ti (101) 0.97 0.24 1 4 MECHANICAL PROPERTIES Table 4 indicates that the largest increase in the mechanical properties was obtained during the first pass through CONFORM SPD. The subsequent passes led to smaller increments. It is important to note that the increase in the ultimate strength and yield stress upon CONFORM SPD is not offset by a decrease in the ductility. The improvement in the mechanical properties is mainly due to the grain refinement and the increased dislocation density, which impede the dislocation move- ment. A further increase in the mechanical properties was obtained with rotary swaging applied after the three passes through the CONFORM SPD machine. In compa- rison with the CONFORM SPD process, rotary swaging leads to a decrease in the ductility as a result of work hardening. Table 4: Mechanical properties in the as-received condition, after CONFORM SPD processing and after RS Condition 0.2 OYS UTS A5 RA MPa MPa % % as received 370 480 25 52 CONSPD – 1 pass 540 580 23 62 CONSPD – 2 passes 560 600 23 62 CONSPD – 3 passes 570 623 20 64 CONSPD – 3 passes + RS (50 %) 830 885 13 54 CONSPD – 3 passes + RS (75 %) 930 1000 12 57 CONSPD – 3 passes + RS (85%) 950 1070 12 58 RS only (85 %) 780 850 12 50 5 DISCUSSION The CONFORM SPD processing led to higher ulti- mate strengths and yield stresses without a reduced ductility. This is characteristic of the process of the for- mation of a UFG equiaxed structure shown in Figure 2. The grain size and distortion are non-uniform due to non-uniform deformation during the forming process. The preferred orientation of crystallites in the [001] direction perpendicular to the surfaces of the investigated samples is not too significant, except for the sample which was processed with CONFORM SPD and the rotary swaging machine where the highest value of the texture index was observed (the basal texture). Major increases in the ultimate strength and yield stress were found after the first pass through the CON- J. PALÁN et al.: EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ... 852 Materiali in tehnologije / Materials and technology 51 (2017) 5, 849–853 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: Diffraction radial profile of the specimen with the strongest texture: CONSPD – 3 passes + rotary swaging (75 %) Figure 4: Diffraction radial profile of the as-received specimen FORM SPD machine, whereas the increments from the subsequent passes were small. This is closely related to the grain size, which did not decrease during the sub- sequent passes. It is possible to say that the additional deformation causes the subgrains to rotate into high- angle grain boundaries, typically with an equiaxed shape of the grains.5,7,9 The subsequent rotary swaging led to a further increase in the mechanical properties, but the strengthening mechanism was different this time. Instead of the formation of equiaxed grains, the grains became elongated in the direction of forming, with a much higher dislocation density (Figure 3).10,11 When com- pared to the CONFORM SPD route, rotary swaging led to a reduced ductility as a consequence of work harden- ing. 6 CONCLUSION This study involved the processing of Ti Grade 2 using CONFORM SPD and rotary swaging with the goal of improving its mechanical properties. The results and findings are described below: • The CONFORM SPD processing substantially refined the grain from its initial size of 5390 nm to 350 nm. The subsequent rotary swaging did not provide for a further grain refinement but led to a grain elongation and to a higher dislocation density. • The processing led to the basal texture. The most significant form of texture was found in the specimen processed with CONFORM SPD and rotary swaging. • Three passes through the CONFORM SPD machine led to a strength of 623 MPa and a yield stress of 570 MPa without any notable decrease in the elon- gation. • After CONFORM SPD, the specimens were rotary swaged, which brought their strength to 1070 MPa and the yield stress to 950 MPa. In the initial con- dition, the ultimate strength was 480 MPa and the yield stress was 370 MPa. Acknowledgment These results were obtained under the project entitled "Development of West-Bohemian Centre of Materials and Metallurgy" No. lo1412, financed by the Ministry of Education of the Czech Republic. 7 REFERENCES 1 R. Z. Valiev, R. K. Islamgaliev, I. V. Alexandrov, Bulk nanostruc- tured materials from severe plastic deformation, Progress in Materials Science, 45 (2000) 2, doi:10.1016/S0079-6425(99)00007-9 2 N. J. Petch, The cleavage strength of polycrystals, J. Iron Steel Ins, 174 (1953), 25–28 3 E. O. Hall, The Deformation and Ageing of Mild Steel: III Dis- cussion of Results, Proc. Phys. Soc. London, 64 (1951), 747–753, doi:10.1088/0370-1301/64/9/303 4 A. Mishra, B. Kad, F. Gregori, M. Meyers, Microstructural evolution in copper subjected to severe plastic deformation, Acta Materialia, 55 (2007), 2, doi:10.1016/j.actamat.2006.07.008 5 M. Duchek, T. Kubina, J. Hodek, J. Dlouhy, Development of the pro- duction of ultrafine-grained titanium with the conform equipment, Mater. Tehnol., 47 (2013), 4 6 G. J. Raab, R. Z. Valiev, T. C. Lowe, Y. T. Zhu, Continuous pro- cessing of ultrafine grained Al by ECAP–Conform, Materials Science and Engineering, 382 (2004) 1/2, doi:10.1016/j.msea. 2004.04.021 7 T. Kubina, J. Dlouhý, M. Kover, Preparation and thermal stability of ultra-fine and nano-grained commercially pure titanium wires using CONFORM equipment, Mater. Tehnol., 49 (2015), 2, doi:10.17222/ mit.2013.226 8 L. Ostrovska, L. Vistejnova, J. Dzugan, P. Slama, T. Kubina, E. Ukraintsev, D. Kubies, M. Kralickova, M. Kalbacova, Biological evaluation of ultra-fine titanium with improved mechanical strength for dental implant engineering, J. Mater. Sci., 23 (2015), doi:10.1007/s10853-015-9619-3 9 M. Zemko, T. Kubina, J. Dlouhý, J. Hodek, Technological aspects of preparation of nanostructured titanium wire using a CONFORM machine, IOP Conference Series: Materials Science and Engineering, 63 (2014), 2, doi:10.1088/1757-899X/63/1/012049 10 D. V. Gunderov, A. V. Polyakov, I. P. Semenova, Evolution of microstructure, macrotexture and mechanical properties of commercially pure Ti during ECAP-conform processing and drawing, Materials Science and Engineering, 562 (2013), 128–136, doi:10.1016/j.msea.2012.11.007 11 I. P. Semenova, R. Z. Valiev, E. B. Yakushima, G. H. Salimgareeva, T. C. Lowe, Strength and fatigue properties enhancement in ultra- fine-grained Ti produced by severe plastic deformation, Journal of Materials Science, 43 (2008), 23–24, doi:10.1007/s10853-008- 2984-4 12 H. Alkhazraji, E. El-Danaf, M. Wollmann, L. Wagner, Enhanced Fatigue Strength of Commercially Pure Ti Processed by Rotary Swaging, Advances in Materials Science and Engineering, (2015), 1–12, doi:10.1155/2015/301837 J. PALÁN et al.: EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 849–853 853 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 4 TEXTURE EVALUATION The preferred orientation of crystallites expressed with the texture index Ti (Equation 1, Table 3) is, in all the cases, in the [001] direction perpendicular to the sample surface analysed. The highest value was found for CONSPD – 3 passes + RS (75 %) titanium sample, i.e., the sample that was processed using the CONFORM SPD device and rotary swaging (basal texture, Figure 5). An increased value of the texture index can be seen in the initial state, due to the previous wire-drawing process (Figure 4). The texture index is lower for the samples that were processed with the CONFORM SPD device. Texture indices for the [101] direction are listed for the sake of comparison. Nevertheless, all their values are lower than 1. It means that no preferred orientation of crystallites in the [101] direction is present, as opposed to the sample surface. Table 3: Texture evaluation using X-ray diffraction Plane (h k l) Texture index as received CONSPD – 1pass CONSPD – 2 passes Ti (002) 3.13 2.49 2.67 Ti (101) 0.66 0.69 0.86 CONSPD – 3 passes CONSPD – 3 passes + RS (75 %) Standard Ti (002) 2.37 3.48 1 Ti (101) 0.97 0.24 1 4 MECHANICAL PROPERTIES Table 4 indicates that the largest increase in the mechanical properties was obtained during the first pass through CONFORM SPD. The subsequent passes led to smaller increments. It is important to note that the increase in the ultimate strength and yield stress upon CONFORM SPD is not offset by a decrease in the ductility. The improvement in the mechanical properties is mainly due to the grain refinement and the increased dislocation density, which impede the dislocation move- ment. A further increase in the mechanical properties was obtained with rotary swaging applied after the three passes through the CONFORM SPD machine. In compa- rison with the CONFORM SPD process, rotary swaging leads to a decrease in the ductility as a result of work hardening. Table 4: Mechanical properties in the as-received condition, after CONFORM SPD processing and after RS Condition 0.2 OYS UTS A5 RA MPa MPa % % as received 370 480 25 52 CONSPD – 1 pass 540 580 23 62 CONSPD – 2 passes 560 600 23 62 CONSPD – 3 passes 570 623 20 64 CONSPD – 3 passes + RS (50 %) 830 885 13 54 CONSPD – 3 passes + RS (75 %) 930 1000 12 57 CONSPD – 3 passes + RS (85%) 950 1070 12 58 RS only (85 %) 780 850 12 50 5 DISCUSSION The CONFORM SPD processing led to higher ulti- mate strengths and yield stresses without a reduced ductility. This is characteristic of the process of the for- mation of a UFG equiaxed structure shown in Figure 2. The grain size and distortion are non-uniform due to non-uniform deformation during the forming process. The preferred orientation of crystallites in the [001] direction perpendicular to the surfaces of the investigated samples is not too significant, except for the sample which was processed with CONFORM SPD and the rotary swaging machine where the highest value of the texture index was observed (the basal texture). Major increases in the ultimate strength and yield stress were found after the first pass through the CON- J. PALÁN et al.: EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ... 852 Materiali in tehnologije / Materials and technology 51 (2017) 5, 849–853 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: Diffraction radial profile of the specimen with the strongest texture: CONSPD – 3 passes + rotary swaging (75 %) Figure 4: Diffraction radial profile of the as-received specimen FORM SPD machine, whereas the increments from the subsequent passes were small. This is closely related to the grain size, which did not decrease during the sub- sequent passes. It is possible to say that the additional deformation causes the subgrains to rotate into high- angle grain boundaries, typically with an equiaxed shape of the grains.5,7,9 The subsequent rotary swaging led to a further increase in the mechanical properties, but the strengthening mechanism was different this time. Instead of the formation of equiaxed grains, the grains became elongated in the direction of forming, with a much higher dislocation density (Figure 3).10,11 When com- pared to the CONFORM SPD route, rotary swaging led to a reduced ductility as a consequence of work harden- ing. 6 CONCLUSION This study involved the processing of Ti Grade 2 using CONFORM SPD and rotary swaging with the goal of improving its mechanical properties. The results and findings are described below: • The CONFORM SPD processing substantially refined the grain from its initial size of 5390 nm to 350 nm. The subsequent rotary swaging did not provide for a further grain refinement but led to a grain elongation and to a higher dislocation density. • The processing led to the basal texture. The most significant form of texture was found in the specimen processed with CONFORM SPD and rotary swaging. • Three passes through the CONFORM SPD machine led to a strength of 623 MPa and a yield stress of 570 MPa without any notable decrease in the elon- gation. • After CONFORM SPD, the specimens were rotary swaged, which brought their strength to 1070 MPa and the yield stress to 950 MPa. In the initial con- dition, the ultimate strength was 480 MPa and the yield stress was 370 MPa. Acknowledgment These results were obtained under the project entitled "Development of West-Bohemian Centre of Materials and Metallurgy" No. lo1412, financed by the Ministry of Education of the Czech Republic. 7 REFERENCES 1 R. Z. Valiev, R. K. Islamgaliev, I. V. Alexandrov, Bulk nanostruc- tured materials from severe plastic deformation, Progress in Materials Science, 45 (2000) 2, doi:10.1016/S0079-6425(99)00007-9 2 N. J. Petch, The cleavage strength of polycrystals, J. Iron Steel Ins, 174 (1953), 25–28 3 E. O. Hall, The Deformation and Ageing of Mild Steel: III Dis- cussion of Results, Proc. Phys. Soc. London, 64 (1951), 747–753, doi:10.1088/0370-1301/64/9/303 4 A. Mishra, B. Kad, F. Gregori, M. Meyers, Microstructural evolution in copper subjected to severe plastic deformation, Acta Materialia, 55 (2007), 2, doi:10.1016/j.actamat.2006.07.008 5 M. Duchek, T. Kubina, J. Hodek, J. Dlouhy, Development of the pro- duction of ultrafine-grained titanium with the conform equipment, Mater. Tehnol., 47 (2013), 4 6 G. J. Raab, R. Z. Valiev, T. C. Lowe, Y. T. Zhu, Continuous pro- cessing of ultrafine grained Al by ECAP–Conform, Materials Science and Engineering, 382 (2004) 1/2, doi:10.1016/j.msea. 2004.04.021 7 T. Kubina, J. Dlouhý, M. Kover, Preparation and thermal stability of ultra-fine and nano-grained commercially pure titanium wires using CONFORM equipment, Mater. Tehnol., 49 (2015), 2, doi:10.17222/ mit.2013.226 8 L. Ostrovska, L. Vistejnova, J. Dzugan, P. Slama, T. Kubina, E. Ukraintsev, D. Kubies, M. Kralickova, M. Kalbacova, Biological evaluation of ultra-fine titanium with improved mechanical strength for dental implant engineering, J. Mater. Sci., 23 (2015), doi:10.1007/s10853-015-9619-3 9 M. Zemko, T. Kubina, J. Dlouhý, J. Hodek, Technological aspects of preparation of nanostructured titanium wire using a CONFORM machine, IOP Conference Series: Materials Science and Engineering, 63 (2014), 2, doi:10.1088/1757-899X/63/1/012049 10 D. V. Gunderov, A. V. Polyakov, I. P. Semenova, Evolution of microstructure, macrotexture and mechanical properties of commercially pure Ti during ECAP-conform processing and drawing, Materials Science and Engineering, 562 (2013), 128–136, doi:10.1016/j.msea.2012.11.007 11 I. P. Semenova, R. Z. Valiev, E. B. Yakushima, G. H. Salimgareeva, T. C. Lowe, Strength and fatigue properties enhancement in ultra- fine-grained Ti produced by severe plastic deformation, Journal of Materials Science, 43 (2008), 23–24, doi:10.1007/s10853-008- 2984-4 12 H. Alkhazraji, E. El-Danaf, M. Wollmann, L. Wagner, Enhanced Fatigue Strength of Commercially Pure Ti Processed by Rotary Swaging, Advances in Materials Science and Engineering, (2015), 1–12, doi:10.1155/2015/301837 J. PALÁN et al.: EFFECT OF SEVERE PLASTIC AND HEAVY COLD DEFORMATION ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 849–853 853 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS A. G. ILLARIONOV et al.: EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE ... 855–859 EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE AND PROPERTIES OF WELD JOINTS OF A TWO-PHASE TITANIUM ALLOY U^INEK MIKROLEGIRANJA ITRIJA IN CIRKONIJA NA STRUKTURO IN LASTNOSTI NA SPOJE ZAVROV DVOFAZNE ZLITINE TITANA Anatoly Illarionov, Artemy Popov, Svetlana Illarionova, Dmitry Gadeev Ural Federal University, 620034, 28 Mira str., Yekaterinburg, Russia d.v.gadeev@urfu.ru Prejem rokopisa – received: 2016-11-10; sprejem za objavo – accepted for publication: 2017-04-19 doi:10.17222/mit.2016.317 The effect of microalloying of the Ti-4.8Al-1.2Mo-2.6V-0.6Cr–0.25Fe titanium alloy with yttrium and zirconium on the phase composition, structure and mechanical properties of welded sheets was studied using light and transmission electron microscopy, X-ray diffraction analysis and microindentation-hardness testing. A differential thermal analysis was employed to model the welding thermal cycle. It was found that alloying with yttrium led to the precipitation of Y2O3 oxide particles resulting in refining the microstructure of the alloy. In addition, yttrium additions were shown to stabilise the -phase of the alloy that decreases the hardness of the alloy. Keywords: microalloying, rare-earth elements, titanium alloy, welding U~inek mikrolegiranja zlitine titana Ti-4.8Al-1.2Mo-2.6V-0.6Cr-0.25Fe z itrijem in cirkonijem na sestavo faze, strukturne in mehanske lastnosti je bil raziskovan s svetlobno elektronsko mikroskopijo, z rentgensko difrakcijsko analizo in s testiranjem trdote mikroindentacije. Uporabili smo diferencialno termi~no analizo za pridobitev modela za varilni termi~ni cikel. Ugotovljeno je bilo, da legiranje z itrijem vodi k razpr{enosti Y2O3 oksidnih delcev, kar se ka`e pri rafiniranju mikrostrukture zlitine. Poleg tega je bilo dokazano, da so dodatki itrija stabilizirali -fazo v zlitini, ki zmanj{uje trdoto zlitine. Klju~ne besede: mikrolegiranje, elementi redke zemlje, zlitina titana, varjenje 1 INTRODUCTION Titanium alloys, due to their high specific strength and good corrosion resistance, are widely used in the aerospace and marine industries as structural materials and it is desirable to further enhance their properties. The latter is possible by means of complex multicomponent alloying.1,2 In addition to the specific strength, good weldability is usually necessary to maximise the material utilisation. Since -titanium and + alloys are considered to have a better weldability compared to metastable -tita- nium alloys,3 a number of +-alloys were developed in Russia to meet these requirements. One of them is Ti-4.8Al-1.2Mo-2.6V-0.6Cr–0.25Fe, which is a typical martensite-type two-phase alloy.4 Despite good weldability, the welding usually negatively affects the mechanical properties of titanium alloys.5 This is partly due to the oxygen absorbed from the shield atmosphere followed by its diffusion to the weld-fusion zone (FZ), partly due to a high heat input resulting in significant grain coarsening. One way to overcome such a problem is to refine the prior grain size by the grain-boundary pinning effect of second-phase particles and promote heterogeneous nucleation in the FZ.6 It was shown that microalloying with transition elements, especially rare-earth metals, is a promising way to refine the structure of titanium alloys.7 One of the most frequently employed alloying elements is yttrium.8–11 Although it was already shown what microalloying of the Ti-4.8Al-1.2Mo-2.6V-0.6Cr–0.25Fe alloy with 0.06 % mass fraction of yttrium decreases the oxygen concentration in - and -solid solutions due to the formation of Y2O3 particles,4 detailed studies of a weld structure were not conducted. This is why our study aimed at investigating the influence of microalloying titanium alloys with zirconium and yttrium on the microstructure formation and mechanical properties of the weld joints of this alloy. 2 EXPERIMENTAL PART Two-millimeter-thick hot-rolled sheets of an (+) martensitic titanium alloy alloyed with yttrium (alloy 1) and zirconium (alloy 2) were used in this study. The chemical compositions of the alloys are shown in Table 1. Prior to their use, the sheets were subjected to vacuum annealing at 750 °C for 1 h followed by arc welding using non-consumable tungsten electrodes. Materiali in tehnologije / Materials and technology 51 (2017) 5, 855–859 855 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 669.794:669.296:67.017:66.046.51 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)855(2017) A. G. ILLARIONOV et al.: EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE ... 855–859 EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE AND PROPERTIES OF WELD JOINTS OF A TWO-PHASE TITANIUM ALLOY U^INEK MIKROLEGIRANJA ITRIJA IN CIRKONIJA NA STRUKTURO IN LASTNOSTI NA SPOJE ZAVROV DVOFAZNE ZLITINE TITANA Anatoly Illarionov, Artemy Popov, Svetlana Illarionova, Dmitry Gadeev Ural Federal University, 620034, 28 Mira str., Yekaterinburg, Russia d.v.gadeev@urfu.ru Prejem rokopisa – received: 2016-11-10; sprejem za objavo – accepted for publication: 2017-04-19 doi:10.17222/mit.2016.317 The effect of microalloying of the Ti-4.8Al-1.2Mo-2.6V-0.6Cr–0.25Fe titanium alloy with yttrium and zirconium on the phase composition, structure and mechanical properties of welded sheets was studied using light and transmission electron microscopy, X-ray diffraction analysis and microindentation-hardness testing. A differential thermal analysis was employed to model the welding thermal cycle. It was found that alloying with yttrium led to the precipitation of Y2O3 oxide particles resulting in refining the microstructure of the alloy. In addition, yttrium additions were shown to stabilise the -phase of the alloy that decreases the hardness of the alloy. Keywords: microalloying, rare-earth elements, titanium alloy, welding U~inek mikrolegiranja zlitine titana Ti-4.8Al-1.2Mo-2.6V-0.6Cr-0.25Fe z itrijem in cirkonijem na sestavo faze, strukturne in mehanske lastnosti je bil raziskovan s svetlobno elektronsko mikroskopijo, z rentgensko difrakcijsko analizo in s testiranjem trdote mikroindentacije. Uporabili smo diferencialno termi~no analizo za pridobitev modela za varilni termi~ni cikel. Ugotovljeno je bilo, da legiranje z itrijem vodi k razpr{enosti Y2O3 oksidnih delcev, kar se ka`e pri rafiniranju mikrostrukture zlitine. Poleg tega je bilo dokazano, da so dodatki itrija stabilizirali -fazo v zlitini, ki zmanj{uje trdoto zlitine. Klju~ne besede: mikrolegiranje, elementi redke zemlje, zlitina titana, varjenje 1 INTRODUCTION Titanium alloys, due to their high specific strength and good corrosion resistance, are widely used in the aerospace and marine industries as structural materials and it is desirable to further enhance their properties. The latter is possible by means of complex multicomponent alloying.1,2 In addition to the specific strength, good weldability is usually necessary to maximise the material utilisation. Since -titanium and + alloys are considered to have a better weldability compared to metastable -tita- nium alloys,3 a number of +-alloys were developed in Russia to meet these requirements. One of them is Ti-4.8Al-1.2Mo-2.6V-0.6Cr–0.25Fe, which is a typical martensite-type two-phase alloy.4 Despite good weldability, the welding usually negatively affects the mechanical properties of titanium alloys.5 This is partly due to the oxygen absorbed from the shield atmosphere followed by its diffusion to the weld-fusion zone (FZ), partly due to a high heat input resulting in significant grain coarsening. One way to overcome such a problem is to refine the prior grain size by the grain-boundary pinning effect of second-phase particles and promote heterogeneous nucleation in the FZ.6 It was shown that microalloying with transition elements, especially rare-earth metals, is a promising way to refine the structure of titanium alloys.7 One of the most frequently employed alloying elements is yttrium.8–11 Although it was already shown what microalloying of the Ti-4.8Al-1.2Mo-2.6V-0.6Cr–0.25Fe alloy with 0.06 % mass fraction of yttrium decreases the oxygen concentration in - and -solid solutions due to the formation of Y2O3 particles,4 detailed studies of a weld structure were not conducted. This is why our study aimed at investigating the influence of microalloying titanium alloys with zirconium and yttrium on the microstructure formation and mechanical properties of the weld joints of this alloy. 2 EXPERIMENTAL PART Two-millimeter-thick hot-rolled sheets of an (+) martensitic titanium alloy alloyed with yttrium (alloy 1) and zirconium (alloy 2) were used in this study. The chemical compositions of the alloys are shown in Table 1. Prior to their use, the sheets were subjected to vacuum annealing at 750 °C for 1 h followed by arc welding using non-consumable tungsten electrodes. Materiali in tehnologije / Materials and technology 51 (2017) 5, 855–859 855 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 669.794:669.296:67.017:66.046.51 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)855(2017) Table 1: Chemical compositions of the test materials, in mass fractions (w/%) Materials Alloying elements, in mass fractions (w/%) Ti Al Mo V Cr Fe Y Zr Alloy 1 Bal 4.80 1.20 2.60 0.60 0.60 0.06 - Alloy 2 - 0.07 The microstructures of the alloys were analyzed with light optical (LOM) and transmission electron (TEM) microscopy on Olympus GX51 and JEM 200C micro- scopes, respectively. A qualitative phase analysis was carried out using X-ray diffraction (XRD) with Cu-K radiation on a DRON-3M powder diffractometer. A differential thermal analysis (DTA) was performed on a Du Pont appliance in DTA-1600 crucibles. The average grain size and grain-size distribution were determined with an Epiquant optical microscope. Vickers microhardness was measured at a 1 N load in accordance with the formula below (ISO 6507) and multiplied with the standard gravity (g = 9.81) to get the values in MPa in Equation (1): HV = 0.1891· (F/d2) (1) with F being the applied load (Newton) and d being the average length of the diagonal of the residual indent (millimeters). Metallographic samples were prepared according to the standard procedures. The steps consisted of grinding with 120–2400 grit SiC paper, polishing with 0.3 μm colloidal alumina, followed by the final polishing with a 0.05 μm colloidal silica suspension. Kroll’ s reagent with a composition of 100 mL water + 2 mL HF + 5 mL HNO3 was used to etch the specimens.12 Thin foils for the TEM investigation were prepared with electrolytic polishing in a methanol-containing electrolyte (300 mL methanol, 175 mL butanol, 30 mL perchloric acid (70–72 %)) at -30 °C.12 3 RESULTS AND DISCUSSION With respect to the differences in the microstructure, weld joints generally consist of three distinct zones, i.e., the base metal, the fusion zone and the heat-affected zone. In the fusion zone (FZ), materials melt during the welding and crystallize during the subsequent cooling. The heat-affected zone (HAZ) corresponds to the narrow transition region between the base metal and the fusion zone where the material is heated up to sub-transus or even superheated above the critical temperature. The latter might obviously cause significant microstructure changes in comparison with the base metal. On the present samples, the zone width was measured and it was around 12 mm and 4.5 mm for the fusion and heat-affected zones, respectively (Figure 1). The base-metal microstructure of both alloys is represented by primary -phase precipitates of a mixed, globular and lamellar, morphology in a -phase matrix (Figure 2a). The yttrium-containing alloy 1 is dis- tinguished by a dispersive oxide precipitation of Y2O3 particles with the average diameter of up to 120 nm (Figure 2b). These particles were found to be almost spherical and to precipitate mainly near the grain boundaries. The TEM micrograph (Figure 2b) shows a specific 'comet-tail-like' contrast around these particles, which is attributed to the elastic-stress field caused by the difference between the thermal-expansion coeffi- cients of the oxides and the matrix. The average -phase grain size decreased with the increasing distance from the FZ (Figure 1) due to the higher temperatures achieved in the HAZ close to the A. G. ILLARIONOV et al.: EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE ... 856 Materiali in tehnologije / Materials and technology 51 (2017) 5, 855–859 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Microstructures of base metals: a) alloy 2 (LOM); b) alloy 1 (TEM) Figure 1: Structures of welded joints: a) alloy 1 with Y; b) alloy 2 with Zr FZ. The coarsest grains of alloy 2 reached about 400 μm in diameter, whereas in alloy 1, they did not exceed 330 μm, i.e., the yttrium-containing alloy 1 was characterized by a 1.2 times finer grain size. A similar relationship was previously observed in the solution-treated and water- quenched alloys of the same composition4 and it was attributed to the presence of the yttrium-oxide particles in the Y-containing alloy that inhibited the grain-boun- dary movement. The X-ray diffraction analysis showed that the vo- lume fractions of the - and -phases varied significantly in different parts of the HAZ (Figure 3). The intensity of -phase peaks in the XRD patterns decreased with the increasing distance from the FZ, clearly indicating that the volume fraction of the -phase was significantly higher in the near-FZ areas. At the same time, the a lattice parameter changed in the opposite direction. It can be seen that the (200)-peak intensity ratio for alloy 1 was considerably higher than that for alloy 2. This confirms the assumption that yttrium binds oxygen, thus increasing the stability of the supercooled -phase upon cooling.4,8–9 Furthermore, such a behaviour resulted in a higher -phase volume fraction, decreasing the hard- ness of the alloy. This was confirmed with the micro- hardness-indentation tests. Alloy 1 had a lower micro- hardness (2800±50 MPa) than alloy 2 (3300±50 MPa). It was also found that the -phase morphology along the HAZ changed from a complex one (Figure 2) to a lamellar one consisting of thin platelets (Figure 4). The analysis of the microstructures of the alloys revealed a noticeable feature of both alloys, i.e., a jagged shape of the grain boundaries. A similar phenomenon was seen previously, for instance, in Ni–Mn–In-based alloys after thermal cycling13 and near-alpha titanium alloys14 and it was attributed to the formation of marten- site crystals, which grow towards the grain boundaries curving them. Figure 4a shows that such a shape coin- cides closely with the lamellar -phase precipitations on both sides of the grain boundaries. Thin -lamellae were formed during the cooling from the single-phase -region. The stress field next to the growing -lamellae could interact with the grain boundaries curving them. The same process occurred in the neighbouring grains resulting in the jag formation. Since this process can occur in preferentially oriented grains, the jags were found only on the corresponding parts of the grains. On the whole, we suppose that the jagged shape of the grain boundaries is associated with the interaction of the boundaries with the products of the -transformation occurring upon the cooling. To summarize the above, it can be concluded that the welding thermal cycle results in significant changes of the grain structure, phase composition and morphology of the phases. Yttrium binds oxygen forming oxide particles that inhibit the grain growth and improve the -phase stability. In turn, microalloying the alloy with zirconium does not result in the formation of additional phases and leads to the solution hardening of the alloy. The average grain sizes for the weld joints of the samples of alloys 1 and 2 did not show any significant difference and was about 480–600 μm (Figure 1). The fine structure of the alloys mainly consists of the -phase, with dislocations perpendicular to the lamellae (Figure 5b). The -lamellae in the FZs of the alloys tend to form small colonies that may cross differently oriented lamellae (Figure 5c). Interlamellar / spacings in A. G. ILLARIONOV et al.: EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 855–859 857 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: -phase lattice parameter (a) and (200) XRD-peak inten- sity ratio (IHAZ /IFZ) as a function of distance from the fusion zone of a welded joint Figure 4: Microstructures of heat-affected zones of weld joints: a) alloy 1, b) alloy 2 Table 1: Chemical compositions of the test materials, in mass fractions (w/%) Materials Alloying elements, in mass fractions (w/%) Ti Al Mo V Cr Fe Y Zr Alloy 1 Bal 4.80 1.20 2.60 0.60 0.60 0.06 - Alloy 2 - 0.07 The microstructures of the alloys were analyzed with light optical (LOM) and transmission electron (TEM) microscopy on Olympus GX51 and JEM 200C micro- scopes, respectively. A qualitative phase analysis was carried out using X-ray diffraction (XRD) with Cu-K radiation on a DRON-3M powder diffractometer. A differential thermal analysis (DTA) was performed on a Du Pont appliance in DTA-1600 crucibles. The average grain size and grain-size distribution were determined with an Epiquant optical microscope. Vickers microhardness was measured at a 1 N load in accordance with the formula below (ISO 6507) and multiplied with the standard gravity (g = 9.81) to get the values in MPa in Equation (1): HV = 0.1891· (F/d2) (1) with F being the applied load (Newton) and d being the average length of the diagonal of the residual indent (millimeters). Metallographic samples were prepared according to the standard procedures. The steps consisted of grinding with 120–2400 grit SiC paper, polishing with 0.3 μm colloidal alumina, followed by the final polishing with a 0.05 μm colloidal silica suspension. Kroll’ s reagent with a composition of 100 mL water + 2 mL HF + 5 mL HNO3 was used to etch the specimens.12 Thin foils for the TEM investigation were prepared with electrolytic polishing in a methanol-containing electrolyte (300 mL methanol, 175 mL butanol, 30 mL perchloric acid (70–72 %)) at -30 °C.12 3 RESULTS AND DISCUSSION With respect to the differences in the microstructure, weld joints generally consist of three distinct zones, i.e., the base metal, the fusion zone and the heat-affected zone. In the fusion zone (FZ), materials melt during the welding and crystallize during the subsequent cooling. The heat-affected zone (HAZ) corresponds to the narrow transition region between the base metal and the fusion zone where the material is heated up to sub-transus or even superheated above the critical temperature. The latter might obviously cause significant microstructure changes in comparison with the base metal. On the present samples, the zone width was measured and it was around 12 mm and 4.5 mm for the fusion and heat-affected zones, respectively (Figure 1). The base-metal microstructure of both alloys is represented by primary -phase precipitates of a mixed, globular and lamellar, morphology in a -phase matrix (Figure 2a). The yttrium-containing alloy 1 is dis- tinguished by a dispersive oxide precipitation of Y2O3 particles with the average diameter of up to 120 nm (Figure 2b). These particles were found to be almost spherical and to precipitate mainly near the grain boundaries. The TEM micrograph (Figure 2b) shows a specific 'comet-tail-like' contrast around these particles, which is attributed to the elastic-stress field caused by the difference between the thermal-expansion coeffi- cients of the oxides and the matrix. The average -phase grain size decreased with the increasing distance from the FZ (Figure 1) due to the higher temperatures achieved in the HAZ close to the A. G. ILLARIONOV et al.: EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE ... 856 Materiali in tehnologije / Materials and technology 51 (2017) 5, 855–859 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Microstructures of base metals: a) alloy 2 (LOM); b) alloy 1 (TEM) Figure 1: Structures of welded joints: a) alloy 1 with Y; b) alloy 2 with Zr FZ. The coarsest grains of alloy 2 reached about 400 μm in diameter, whereas in alloy 1, they did not exceed 330 μm, i.e., the yttrium-containing alloy 1 was characterized by a 1.2 times finer grain size. A similar relationship was previously observed in the solution-treated and water- quenched alloys of the same composition4 and it was attributed to the presence of the yttrium-oxide particles in the Y-containing alloy that inhibited the grain-boun- dary movement. The X-ray diffraction analysis showed that the vo- lume fractions of the - and -phases varied significantly in different parts of the HAZ (Figure 3). The intensity of -phase peaks in the XRD patterns decreased with the increasing distance from the FZ, clearly indicating that the volume fraction of the -phase was significantly higher in the near-FZ areas. At the same time, the a lattice parameter changed in the opposite direction. It can be seen that the (200)-peak intensity ratio for alloy 1 was considerably higher than that for alloy 2. This confirms the assumption that yttrium binds oxygen, thus increasing the stability of the supercooled -phase upon cooling.4,8–9 Furthermore, such a behaviour resulted in a higher -phase volume fraction, decreasing the hard- ness of the alloy. This was confirmed with the micro- hardness-indentation tests. Alloy 1 had a lower micro- hardness (2800±50 MPa) than alloy 2 (3300±50 MPa). It was also found that the -phase morphology along the HAZ changed from a complex one (Figure 2) to a lamellar one consisting of thin platelets (Figure 4). The analysis of the microstructures of the alloys revealed a noticeable feature of both alloys, i.e., a jagged shape of the grain boundaries. A similar phenomenon was seen previously, for instance, in Ni–Mn–In-based alloys after thermal cycling13 and near-alpha titanium alloys14 and it was attributed to the formation of marten- site crystals, which grow towards the grain boundaries curving them. Figure 4a shows that such a shape coin- cides closely with the lamellar -phase precipitations on both sides of the grain boundaries. Thin -lamellae were formed during the cooling from the single-phase -region. The stress field next to the growing -lamellae could interact with the grain boundaries curving them. The same process occurred in the neighbouring grains resulting in the jag formation. Since this process can occur in preferentially oriented grains, the jags were found only on the corresponding parts of the grains. On the whole, we suppose that the jagged shape of the grain boundaries is associated with the interaction of the boundaries with the products of the -transformation occurring upon the cooling. To summarize the above, it can be concluded that the welding thermal cycle results in significant changes of the grain structure, phase composition and morphology of the phases. Yttrium binds oxygen forming oxide particles that inhibit the grain growth and improve the -phase stability. In turn, microalloying the alloy with zirconium does not result in the formation of additional phases and leads to the solution hardening of the alloy. The average grain sizes for the weld joints of the samples of alloys 1 and 2 did not show any significant difference and was about 480–600 μm (Figure 1). The fine structure of the alloys mainly consists of the -phase, with dislocations perpendicular to the lamellae (Figure 5b). The -lamellae in the FZs of the alloys tend to form small colonies that may cross differently oriented lamellae (Figure 5c). Interlamellar / spacings in A. G. ILLARIONOV et al.: EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 855–859 857 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: -phase lattice parameter (a) and (200) XRD-peak inten- sity ratio (IHAZ /IFZ) as a function of distance from the fusion zone of a welded joint Figure 4: Microstructures of heat-affected zones of weld joints: a) alloy 1, b) alloy 2 alloy 2 have a complex structure and high dislocation density (Figure 5d). -phase interlayers are relatively wide (up to 200 nm) and non-uniform throughout the bulk of the alloy. Secondary -phase precipitations were observed in the interlamellar spacings, both along and across the  interlayers as well as at the /-interphase boundaries (Figure 5d). Generally, the coarse -platelets with a small dislo- cation density prevail in both yttrium- and zirconium- containing alloys (Figures 5e, 5f). However, alloy 1 is distinguished by a more conventional structure of inter- lamellar spacing. Although the thin lenticular platelets of the secondary -phase were still present, no highly dispersive precipitates were found on the interphase boundaries and within the -layers. It should be noted that no Y2O3 particles were detected in the FZ. This might be connected with the intense overheating of the alloy caused by welding that resulted in a dissociation of the oxide. A. G. ILLARIONOV et al.: EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE ... 858 Materiali in tehnologije / Materials and technology 51 (2017) 5, 855–859 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: Thin structures of the weld joints of alloys 1 and 2: a), c), d) alloy 1; b), e), f) alloy 2 Figure 6: DTA curves for continuous cooling (20 °C/min cooling rate) from 1200 °C: a) alloy 1, b) alloy 2 In order to model the welding thermal cycle, a num- ber of DTA experiments involving continuous heating up to 1200 °C and the subsequent cooling to 20 °C/min were carried out. Continuous-cooling DTA curves allowed us to determine the temperature ranges of the -transformation (Figure 6). A comparative analysis of the curves showed that the -phase decomposition process in alloy 2 consisted of three distinctive stages, whereas only two exothermal effects were found for alloy 1. Stages I and II (at 920–750 and 700–550 °C, respectively) were attributed to the formation of coarse primary -colonies with a low defect density and secondary -platelets within the -matrix. Stage III (500–400 °C) was assigned to the precipitation of the highly dispersed -precipitates mentioned above (Figure 5 d). The formation of the -phase upon supercooling below the transus temperature of the titanium alloys is initiated by the crystallographic defects of the structure such as grain boundaries, dislocations, etc. Yttrium binds oxygen and thus refines the alloy. The products of the -phase decomposition were quite free of precipitates. On the other hand, the precipitation of the secondary -platelets was possible in both alloys studied during supercooling of up to about 600 °C. At lower tem- peratures, the -phase was stabilized in alloy 1, whereas its further decomposition occurred in alloy 2 resulting in the formation of highly dispersed -precipitates. We believe that all these differences in the behaviour of the alloys are associated with different -phase stability values, which are primarily defined by the concentration of impurities in the alloys. Specifically, the interstitial elements in alloy 1, mainly oxygen, are bound to yttrium in the form of oxides. Zirconium has a lower affinity for oxygen and thus oxygen atoms remain in the solid solution leading to a more complete -phase decom- position upon the cooling (Figure 2) resulting in a lower -phase lattice parameter. With an increase in the volume fraction of the body-centred cubic -phase, the strength of the titanium alloys usually decreases.12 This is why the microhardness of Y-containing alloy 1 was found to be lower (3300±100 MPa for alloy 2) and 3000±100 MPa for alloy 1). 4 CONCLUSIONS The influence of the microalloying of titanium alloy Ti–4.8À1–1.2Ìî–2.6V–0.6Ñr–0.25Fe with yttrium and zirconium on the structure and microhardness was considered. It was found that the alloying with yttrium has the most noticeable effect on the structure of a weld joint and the phase transformation of the alloy. A yttrium addition refines the alloy by means of binding with oxygen atoms and forming Y2O3 particles. These particles pin the grain boundaries in both the heat-affected and fusion zones, inhibiting the grain growth during a welding thermal cycle. In addition, due to a lower concentration of oxygen in the -phase, yttrium decreases the -transus temperature of the alloy leading to the stabilization of the supercooled -phase. Due to a higher volume fraction of the "soft" -phase in the yttrium-containing alloy, its microhardness was shown to be lower than that of the alloy with zirconium. Acknowledgment We hereby acknowledge the support of the Ministry of Science and Education of the Russian Federation, in accordance with the decree of the Government of 9 April 2010, No. 218, project No. 03.G25.31.0234. 5 REFERENCES 1 A I. Khorev, Development of structural titanium alloys for components and sections in aerospace technology, 10 (2010), 13–23, doi:10.1080/09507116.2010.486188 2 A. I. Khorev, Complex alloying and heat treatment of titanium alloys, 6 (2008), 364–368, doi:10.1080/09507110802288312 3 M. J. Donachie, Titanium, A Technical Guide, 2nd ed., ASM Interna- tional, 2000 4 A. G. Illarionov, A. A. Popov, S. M. Illarionova, Effect of Micro- alloying, with Rem Inclusively, on the Structure, Phase Composition and Properties of ( + )-Titanium Alloy, Metal Science and Heat Treatment, 57 (2016), 719–725, doi:10.1007/s11041-016-9948-0 5 E. W. Collings, The physical metallurgy of titanium alloys, American Society for Metals, 1984, 261 6 Rios PR. Overview No. 62: A theory for grain boundary pinning by particles, Acta Metall., 35 (1987) 12, 2805–2814 7 A. I. Khorev, Theory and practices of microalloying of near- and + titanium alloys with REE, zirconium, hafnium and rhenium, Techologii machinostroeniya, 1 (2015), 5–10 8 H. Wu, Y. Han, X. Chen, Effects of Yttrium on Mechanical Pro- perties and Microstructures of Ti-Si Eutectic Alloy, Chinese Journal of Aeronautics, 18 (2005), 171–174, doi:10.1016/S1000-9361(11) 60324-5 9 R. P. Kolli, A. A. Herzing, S. Ankem, Characterization of yttrium- rich precipitates in a titanium alloy weld, Materials Characterization, 122 (2016), 30–35, doi:10.1016/j.matchar.2016.10.014 10 B. Poorganji, A. Kazahari, T. Narushima, C. Ouchi, T. Furuhara, Effect of yttrium addition on grain growth of ,  and + titanium alloys, Journal of Physics: Conference Series, 240 (2010), 12170, doi:10.1088/1742-6596/240/1/012170 11 W. F. Cui, C. M. Liu, L. Zhou, G. Z. Luo, Characteristics of microstructures and second-phase particles in Y-bearing Ti-1100 alloy, Materials Science and Engineering: A, 323 (2002), 192–197, doi:10.1016/S0921-5093(01)01362-4 12 G. Lütjering, J. C. Williams, Titanium, 2nd ed., Springer, Berlin, 2007, 442 13 Y. V. Kaletina, E. D. Efimova, E. G. Gerasimov, A. Y. Kaletin, Effect of thermal cycling on structure and properties of Ni–Mn–In-based alloys, Technical Physics, 12 (2016), 1894–1897 14 H. M. Flower, P. R. Swann, D. R. F. West, The effect of Si, Zr, Al and Mo on the structure and strength of Ti martensite, Journal of Materials Science, 8 (1972), 929–938 A. G. ILLARIONOV et al.: EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 855–859 859 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS alloy 2 have a complex structure and high dislocation density (Figure 5d). -phase interlayers are relatively wide (up to 200 nm) and non-uniform throughout the bulk of the alloy. Secondary -phase precipitations were observed in the interlamellar spacings, both along and across the  interlayers as well as at the /-interphase boundaries (Figure 5d). Generally, the coarse -platelets with a small dislo- cation density prevail in both yttrium- and zirconium- containing alloys (Figures 5e, 5f). However, alloy 1 is distinguished by a more conventional structure of inter- lamellar spacing. Although the thin lenticular platelets of the secondary -phase were still present, no highly dispersive precipitates were found on the interphase boundaries and within the -layers. It should be noted that no Y2O3 particles were detected in the FZ. This might be connected with the intense overheating of the alloy caused by welding that resulted in a dissociation of the oxide. A. G. ILLARIONOV et al.: EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE ... 858 Materiali in tehnologije / Materials and technology 51 (2017) 5, 855–859 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: Thin structures of the weld joints of alloys 1 and 2: a), c), d) alloy 1; b), e), f) alloy 2 Figure 6: DTA curves for continuous cooling (20 °C/min cooling rate) from 1200 °C: a) alloy 1, b) alloy 2 In order to model the welding thermal cycle, a num- ber of DTA experiments involving continuous heating up to 1200 °C and the subsequent cooling to 20 °C/min were carried out. Continuous-cooling DTA curves allowed us to determine the temperature ranges of the -transformation (Figure 6). A comparative analysis of the curves showed that the -phase decomposition process in alloy 2 consisted of three distinctive stages, whereas only two exothermal effects were found for alloy 1. Stages I and II (at 920–750 and 700–550 °C, respectively) were attributed to the formation of coarse primary -colonies with a low defect density and secondary -platelets within the -matrix. Stage III (500–400 °C) was assigned to the precipitation of the highly dispersed -precipitates mentioned above (Figure 5 d). The formation of the -phase upon supercooling below the transus temperature of the titanium alloys is initiated by the crystallographic defects of the structure such as grain boundaries, dislocations, etc. Yttrium binds oxygen and thus refines the alloy. The products of the -phase decomposition were quite free of precipitates. On the other hand, the precipitation of the secondary -platelets was possible in both alloys studied during supercooling of up to about 600 °C. At lower tem- peratures, the -phase was stabilized in alloy 1, whereas its further decomposition occurred in alloy 2 resulting in the formation of highly dispersed -precipitates. We believe that all these differences in the behaviour of the alloys are associated with different -phase stability values, which are primarily defined by the concentration of impurities in the alloys. Specifically, the interstitial elements in alloy 1, mainly oxygen, are bound to yttrium in the form of oxides. Zirconium has a lower affinity for oxygen and thus oxygen atoms remain in the solid solution leading to a more complete -phase decom- position upon the cooling (Figure 2) resulting in a lower -phase lattice parameter. With an increase in the volume fraction of the body-centred cubic -phase, the strength of the titanium alloys usually decreases.12 This is why the microhardness of Y-containing alloy 1 was found to be lower (3300±100 MPa for alloy 2) and 3000±100 MPa for alloy 1). 4 CONCLUSIONS The influence of the microalloying of titanium alloy Ti–4.8À1–1.2Ìî–2.6V–0.6Ñr–0.25Fe with yttrium and zirconium on the structure and microhardness was considered. It was found that the alloying with yttrium has the most noticeable effect on the structure of a weld joint and the phase transformation of the alloy. A yttrium addition refines the alloy by means of binding with oxygen atoms and forming Y2O3 particles. These particles pin the grain boundaries in both the heat-affected and fusion zones, inhibiting the grain growth during a welding thermal cycle. In addition, due to a lower concentration of oxygen in the -phase, yttrium decreases the -transus temperature of the alloy leading to the stabilization of the supercooled -phase. Due to a higher volume fraction of the "soft" -phase in the yttrium-containing alloy, its microhardness was shown to be lower than that of the alloy with zirconium. Acknowledgment We hereby acknowledge the support of the Ministry of Science and Education of the Russian Federation, in accordance with the decree of the Government of 9 April 2010, No. 218, project No. 03.G25.31.0234. 5 REFERENCES 1 A I. Khorev, Development of structural titanium alloys for components and sections in aerospace technology, 10 (2010), 13–23, doi:10.1080/09507116.2010.486188 2 A. I. Khorev, Complex alloying and heat treatment of titanium alloys, 6 (2008), 364–368, doi:10.1080/09507110802288312 3 M. J. Donachie, Titanium, A Technical Guide, 2nd ed., ASM Interna- tional, 2000 4 A. G. Illarionov, A. A. Popov, S. M. Illarionova, Effect of Micro- alloying, with Rem Inclusively, on the Structure, Phase Composition and Properties of ( + )-Titanium Alloy, Metal Science and Heat Treatment, 57 (2016), 719–725, doi:10.1007/s11041-016-9948-0 5 E. W. Collings, The physical metallurgy of titanium alloys, American Society for Metals, 1984, 261 6 Rios PR. Overview No. 62: A theory for grain boundary pinning by particles, Acta Metall., 35 (1987) 12, 2805–2814 7 A. I. Khorev, Theory and practices of microalloying of near- and + titanium alloys with REE, zirconium, hafnium and rhenium, Techologii machinostroeniya, 1 (2015), 5–10 8 H. Wu, Y. Han, X. Chen, Effects of Yttrium on Mechanical Pro- perties and Microstructures of Ti-Si Eutectic Alloy, Chinese Journal of Aeronautics, 18 (2005), 171–174, doi:10.1016/S1000-9361(11) 60324-5 9 R. P. Kolli, A. A. Herzing, S. Ankem, Characterization of yttrium- rich precipitates in a titanium alloy weld, Materials Characterization, 122 (2016), 30–35, doi:10.1016/j.matchar.2016.10.014 10 B. Poorganji, A. Kazahari, T. Narushima, C. Ouchi, T. Furuhara, Effect of yttrium addition on grain growth of ,  and + titanium alloys, Journal of Physics: Conference Series, 240 (2010), 12170, doi:10.1088/1742-6596/240/1/012170 11 W. F. Cui, C. M. Liu, L. Zhou, G. Z. Luo, Characteristics of microstructures and second-phase particles in Y-bearing Ti-1100 alloy, Materials Science and Engineering: A, 323 (2002), 192–197, doi:10.1016/S0921-5093(01)01362-4 12 G. Lütjering, J. C. Williams, Titanium, 2nd ed., Springer, Berlin, 2007, 442 13 Y. V. Kaletina, E. D. Efimova, E. G. Gerasimov, A. Y. Kaletin, Effect of thermal cycling on structure and properties of Ni–Mn–In-based alloys, Technical Physics, 12 (2016), 1894–1897 14 H. M. Flower, P. R. Swann, D. R. F. West, The effect of Si, Zr, Al and Mo on the structure and strength of Ti martensite, Journal of Materials Science, 8 (1972), 929–938 A. G. ILLARIONOV et al.: EFFECT OF YTTRIUM AND ZIRCONIUM MICROALLOYING ON THE STRUCTURE ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 855–859 859 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... 861–869 MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu FRICTION-STIR SPOT WELDS RAZVOJ MIKROSTRUKTURE IN STATISTI^NA ANALIZA VRTILNO-TORNIH TO^KASTIH ZVAROV Al/Cu Mukuna Patrick Mubiayi1, Esther Titilayo Akinlabi1, Mamookho Elizabeth Makhatha2 1University of Johannesburg, Department of Mechanical Engineering Science, Auckland Park Kingsway Campus, 2006 Johannesburg, South Africa 2University of Johannesburg, Department of Metallurgy, School of Mining, Metallurgy and Chemical Engineering, Doornfontein Campus, 2028 Doornfontein, South Africa patrickmubiayi@gmail.com Prejem rokopisa – received: 2016-11-18; sprejem za objavo – accepted for publication: 2017-03-16 doi:10.17222/mit.2016.320 In this paper, friction-stir spot welding (FSSW) is performed on 3 mm thick AA1060 and C11000 using different process parameters and tool geometries. The microstructure- and the microhardness-profile analyses were conducted and the probability distribution function (PDF) of the obtained microhardness values was determined. Optical images showed a good material mixing in most of the spot welds produced, whereas the energy-dispersive-spectroscopy (EDS) analysis showed the presence of intermetallic compounds. Microhardness results revealed that process parameters and tool geometries have significant effects on the distribution of microhardness values in different locations of the produced spot welds. Furthermore, goodness-of-fit values showed that most of the R2 values ranged between 0.8842 and 0.9999, which indicated how well the model fits with the experimental data. On the other hand, the residuals comprised positive and negative runs which also indicated the existence of a certain correlation with the experimentation. Keywords: aluminium, copper, friction-stir spot welding, statistical analysis V prispevku je predstavljeno vrtilno-torno to~kovno varjenje (angl. FSSW) na 3 mm debeli plo~evini AA1060 in C11000 z uporabo razli~nih procesnih parametrov in razli~no geometrijo orodij. Izvedene so bile analize profila mikrostrukture in mikrotrdote in dolo~ena je bila porazdelitvena funkcija verjetnosti za dobljene vrednosti mikrotrdote. Posnetki so pokazali dobro me{anje materiala na ve~ini izdelanih to~kovnih zvarov, medtem ko je EDS-analiza pokazala prisotnost intermetalnih spojin. Meritve mikrotrdote so pokazale, da imajo procesni parametri in geometrija orodij pomemben vpliv na porazdelitev mikrotrdote glede na razli~ne lokacije izdelanih to~kovnih zvarov. Nadalje je serija opazovanj oz. ocena pokazala, da je ve~ina R2 vrednosti rangiranih med 0,8842 in 0,9999. To potrjuje, da se model dobro ujema z eksperimentalnimi podatki. Po drugi strani pa razlika med pozitivnimi in negativnimi cikli ka`e na obstoj dolo~ene korelacije s preizku{anjem. Klju~ne besede: aluminij, baker, vrtilno-torno to~kovno varjenje, statisti~na analiza 1 INTRODUCTION Friction-stir welding (FSW) is a fairly new solid-state joining technique created and patented by The Welding Institute (TWI) in 1991 for butt and lap welding of ferrous and non-ferrous metals and plastics.1 Friction-stir spot welding (FSSW) is a novel variant of linear friction-stir welding (FSW) used for spot-welding applications.2 The FSSW process uses a non-consumable rotating tool which is plunged into the workpieces to be welded. Before attaining the selected plunge depth, the rotating tool is held at the same position for a fixed time, which is defined as the dwell time. Consequently, the rotating tool is withdrawn from the welded joint, leaving a solid-phase friction-stir spot weld behind. Throughout the FSSW process, the tool penetration depth and the Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 861 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 620.1:67.017:669.71:621.791 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)861(2017) Figure 1: Schematic diagram of the friction-stir spot-welding process3 M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... 861–869 MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu FRICTION-STIR SPOT WELDS RAZVOJ MIKROSTRUKTURE IN STATISTI^NA ANALIZA VRTILNO-TORNIH TO^KASTIH ZVAROV Al/Cu Mukuna Patrick Mubiayi1, Esther Titilayo Akinlabi1, Mamookho Elizabeth Makhatha2 1University of Johannesburg, Department of Mechanical Engineering Science, Auckland Park Kingsway Campus, 2006 Johannesburg, South Africa 2University of Johannesburg, Department of Metallurgy, School of Mining, Metallurgy and Chemical Engineering, Doornfontein Campus, 2028 Doornfontein, South Africa patrickmubiayi@gmail.com Prejem rokopisa – received: 2016-11-18; sprejem za objavo – accepted for publication: 2017-03-16 doi:10.17222/mit.2016.320 In this paper, friction-stir spot welding (FSSW) is performed on 3 mm thick AA1060 and C11000 using different process parameters and tool geometries. The microstructure- and the microhardness-profile analyses were conducted and the probability distribution function (PDF) of the obtained microhardness values was determined. Optical images showed a good material mixing in most of the spot welds produced, whereas the energy-dispersive-spectroscopy (EDS) analysis showed the presence of intermetallic compounds. Microhardness results revealed that process parameters and tool geometries have significant effects on the distribution of microhardness values in different locations of the produced spot welds. Furthermore, goodness-of-fit values showed that most of the R2 values ranged between 0.8842 and 0.9999, which indicated how well the model fits with the experimental data. On the other hand, the residuals comprised positive and negative runs which also indicated the existence of a certain correlation with the experimentation. Keywords: aluminium, copper, friction-stir spot welding, statistical analysis V prispevku je predstavljeno vrtilno-torno to~kovno varjenje (angl. FSSW) na 3 mm debeli plo~evini AA1060 in C11000 z uporabo razli~nih procesnih parametrov in razli~no geometrijo orodij. Izvedene so bile analize profila mikrostrukture in mikrotrdote in dolo~ena je bila porazdelitvena funkcija verjetnosti za dobljene vrednosti mikrotrdote. Posnetki so pokazali dobro me{anje materiala na ve~ini izdelanih to~kovnih zvarov, medtem ko je EDS-analiza pokazala prisotnost intermetalnih spojin. Meritve mikrotrdote so pokazale, da imajo procesni parametri in geometrija orodij pomemben vpliv na porazdelitev mikrotrdote glede na razli~ne lokacije izdelanih to~kovnih zvarov. Nadalje je serija opazovanj oz. ocena pokazala, da je ve~ina R2 vrednosti rangiranih med 0,8842 in 0,9999. To potrjuje, da se model dobro ujema z eksperimentalnimi podatki. Po drugi strani pa razlika med pozitivnimi in negativnimi cikli ka`e na obstoj dolo~ene korelacije s preizku{anjem. Klju~ne besede: aluminij, baker, vrtilno-torno to~kovno varjenje, statisti~na analiza 1 INTRODUCTION Friction-stir welding (FSW) is a fairly new solid-state joining technique created and patented by The Welding Institute (TWI) in 1991 for butt and lap welding of ferrous and non-ferrous metals and plastics.1 Friction-stir spot welding (FSSW) is a novel variant of linear friction-stir welding (FSW) used for spot-welding applications.2 The FSSW process uses a non-consumable rotating tool which is plunged into the workpieces to be welded. Before attaining the selected plunge depth, the rotating tool is held at the same position for a fixed time, which is defined as the dwell time. Consequently, the rotating tool is withdrawn from the welded joint, leaving a solid-phase friction-stir spot weld behind. Throughout the FSSW process, the tool penetration depth and the Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 861 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 620.1:67.017:669.71:621.791 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)861(2017) Figure 1: Schematic diagram of the friction-stir spot-welding process3 dwell time fundamentally determine the heat generation, material plasticization around the tool’s pin, weld geometry and, hence, the mechanical properties of the welded joint.2 Figure 1 depicts a schematic illustration of the FSSW technique. It should be noted that the FSSW process uses a non-consumable tool, which is similar to the FSW tool.4 The shoulder produces the frictional heat, whereas the pin creates the material flow between the work sheets.2,3 In addition to the tool, the other parameters involved in FSSW include the tool rotation speed, the tool plunge depth and the tool dwell time. These parameters have an effect on the strength, the surface texture of the welded joints2 and the existence of weld defects.3 A typical cross-section of a friction-stir spot weld displays five different structures including the parent material (PM), the heat-affected zone (HAZ), the thermomechanically affected zone (TMAZ), the stir zone (SZ) and the hook.2 Friction-stir welding (FSW) and friction-stir spot welding (FSSW) of aluminium and copper have thus far not been fully investigated due to the huge difference between their melting temperatures and the high chemi- cal affinity of both materials which facilitate the forma- tion of brittle intermetallic Al/Cu phases.6–18 Vickers hardness measurement is a common technique used to characterise the hardness of materials and it was reported that the presence of intermetallics affects the hardness values of the produced FS welds and FSS welds.11,17 On the other hand, statistical analyses of Al/Cu friction-stir welds and spot welds have not been well researched. S. Akinlabi and E. T. Akinlabi19 conducted statistical ana- lyses on the data obtained from dissimilar friction-stir butt welds of aluminium (AA5754) and copper (C11000) to understand the link between the process parameters and the properties of the resulting welds. They concluded that the downward vertical force has a significant effect on the ultimate tensile strength (UTS) of the produced welds. A robust relationship between the electrical resis- tivity and the heat input into the welds was also ob- served.19 Statistical analyses evaluate the effects of process parameters on the properties of the produced spot welds and establish the relationships amongst the process parameters. Therefore, in the current study, an effort was made to further understand the relationships between the process parameters and the microhardness profiles of FSSW welds between copper and aluminium using the proba- bility distribution function (PDF). Furthermore, the microstructure and the chemical analyses of the pro- duced spot welds were also studied. 2 MATERIALS AND WELDING PARAMETERS In this study, AA1060 and C11000 base materials with dimensions of 3 mm thickness, 600 mm length and 120 mm width were friction-stir spot lap welded. The chemical compositions of the two parent materials were determined using a spectrometer. The chemical compo- sition of the aluminium sheet is as follows: 0.058 % mass fraction of Si, 0.481 % mass fraction of Fe, 0.011 % mass fraction of Ga, 0.05 % mass fraction of other elements and the rest is Al. The chemical com- position of the copper sheet is: 0.137 % mass fraction of Zn, <0.1 % mass fractions of Pb, 0.02 % mass fraction of Ni, 0.023 % mass fraction of Al, 0.012 % mass fraction of Co, 0.077 % mass fraction of B, 0.036 % mass fraction of Sb, 0.043 % mass fraction of Nb, <0.492 % mass fraction of other elements and the rest is Cu. The sheets were friction-stir spot welded in a 30 mm overlap configuration. The spot welds were produced at the eNtsa of Nelson Mandela Metropolitan University, Port Elizabeth, South Africa using an MTS PDS I-Stir. The tool material used was H13 tool steel hardened to 50–52 HRC with a 4 mm tool pin, 5 mm tool diameter and 15 mm tool shoulder. The friction-stir spot welds were produced at rotational speeds of 800 min–1 and 1200 min–1, the tool-shoulder-plunge depths employed were 0.5 mm and 1 mm at a constant dwell time of 10 s. The two different tool profiles used in the current study were flat pin/flat shoulder and conical pin/concave shoulder tool, designated as FPS and CCS, respectively. The produced welds were designated as XX_XX_XX with the first part describing the tool geometry, the second part indicating the rotational speed and the third part indicating the shoulder plunge depth. The weld samples were sectioned using wire elec- trical discharge machining (WEDM), grinded and polished, mounted and prepared, using the ASTM standard metallographic procedure and ASTM Standard E3-11.20 A solution of FeCl3 (10g) + HCl (6 mL) + ethanol (C2H5OH) (20 mL) + H2O (80 mL) was used to etch the copper side of the spot welds, while the aluminium side was etched with H2O (190 mL) + HNO3 (5 mL) + HCl (10 mL) + HF (2 mL). The microstructure of the spot welds was studied using an optical microscope (Olympus BX51M) equipped with the Stream software. Scanning electron microscopy combined with energy dispersive spectroscopy (SEM/EDS) was used to further examine the microstructure and the chemical analyses, respec- tively. A TESCAN equipped with Oxford Instruments X-Max was used for the SEM/EDS analyses. The Vickers microhardness was determined using a dia- M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... 862 Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Representation of a spot weld with dashed lines illustrating the location of the microhardness-profile measurements mond-pyramid-indenter EMCO Test DuraScan tester. Two different locations on the spot welds were used, viz., the top and the bottom, as illustrated in Figure 2. The microhardness measurements were carried out from the keyhole for all the different parameters and tool geometries, to find the probability density function (PDF) of each one. The Matlab 2014 software program was used to determine the PDF. 3 RESULTS AND DISCUSSION The microstructure of the friction-stir spot welds was studied using an optical microscope and the results are depicted in Figures 3 and 4. Figure 3 illustrates the microstructures of the spot welds produced at: a) 800 min–1 and b) 1200 min–1, a 1 mm shoulder plunge depth using a flat pin and a flat shoulder tool. On the other hand, Figure 4 shows the microstructure of the spot weld produced at: a) 800 min–1 and b) 1200 min–1, a 0.5 mm shoulder plunge depth using a conical pin and a concave shoulder tool. It can be seen in Figure 3a that there are a copper ring and a mixture of Al/Cu particles in the stir zone. There is no palpable welding defect in the weld and copper is disseminated in this zone with different shapes. In the upper part of the joint, a large bulk of copper with irregular shapes can be observed (Figure 3a). The tool pin was inserted in the aluminium plate and the copper ring extruded upward from the lower copper plate into the aluminium plate was observed. This was in agree- ment with the previous work.15 Additionally, the intermixing of copper and alumi- nium was not homogenous for different spot welds and different microstructures were formed in different regions of the welds. It was reported that the FSW of dissimilar materials is different from that of similar materials due to the formation of a complex, intercalated vortex-like and related flow pattern.21 In Figure 4b, a good interlaced structure can be seen. This is formed by aluminium and copper, thereby indicating that the two plates are bonded firmly in this region, which is composed of a lamellar structure of copper particles with a streamlined shape of aluminium strips. In this region, a few disseminated copper particles were also observed. The energy-dispersive-spectrometry analyses at the selected points in the stir zone were recorded using SEM micrographs (Table 1). Intermetallic compounds were found in most of the produced welds. Two intermetallic compounds, viz., Al2Cu and Al3Cu4 were found in the weld produced at 1200 min–1, 0.5 mm shoulder plunge M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 863 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 4: Optical microscope images showing the macrostructures of the joints at: a) 800 min–1 and b) 1200 min–1, 0.5 mm shoulder plunge depth using a conical pin and concave shoulder tool Figure 3: Optical microscope images showing the macrostructures of the joints at: a) 800 min–1 and b) 1200 min–1, 1 mm shoulder plunge depth using a flat pin and flat shoulder tool dwell time fundamentally determine the heat generation, material plasticization around the tool’s pin, weld geometry and, hence, the mechanical properties of the welded joint.2 Figure 1 depicts a schematic illustration of the FSSW technique. It should be noted that the FSSW process uses a non-consumable tool, which is similar to the FSW tool.4 The shoulder produces the frictional heat, whereas the pin creates the material flow between the work sheets.2,3 In addition to the tool, the other parameters involved in FSSW include the tool rotation speed, the tool plunge depth and the tool dwell time. These parameters have an effect on the strength, the surface texture of the welded joints2 and the existence of weld defects.3 A typical cross-section of a friction-stir spot weld displays five different structures including the parent material (PM), the heat-affected zone (HAZ), the thermomechanically affected zone (TMAZ), the stir zone (SZ) and the hook.2 Friction-stir welding (FSW) and friction-stir spot welding (FSSW) of aluminium and copper have thus far not been fully investigated due to the huge difference between their melting temperatures and the high chemi- cal affinity of both materials which facilitate the forma- tion of brittle intermetallic Al/Cu phases.6–18 Vickers hardness measurement is a common technique used to characterise the hardness of materials and it was reported that the presence of intermetallics affects the hardness values of the produced FS welds and FSS welds.11,17 On the other hand, statistical analyses of Al/Cu friction-stir welds and spot welds have not been well researched. S. Akinlabi and E. T. Akinlabi19 conducted statistical ana- lyses on the data obtained from dissimilar friction-stir butt welds of aluminium (AA5754) and copper (C11000) to understand the link between the process parameters and the properties of the resulting welds. They concluded that the downward vertical force has a significant effect on the ultimate tensile strength (UTS) of the produced welds. A robust relationship between the electrical resis- tivity and the heat input into the welds was also ob- served.19 Statistical analyses evaluate the effects of process parameters on the properties of the produced spot welds and establish the relationships amongst the process parameters. Therefore, in the current study, an effort was made to further understand the relationships between the process parameters and the microhardness profiles of FSSW welds between copper and aluminium using the proba- bility distribution function (PDF). Furthermore, the microstructure and the chemical analyses of the pro- duced spot welds were also studied. 2 MATERIALS AND WELDING PARAMETERS In this study, AA1060 and C11000 base materials with dimensions of 3 mm thickness, 600 mm length and 120 mm width were friction-stir spot lap welded. The chemical compositions of the two parent materials were determined using a spectrometer. The chemical compo- sition of the aluminium sheet is as follows: 0.058 % mass fraction of Si, 0.481 % mass fraction of Fe, 0.011 % mass fraction of Ga, 0.05 % mass fraction of other elements and the rest is Al. The chemical com- position of the copper sheet is: 0.137 % mass fraction of Zn, <0.1 % mass fractions of Pb, 0.02 % mass fraction of Ni, 0.023 % mass fraction of Al, 0.012 % mass fraction of Co, 0.077 % mass fraction of B, 0.036 % mass fraction of Sb, 0.043 % mass fraction of Nb, <0.492 % mass fraction of other elements and the rest is Cu. The sheets were friction-stir spot welded in a 30 mm overlap configuration. The spot welds were produced at the eNtsa of Nelson Mandela Metropolitan University, Port Elizabeth, South Africa using an MTS PDS I-Stir. The tool material used was H13 tool steel hardened to 50–52 HRC with a 4 mm tool pin, 5 mm tool diameter and 15 mm tool shoulder. The friction-stir spot welds were produced at rotational speeds of 800 min–1 and 1200 min–1, the tool-shoulder-plunge depths employed were 0.5 mm and 1 mm at a constant dwell time of 10 s. The two different tool profiles used in the current study were flat pin/flat shoulder and conical pin/concave shoulder tool, designated as FPS and CCS, respectively. The produced welds were designated as XX_XX_XX with the first part describing the tool geometry, the second part indicating the rotational speed and the third part indicating the shoulder plunge depth. The weld samples were sectioned using wire elec- trical discharge machining (WEDM), grinded and polished, mounted and prepared, using the ASTM standard metallographic procedure and ASTM Standard E3-11.20 A solution of FeCl3 (10g) + HCl (6 mL) + ethanol (C2H5OH) (20 mL) + H2O (80 mL) was used to etch the copper side of the spot welds, while the aluminium side was etched with H2O (190 mL) + HNO3 (5 mL) + HCl (10 mL) + HF (2 mL). The microstructure of the spot welds was studied using an optical microscope (Olympus BX51M) equipped with the Stream software. Scanning electron microscopy combined with energy dispersive spectroscopy (SEM/EDS) was used to further examine the microstructure and the chemical analyses, respec- tively. A TESCAN equipped with Oxford Instruments X-Max was used for the SEM/EDS analyses. The Vickers microhardness was determined using a dia- M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... 862 Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 2: Representation of a spot weld with dashed lines illustrating the location of the microhardness-profile measurements mond-pyramid-indenter EMCO Test DuraScan tester. Two different locations on the spot welds were used, viz., the top and the bottom, as illustrated in Figure 2. The microhardness measurements were carried out from the keyhole for all the different parameters and tool geometries, to find the probability density function (PDF) of each one. The Matlab 2014 software program was used to determine the PDF. 3 RESULTS AND DISCUSSION The microstructure of the friction-stir spot welds was studied using an optical microscope and the results are depicted in Figures 3 and 4. Figure 3 illustrates the microstructures of the spot welds produced at: a) 800 min–1 and b) 1200 min–1, a 1 mm shoulder plunge depth using a flat pin and a flat shoulder tool. On the other hand, Figure 4 shows the microstructure of the spot weld produced at: a) 800 min–1 and b) 1200 min–1, a 0.5 mm shoulder plunge depth using a conical pin and a concave shoulder tool. It can be seen in Figure 3a that there are a copper ring and a mixture of Al/Cu particles in the stir zone. There is no palpable welding defect in the weld and copper is disseminated in this zone with different shapes. In the upper part of the joint, a large bulk of copper with irregular shapes can be observed (Figure 3a). The tool pin was inserted in the aluminium plate and the copper ring extruded upward from the lower copper plate into the aluminium plate was observed. This was in agree- ment with the previous work.15 Additionally, the intermixing of copper and alumi- nium was not homogenous for different spot welds and different microstructures were formed in different regions of the welds. It was reported that the FSW of dissimilar materials is different from that of similar materials due to the formation of a complex, intercalated vortex-like and related flow pattern.21 In Figure 4b, a good interlaced structure can be seen. This is formed by aluminium and copper, thereby indicating that the two plates are bonded firmly in this region, which is composed of a lamellar structure of copper particles with a streamlined shape of aluminium strips. In this region, a few disseminated copper particles were also observed. The energy-dispersive-spectrometry analyses at the selected points in the stir zone were recorded using SEM micrographs (Table 1). Intermetallic compounds were found in most of the produced welds. Two intermetallic compounds, viz., Al2Cu and Al3Cu4 were found in the weld produced at 1200 min–1, 0.5 mm shoulder plunge M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 863 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 4: Optical microscope images showing the macrostructures of the joints at: a) 800 min–1 and b) 1200 min–1, 0.5 mm shoulder plunge depth using a conical pin and concave shoulder tool Figure 3: Optical microscope images showing the macrostructures of the joints at: a) 800 min–1 and b) 1200 min–1, 1 mm shoulder plunge depth using a flat pin and flat shoulder tool M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... 864 Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Table 1 displays SEM micrographs and the EDS analysis of selected points on the produced friction-spot welds: a) (FFS_1200_0.5), b) (FFS_1200_1), c) (CCS_1200_0.5) and d) (CCS_1200_1). Point Composition Intermetallic compoundAl Cu 1 2 3 4 5 54.94 48.5 4.34 24.55 41.98 45.06 51.5 95.66 75.45 58.02 - Al2Cu - Al3Cu4 Al2Cu Point Composition Intermetallic compoundAl Cu 1 2 3 4 5 1.46 17.71 36.27 94.53 16.98 98.54 82.29 63.73 5.07 83.02 - Al4Cu9 AlCu - Al4Cu9 Point Composition Intermetallic compoundAl Cu 1 2 3 4 5 3.1 7.82 49.49 29.22 13.77 96.9 92.18 50.27 70.78 86.23 - - Al2Cu AlCu AlCu3 Point Composition Intermetallic compoundAl Cu 1 2 3 4 5 2.26 1.33 92.77 37.35 9.15 97.74 98.67 6.97 62.65 90.85 - - - AlCu - depth, whereas the weld produced at 1200 min–1 and 1 mm shoulder plunge depth contained Al4Cu9 and AlCu intermetallic compounds. These intermetallics were found in the welds produced using a flat pin and flat shoulder tool. On the other hand, the spot welds pro- duced using a conical pin and concave shoulder con- tained Al2Cu, AlCu3 and AlCu intermetallic compounds. These intermetallic compounds were found in the weld produced at 1200 min–1 and 0.5 mm shoulder plunge depth. Only the AlCu intermetallic was found in the weld produced at 1200 min–1 and 1 mm shoulder plunge depth, but the concentration of this intermetallic com- pound was relatively small. It was reported that the presence of intermetallic compounds could affect the microhardness profile.17 Figure 5 (a, b, c and d) shows the microhardness values of the spot welds produced using a flat pin and a flat shoulder tool, or a conical pin and a concave shoulder at different process parameters. The micro- hardness values of the parent materials were in the range of 86.7–96.3 HV for Cu while for Al, the range was between 34.6–40.3 HV. In all the samples, high micro- hardness values were recorded at the top, in the region close to the keyhole. It was reported that all the mechanical tests are subject to large statistical variations, which should be evaluated.22 The probability distribution function (PDF) of the Vickers hardness was reported in the literature to correspond to the Gaussian (or normal) distribution23 and log-normal distribution.24 A. M. Hassan et al.25 studied the significance of the process parameters of friction-stir welding of aluminium-matrix composites to set the optimum level for each of these parameters and to further predict which responses are affected when using analyses of variance (ANOVA).25 The present study used the Matlab 2014a statistical toolbox to analyse the pro- bability density function (PDF) of the obtained micro- hardness results. This was done to understand how different parameters and tool geometries affect the probability of obtaining specific microhardness values. A probability density function (PDF) is a function that defines the relative possibility for a random variable to take on a given value. The probability of the random variable falling within a particular range [a, b] of values is given with a finite integral of the PDF within that range [a, b], Equation (1): p x f x e b a x b a ( ) ( , , ) ( ) = =∫ ∫ − −μ    1 2 2 2 π (1) M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 865 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: Microhardness distributions along the welds produced using different tools and process parameters: a) flat pin/flat shoulder (FPS), top; b) bottom; c) conical pin/concave shoulder (CCS), top; d) bottom M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... 864 Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Table 1 displays SEM micrographs and the EDS analysis of selected points on the produced friction-spot welds: a) (FFS_1200_0.5), b) (FFS_1200_1), c) (CCS_1200_0.5) and d) (CCS_1200_1). Point Composition Intermetallic compoundAl Cu 1 2 3 4 5 54.94 48.5 4.34 24.55 41.98 45.06 51.5 95.66 75.45 58.02 - Al2Cu - Al3Cu4 Al2Cu Point Composition Intermetallic compoundAl Cu 1 2 3 4 5 1.46 17.71 36.27 94.53 16.98 98.54 82.29 63.73 5.07 83.02 - Al4Cu9 AlCu - Al4Cu9 Point Composition Intermetallic compoundAl Cu 1 2 3 4 5 3.1 7.82 49.49 29.22 13.77 96.9 92.18 50.27 70.78 86.23 - - Al2Cu AlCu AlCu3 Point Composition Intermetallic compoundAl Cu 1 2 3 4 5 2.26 1.33 92.77 37.35 9.15 97.74 98.67 6.97 62.65 90.85 - - - AlCu - depth, whereas the weld produced at 1200 min–1 and 1 mm shoulder plunge depth contained Al4Cu9 and AlCu intermetallic compounds. These intermetallics were found in the welds produced using a flat pin and flat shoulder tool. On the other hand, the spot welds pro- duced using a conical pin and concave shoulder con- tained Al2Cu, AlCu3 and AlCu intermetallic compounds. These intermetallic compounds were found in the weld produced at 1200 min–1 and 0.5 mm shoulder plunge depth. Only the AlCu intermetallic was found in the weld produced at 1200 min–1 and 1 mm shoulder plunge depth, but the concentration of this intermetallic com- pound was relatively small. It was reported that the presence of intermetallic compounds could affect the microhardness profile.17 Figure 5 (a, b, c and d) shows the microhardness values of the spot welds produced using a flat pin and a flat shoulder tool, or a conical pin and a concave shoulder at different process parameters. The micro- hardness values of the parent materials were in the range of 86.7–96.3 HV for Cu while for Al, the range was between 34.6–40.3 HV. In all the samples, high micro- hardness values were recorded at the top, in the region close to the keyhole. It was reported that all the mechanical tests are subject to large statistical variations, which should be evaluated.22 The probability distribution function (PDF) of the Vickers hardness was reported in the literature to correspond to the Gaussian (or normal) distribution23 and log-normal distribution.24 A. M. Hassan et al.25 studied the significance of the process parameters of friction-stir welding of aluminium-matrix composites to set the optimum level for each of these parameters and to further predict which responses are affected when using analyses of variance (ANOVA).25 The present study used the Matlab 2014a statistical toolbox to analyse the pro- bability density function (PDF) of the obtained micro- hardness results. This was done to understand how different parameters and tool geometries affect the probability of obtaining specific microhardness values. A probability density function (PDF) is a function that defines the relative possibility for a random variable to take on a given value. The probability of the random variable falling within a particular range [a, b] of values is given with a finite integral of the PDF within that range [a, b], Equation (1): p x f x e b a x b a ( ) ( , , ) ( ) = =∫ ∫ − −μ    1 2 2 2 π (1) M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 865 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 5: Microhardness distributions along the welds produced using different tools and process parameters: a) flat pin/flat shoulder (FPS), top; b) bottom; c) conical pin/concave shoulder (CCS), top; d) bottom where  is the standard deviation, 2 is the variance and μ is the mean. It is given by the area under the density function, nonetheless above the horizontal axis and in between the lowest and highest values of the range. The probability density function is a non-negative value and its integral over the entire space is equal to one. PDF histograms of the microhardness and their fits for the parent materials, namely, aluminium and copper, are depicted in Figures 6a and 7b, respectively. The PDFs of the top and bottom hardness measurements were investigated, Figures 7a and 7b depict the PDF histograms of the microhardness for the weld produced at 800 min–1, 0.5 mm shoulder plunge depth, using a flat pin and a flat shoulder tool. It can be seen that the probability of having the microhardness values between 40 and 45 HV is high in the histogram, as shown in Figure 7a, which represents the microhardness measured at the top of the spot weld. This corresponds mostly to the microhardness of the aluminium parent material, whereas the possibility of getting high microhardness values between 50 and 60 HV is low. On the other hand, the PDF of the bottom measu- rement (Figure 7b) shows that there is a high possibility of getting microhardness values between 85 and 95 HV. This corresponds to the microhardness of the copper- parent sheets, while the microhardness values between M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... 866 Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: Depicting the microhardness PDF histograms of the parent materials: a) aluminium, b) copper Figure 8: PDF histogram of the microhardness of CCS_800_0.5 spot weld: a) top and b) bottom Figure 7: PDF histograms of the microhardness of FPS_800_0.5 spot weld: a) top and b) bottom 100 HV and 105 HV are likely to show a lower possi- bility of being obtained when using the same process parameters as those used in this research work. The possibility of having higher microhardness values compared to the values of the parent materials in the two different sheets (copper and aluminium) was observed to be due to the presence of a mixture of copper and aluminium in the vicinity of the keyhole. Additionally, Figure 8 depicts a PDF histogram of the microhardness measurements ((a) top and (b) bottom) for the weld produced at 800 min–1, 0.5 mm shoulder plunge depth, using a conical pin and a concave shoulder. The results show that there is a higher possibility for obtaining microhardness values between 30 50 HV and 50 HV, whereas the possibility for high values between 120 and 130 HV is low (Figure 8a). The trend for the bottom area (Figure 8b) is similar to the one discussed above for the PDF of the bottom microhardness obtained using a flat pin and a flat shoulder. Moreover, when the rotational speed of 1200 min–1 is increased, the possibility of getting high microhardness values ranging between 100–110 HV and 90–100 HV increases for the top area of the spot weld produced using a conical pin and a concave shoulder, and for the top and bottom microhardness values, respectively. It can be seen that the rotational speed and the tool geometry may influence the possibility of different probability distributions. The model shows that, in order to get the probability in a specific region, the integral of the PDF for the region of interest need to be computed. The PDF found in the current research work is a normal distribution (called a Gaussian distribution as well). In order to get any pro- bability, we can compute the finite integral of the normal distribution equation (1). The goodness of fit and the residuals were also analysed. The results show that most of the R2 values range between 0.8842 and 0.9999, which is an indication of how well the model fits with the experimental data shown in Table 2. Table 2: R2 and adjusted R2 of the welds produced using a flat pin/flat shoulder tool and conical pin/concave shoulder tool for the micro- hardness measured on the top and at the bottom Sample ID R2 square R2 square FPS_800_0.5 0.9896 0.8842 FPS_800_1 0.9996 0.9999 FPS_1200_0.5 0.9999 0.9937 FPS_1200_1 0.9993 0.9298 CCS_800_0.5 0.9924 0.9818 CCS_800_1 0.9976 0.9424 CCS_1200_0.5 0.9893 0.987 CCS_1200_1 0.9997 0.9866 Figures 9a to 9b depict the goodness of fit and the residuals for the weld produced at 1200 min–1, 1 mm shoulder plunge depth, using a flat pin and a flat shoulder (the top-microhardness measurement). In M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 867 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 9: a) goodness of fit and b) the residuals for the spot weld produced at 1200 min–1, 1 mm shoulder plunge depth, using a flat pin and a flat shoulder (top-microhardness measurements) Figure 10: a) goodness of fit and b) the residuals for the spot weld produced at 1200 min–1, 1 mm shoulder plunge depth, using a conical pin and a concave shoulder (top-microhardness measurements) where  is the standard deviation, 2 is the variance and μ is the mean. It is given by the area under the density function, nonetheless above the horizontal axis and in between the lowest and highest values of the range. The probability density function is a non-negative value and its integral over the entire space is equal to one. PDF histograms of the microhardness and their fits for the parent materials, namely, aluminium and copper, are depicted in Figures 6a and 7b, respectively. The PDFs of the top and bottom hardness measurements were investigated, Figures 7a and 7b depict the PDF histograms of the microhardness for the weld produced at 800 min–1, 0.5 mm shoulder plunge depth, using a flat pin and a flat shoulder tool. It can be seen that the probability of having the microhardness values between 40 and 45 HV is high in the histogram, as shown in Figure 7a, which represents the microhardness measured at the top of the spot weld. This corresponds mostly to the microhardness of the aluminium parent material, whereas the possibility of getting high microhardness values between 50 and 60 HV is low. On the other hand, the PDF of the bottom measu- rement (Figure 7b) shows that there is a high possibility of getting microhardness values between 85 and 95 HV. This corresponds to the microhardness of the copper- parent sheets, while the microhardness values between M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... 866 Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: Depicting the microhardness PDF histograms of the parent materials: a) aluminium, b) copper Figure 8: PDF histogram of the microhardness of CCS_800_0.5 spot weld: a) top and b) bottom Figure 7: PDF histograms of the microhardness of FPS_800_0.5 spot weld: a) top and b) bottom 100 HV and 105 HV are likely to show a lower possi- bility of being obtained when using the same process parameters as those used in this research work. The possibility of having higher microhardness values compared to the values of the parent materials in the two different sheets (copper and aluminium) was observed to be due to the presence of a mixture of copper and aluminium in the vicinity of the keyhole. Additionally, Figure 8 depicts a PDF histogram of the microhardness measurements ((a) top and (b) bottom) for the weld produced at 800 min–1, 0.5 mm shoulder plunge depth, using a conical pin and a concave shoulder. The results show that there is a higher possibility for obtaining microhardness values between 30 50 HV and 50 HV, whereas the possibility for high values between 120 and 130 HV is low (Figure 8a). The trend for the bottom area (Figure 8b) is similar to the one discussed above for the PDF of the bottom microhardness obtained using a flat pin and a flat shoulder. Moreover, when the rotational speed of 1200 min–1 is increased, the possibility of getting high microhardness values ranging between 100–110 HV and 90–100 HV increases for the top area of the spot weld produced using a conical pin and a concave shoulder, and for the top and bottom microhardness values, respectively. It can be seen that the rotational speed and the tool geometry may influence the possibility of different probability distributions. The model shows that, in order to get the probability in a specific region, the integral of the PDF for the region of interest need to be computed. The PDF found in the current research work is a normal distribution (called a Gaussian distribution as well). In order to get any pro- bability, we can compute the finite integral of the normal distribution equation (1). The goodness of fit and the residuals were also analysed. The results show that most of the R2 values range between 0.8842 and 0.9999, which is an indication of how well the model fits with the experimental data shown in Table 2. Table 2: R2 and adjusted R2 of the welds produced using a flat pin/flat shoulder tool and conical pin/concave shoulder tool for the micro- hardness measured on the top and at the bottom Sample ID R2 square R2 square FPS_800_0.5 0.9896 0.8842 FPS_800_1 0.9996 0.9999 FPS_1200_0.5 0.9999 0.9937 FPS_1200_1 0.9993 0.9298 CCS_800_0.5 0.9924 0.9818 CCS_800_1 0.9976 0.9424 CCS_1200_0.5 0.9893 0.987 CCS_1200_1 0.9997 0.9866 Figures 9a to 9b depict the goodness of fit and the residuals for the weld produced at 1200 min–1, 1 mm shoulder plunge depth, using a flat pin and a flat shoulder (the top-microhardness measurement). In M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 867 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 9: a) goodness of fit and b) the residuals for the spot weld produced at 1200 min–1, 1 mm shoulder plunge depth, using a flat pin and a flat shoulder (top-microhardness measurements) Figure 10: a) goodness of fit and b) the residuals for the spot weld produced at 1200 min–1, 1 mm shoulder plunge depth, using a conical pin and a concave shoulder (top-microhardness measurements) addition, Figures 10a to 10b depict the goodness of fit and the residuals for the weld produced at 1200 min–1, 1 mm shoulder plunge depth, using a conical pin and a concave shoulder (the top-microhardness measurement). Besides the R2 values, the residual analysis was also employed in the study in order to check the adequacy of the models. Figures 9b and 10b show the residual plots for the microhardness values obtained at the top, using a flat pin/flat shoulder and a conical pin/concave shoulder, using the 1200 min–1 speed and 1 mm shoulder plunge depth, respectively. It was reported that the tendencies to have runs of positive and negative residuals indicate the existence of a certain correlation with the experimen- tation.26 Tables 3 and 4 present the standard deviation, the variance and the mean obtained from the statistical analyses of the measured microhardness values for the two different positions, namely, the top and bottom of the spot welds produced with different tools and using different process parameters. Table 3: Mean, variance, mu and sigma of the spot samples produced using a flat pin/flat shoulder tool and a conical pin/concave shoulder tool for the microhardness taken on the top Sample ID Mean 2 μ  FPS_800_0.5 46.543 82.8442 46.5429 9.102 FPS_800_1 75.099 4994.97 4.00167 0.796 FPS_1200_0.5 45.121 441.806 3.71115 0.443 FPS_1200_1 51.731 672.845 3.83391 0.474 CCS_800_0.5 44.575 409.802 3.70341 0.433 CCS_800_1 55.887 813.763 3.90757 0.481 CCS_1200_0.5 45.453 482.889 3.71166 0.458 CCS_1200_1 40.942 248.04 3.64315 0.371 Table 4: Mean, variance, mu and sigma of the spot samples produced using a flat pin/flat shoulder tool and a conical pin/concave shoulder tool for the microhardness taken at the bottom Sample ID Mean 2 μ  FPS_800_0.5 92.234 32.7196 4.52241 0.062 FPS_800_1 84.7 119.974 84.7 10.95 FPS_1200_0.5 84.7 119.974 84.7 10.95 FPS_1200_1 78.429 58.0637 78.4286 7.62 CCS_800_0.5 89.986 14.2859 89.9857 3.78 CCS_800_1 88.564 102.307 88.5643 10.11 CCS_1200_0.5 86.252 89.6604 4.45128 0.109 CCS_1200_1 82.043 54.8519 4.40318 0.09 It should be noted that the mean is equal to μ if the distribution is normal. It can be seen that in some cases in the current work μ and the mean have different values, which shows that the distributions in some of the analyses were not normal. The standard deviation  should be close to zero, but in the current work, the value of the standard deviation is not close to zero in some of the cases. This shows that the microhardness values are not close to the expected values and this could be due to the microhardness values measured in different locations of the weld samples, far apart from each other. This was further suspected due to the presence of intermetallics, which could have been the cause of the high microhardness values since intermetallics are invariably hard and brittle. Each figure contains residuals versus the distance (the distance from the keyhole), taking into account the data and the constant variance of the residuals. In the plot of residuals versus distance, it is shown that the models are adequate to predict the responses in an acceptable manner. 4 CONCLUSIONS According to the presented results, some conclusions can be drawn: • The microstructure of the produced spot welds shows a good material mixing, the presence of a copper ring and a mixture of Al/Cu particles present in the stir zone. • The EDS analyses of the produced friction-stir spot welds exhibited the presence of intermetallic com- pounds, which are known to affect the microhard- ness. • The microhardness values obtained at the top were high for all the samples and this was in the region close to the spot-weld keyhole. In addition, all the microhardness values obtained at the bottom of the samples, in the region close to the keyhole, have lower values, which were close to the average value of the copper base material. This occurred for all the spot welds produced using a conical pin and concave shoulder. • The probability-density-function (PDF) histograms of the microhardness results revealed that the process parameters and the tool geometries have significant effects on the distribution of the microhardness values in different locations of the produced spot welds. • Additionally, goodness-of-fit values were also analysed and these showed that most of the R2 values ranged between 0.8842 and 0.9999, which is an indication of how well the model fits with the produced experimental data. Acknowledgements The authors wish to acknowledge the financial support of the University of Johannesburg and the assistance from Mr. Riaan Brown (eNtsa at Nelson Mandela Metropolitan University). 5 REFERENCES 1 W. M. Thomas, E. D. Nicholas, J. C. Needham, M. G. Murch, P. Temple-Smith, C. J. Dawes, International Patent No. PCT/GB92/ 02203, GB patent application No. 9125978.8, 1991 2 H. Badarinarayan, Fundamentals of Friction Stir Spot Welding, PhD Thesis, 2009, Missouri M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... 868 Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 3 W. Yuan, R. S. Mishra, B. Carlson, R. Verma, R. K. Mishra, Material flow and microstructural evolution during friction stir spot welding of AZ31 magnesium alloy, Materials Science and Engineering, A 543 (2012), 200–209 4 J. Podr`aj, B. Jerman, D. Klob~ar, Welding defects at friction stir welding, Metalurgija, 54 (2015) 2, 387–389 5 J. M. Timothy, Friction Stir Welding of Commercially Available Superplastic Aluminium, PhD thesis, 2008, Department of Engi- neering and Design, Brunel University, Brunel 6 J. Ouyang, E. Yarrapareddy, R. Kovacevic, Microstructural evolution in the friction stir welded 6061 aluminum alloy (T6-temper condi- tion) to copper, Journal of Materials Processing Technology, 172 (2006), 110–122 7 P. Liu, Q. Shi, W. Wang, X. Wang, Z. Zhang, Microstructure and XRD analysis of FSW joints for copper T2/aluminium 5A06 dissimilar materials, Materials Letters, 62 (2008), 4106–4108 8 P. Xue, B. L. Xiao, D. R. Ni, Z. Y. Ma, Enhanced mechanical pro- perties of friction stir welded dissimilar Al–Cu joint by intermetallic compounds, Materials Science and Engineering A, 527 (2010), 5723–5727 9 A. Esmaeili, M. K. Besharati Givi, H. R. Zareie Rajani, A metallur- gical and mechanical study on dissimilar friction stir welding of aluminum 1050 to brass (CuZn30), Materials Science and Engi- neering, A 528 (2011), 7093–7102 10 A. Abdollah-Zadeh, T. Saeid, B. Sazgari, Microstructural and mecha- nical properties of friction stir welded aluminum/copper lap joints, Journal of Alloys and Compounds, 460 (2008), 535–538 11 E. T. Akinlabi, Characterisation of Dissimilar Friction Stir Welds between 5754 Aluminium Alloy and C11000 Copper, D-Tech thesis, Nelson Mandela Metropolitan University, South Africa, 2010 12 M. N. Avettand-Fenoel, R. Taillard, G. Ji, D. Goran, Multiscale Study of Interfacial Intermetallic Compounds in a Dissimilar Al 6082-T6/Cu Friction-Stir Weld, Metallurgical and Materials Transactions A, (2012), 4655–4666 13 M. F. X Muthu, V. Jayabalan, Tool Travel Speed Effects on the Microstructure of Friction Stir Welded Aluminium – Copper Joints, Journal of Materials Processing Technology, 217 (2015), 105–113 14 M. P. Mubiayi, E. T Akinlabi, Friction Stir Spot Welding between Copper and Aluminium: Microstructural Evolution, Proceedings of the International MultiConference of Engineers and Computer Scientists, Vol II, IMECS 2015, March 18–20, 2015, Hong Kong 15 R. Heideman, C. Johnson, S. Kou, Metallurgical analysis of Al/Cu friction stir spot welding, Science and Technology of Welding and Joining, 15 (2010) 7, 597–604 16 U. Özdemir, S. Sayer, Ç. Yeni, Bornova-Izmir, Effect of Pin Pene- tration Depth on the Mechanical Properties of Friction Stir Spot Welded Aluminum and Copper, Materials Testing in Joining Techno- logy, 54 (2012) 4, 233–239 17 M. Shiraly, M. Shamanian, M. R. Toroghinejad, M. Ahmadi Jazani, Effect of Tool Rotation Rate on Microstructure and Mechanical Behavior of Friction Stir Spot-Welded Al/Cu Composite, Journal of Materials Engineering and Performance, 23 (2014) 2, 413–420 18 M. P. Mubiayi, E. T. Akinlabi, Evolving properties of friction stir spot welds between AA1060 and commercially pure copper C11000, Transactions of Nonferrous Metals Society of China, 26 (2016), 1852–1862 19 E. T Akinlabi, S. A. Akinlabi, Friction stir welding of dissimilar materials – statistical analysis of the weld data, Proceedings of the International MultiConference of Engineers and Computer Scientists, 2012, 1368–1373 20 ASTM Standard E3-11, Standard guide for preparation of metallographic specimens, ASTM International, West Consho- hocken, PA, 2011, doi: 10.1520/E0003-11, www.astm.org 21 L. E Murr, A Review of FSW Research on Dissimilar Metal and Alloy Systems, Journal of Materials Engineering and Performance, 8 (2010)19, 1071–108 22 J. M. Schneider, M. Bigerelle, A. Iost, Statistical analysis of the Vickers hardness, Materials Science and Engineering, A 262 (1999), 256–263 23 A. L. Yurkov, N. V. Jhuravleva, E. S. Lukin, Kinetic microhardness measurements of sialon-based ceramics, Journal of Materials Science, 29 (1994), 6551–6560 24 I. Y. Yanchev, E. P. Trifonova, Analysis of microhardness data in Tlx ln l-x Se, Journal of Materials Science, 30 (1995), 5576–5580 25 A. M. Hassan, M. Almomani, T. Qasim, A. Ghaithan, Statistical analysis of some mechanical properties of friction stir welded aluminium matrix composite, Int. J. Experimental Design and Process Optimisation, 3 (2012)1, 91–109 26 K. Palanikumar, R. Karthikeyan, Optimal machining conditions for turning of particulate metal matrix composites using Taguchi and response surface methodologies, Machining Science and Technology, 10 (2006) 4, 417–433 M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 869 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS addition, Figures 10a to 10b depict the goodness of fit and the residuals for the weld produced at 1200 min–1, 1 mm shoulder plunge depth, using a conical pin and a concave shoulder (the top-microhardness measurement). Besides the R2 values, the residual analysis was also employed in the study in order to check the adequacy of the models. Figures 9b and 10b show the residual plots for the microhardness values obtained at the top, using a flat pin/flat shoulder and a conical pin/concave shoulder, using the 1200 min–1 speed and 1 mm shoulder plunge depth, respectively. It was reported that the tendencies to have runs of positive and negative residuals indicate the existence of a certain correlation with the experimen- tation.26 Tables 3 and 4 present the standard deviation, the variance and the mean obtained from the statistical analyses of the measured microhardness values for the two different positions, namely, the top and bottom of the spot welds produced with different tools and using different process parameters. Table 3: Mean, variance, mu and sigma of the spot samples produced using a flat pin/flat shoulder tool and a conical pin/concave shoulder tool for the microhardness taken on the top Sample ID Mean 2 μ  FPS_800_0.5 46.543 82.8442 46.5429 9.102 FPS_800_1 75.099 4994.97 4.00167 0.796 FPS_1200_0.5 45.121 441.806 3.71115 0.443 FPS_1200_1 51.731 672.845 3.83391 0.474 CCS_800_0.5 44.575 409.802 3.70341 0.433 CCS_800_1 55.887 813.763 3.90757 0.481 CCS_1200_0.5 45.453 482.889 3.71166 0.458 CCS_1200_1 40.942 248.04 3.64315 0.371 Table 4: Mean, variance, mu and sigma of the spot samples produced using a flat pin/flat shoulder tool and a conical pin/concave shoulder tool for the microhardness taken at the bottom Sample ID Mean 2 μ  FPS_800_0.5 92.234 32.7196 4.52241 0.062 FPS_800_1 84.7 119.974 84.7 10.95 FPS_1200_0.5 84.7 119.974 84.7 10.95 FPS_1200_1 78.429 58.0637 78.4286 7.62 CCS_800_0.5 89.986 14.2859 89.9857 3.78 CCS_800_1 88.564 102.307 88.5643 10.11 CCS_1200_0.5 86.252 89.6604 4.45128 0.109 CCS_1200_1 82.043 54.8519 4.40318 0.09 It should be noted that the mean is equal to μ if the distribution is normal. It can be seen that in some cases in the current work μ and the mean have different values, which shows that the distributions in some of the analyses were not normal. The standard deviation  should be close to zero, but in the current work, the value of the standard deviation is not close to zero in some of the cases. This shows that the microhardness values are not close to the expected values and this could be due to the microhardness values measured in different locations of the weld samples, far apart from each other. This was further suspected due to the presence of intermetallics, which could have been the cause of the high microhardness values since intermetallics are invariably hard and brittle. Each figure contains residuals versus the distance (the distance from the keyhole), taking into account the data and the constant variance of the residuals. In the plot of residuals versus distance, it is shown that the models are adequate to predict the responses in an acceptable manner. 4 CONCLUSIONS According to the presented results, some conclusions can be drawn: • The microstructure of the produced spot welds shows a good material mixing, the presence of a copper ring and a mixture of Al/Cu particles present in the stir zone. • The EDS analyses of the produced friction-stir spot welds exhibited the presence of intermetallic com- pounds, which are known to affect the microhard- ness. • The microhardness values obtained at the top were high for all the samples and this was in the region close to the spot-weld keyhole. In addition, all the microhardness values obtained at the bottom of the samples, in the region close to the keyhole, have lower values, which were close to the average value of the copper base material. This occurred for all the spot welds produced using a conical pin and concave shoulder. • The probability-density-function (PDF) histograms of the microhardness results revealed that the process parameters and the tool geometries have significant effects on the distribution of the microhardness values in different locations of the produced spot welds. • Additionally, goodness-of-fit values were also analysed and these showed that most of the R2 values ranged between 0.8842 and 0.9999, which is an indication of how well the model fits with the produced experimental data. Acknowledgements The authors wish to acknowledge the financial support of the University of Johannesburg and the assistance from Mr. Riaan Brown (eNtsa at Nelson Mandela Metropolitan University). 5 REFERENCES 1 W. M. Thomas, E. D. Nicholas, J. C. Needham, M. G. Murch, P. Temple-Smith, C. J. Dawes, International Patent No. PCT/GB92/ 02203, GB patent application No. 9125978.8, 1991 2 H. Badarinarayan, Fundamentals of Friction Stir Spot Welding, PhD Thesis, 2009, Missouri M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... 868 Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 3 W. Yuan, R. S. Mishra, B. Carlson, R. Verma, R. K. Mishra, Material flow and microstructural evolution during friction stir spot welding of AZ31 magnesium alloy, Materials Science and Engineering, A 543 (2012), 200–209 4 J. Podr`aj, B. Jerman, D. Klob~ar, Welding defects at friction stir welding, Metalurgija, 54 (2015) 2, 387–389 5 J. M. Timothy, Friction Stir Welding of Commercially Available Superplastic Aluminium, PhD thesis, 2008, Department of Engi- neering and Design, Brunel University, Brunel 6 J. Ouyang, E. Yarrapareddy, R. Kovacevic, Microstructural evolution in the friction stir welded 6061 aluminum alloy (T6-temper condi- tion) to copper, Journal of Materials Processing Technology, 172 (2006), 110–122 7 P. Liu, Q. Shi, W. Wang, X. Wang, Z. Zhang, Microstructure and XRD analysis of FSW joints for copper T2/aluminium 5A06 dissimilar materials, Materials Letters, 62 (2008), 4106–4108 8 P. Xue, B. L. Xiao, D. R. Ni, Z. Y. Ma, Enhanced mechanical pro- perties of friction stir welded dissimilar Al–Cu joint by intermetallic compounds, Materials Science and Engineering A, 527 (2010), 5723–5727 9 A. Esmaeili, M. K. Besharati Givi, H. R. Zareie Rajani, A metallur- gical and mechanical study on dissimilar friction stir welding of aluminum 1050 to brass (CuZn30), Materials Science and Engi- neering, A 528 (2011), 7093–7102 10 A. Abdollah-Zadeh, T. Saeid, B. Sazgari, Microstructural and mecha- nical properties of friction stir welded aluminum/copper lap joints, Journal of Alloys and Compounds, 460 (2008), 535–538 11 E. T. Akinlabi, Characterisation of Dissimilar Friction Stir Welds between 5754 Aluminium Alloy and C11000 Copper, D-Tech thesis, Nelson Mandela Metropolitan University, South Africa, 2010 12 M. N. Avettand-Fenoel, R. Taillard, G. Ji, D. Goran, Multiscale Study of Interfacial Intermetallic Compounds in a Dissimilar Al 6082-T6/Cu Friction-Stir Weld, Metallurgical and Materials Transactions A, (2012), 4655–4666 13 M. F. X Muthu, V. Jayabalan, Tool Travel Speed Effects on the Microstructure of Friction Stir Welded Aluminium – Copper Joints, Journal of Materials Processing Technology, 217 (2015), 105–113 14 M. P. Mubiayi, E. T Akinlabi, Friction Stir Spot Welding between Copper and Aluminium: Microstructural Evolution, Proceedings of the International MultiConference of Engineers and Computer Scientists, Vol II, IMECS 2015, March 18–20, 2015, Hong Kong 15 R. Heideman, C. Johnson, S. Kou, Metallurgical analysis of Al/Cu friction stir spot welding, Science and Technology of Welding and Joining, 15 (2010) 7, 597–604 16 U. Özdemir, S. Sayer, Ç. Yeni, Bornova-Izmir, Effect of Pin Pene- tration Depth on the Mechanical Properties of Friction Stir Spot Welded Aluminum and Copper, Materials Testing in Joining Techno- logy, 54 (2012) 4, 233–239 17 M. Shiraly, M. Shamanian, M. R. Toroghinejad, M. Ahmadi Jazani, Effect of Tool Rotation Rate on Microstructure and Mechanical Behavior of Friction Stir Spot-Welded Al/Cu Composite, Journal of Materials Engineering and Performance, 23 (2014) 2, 413–420 18 M. P. Mubiayi, E. T. Akinlabi, Evolving properties of friction stir spot welds between AA1060 and commercially pure copper C11000, Transactions of Nonferrous Metals Society of China, 26 (2016), 1852–1862 19 E. T Akinlabi, S. A. Akinlabi, Friction stir welding of dissimilar materials – statistical analysis of the weld data, Proceedings of the International MultiConference of Engineers and Computer Scientists, 2012, 1368–1373 20 ASTM Standard E3-11, Standard guide for preparation of metallographic specimens, ASTM International, West Consho- hocken, PA, 2011, doi: 10.1520/E0003-11, www.astm.org 21 L. E Murr, A Review of FSW Research on Dissimilar Metal and Alloy Systems, Journal of Materials Engineering and Performance, 8 (2010)19, 1071–108 22 J. M. Schneider, M. Bigerelle, A. Iost, Statistical analysis of the Vickers hardness, Materials Science and Engineering, A 262 (1999), 256–263 23 A. L. Yurkov, N. V. Jhuravleva, E. S. Lukin, Kinetic microhardness measurements of sialon-based ceramics, Journal of Materials Science, 29 (1994), 6551–6560 24 I. Y. Yanchev, E. P. Trifonova, Analysis of microhardness data in Tlx ln l-x Se, Journal of Materials Science, 30 (1995), 5576–5580 25 A. M. Hassan, M. Almomani, T. Qasim, A. Ghaithan, Statistical analysis of some mechanical properties of friction stir welded aluminium matrix composite, Int. J. Experimental Design and Process Optimisation, 3 (2012)1, 91–109 26 K. Palanikumar, R. Karthikeyan, Optimal machining conditions for turning of particulate metal matrix composites using Taguchi and response surface methodologies, Machining Science and Technology, 10 (2006) 4, 417–433 M. P. MUBIAYI et al.: MICROSTRUCTURE EVOLUTION AND STATISTICAL ANALYSIS OF Al/Cu ... Materiali in tehnologije / Materials and technology 51 (2017) 5, 861–869 869 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH 871–878 SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH SINTEZA PMMA/ZnO NANODELCEV KOMPOZITOV ZA IZDELAVO ZOB IZ UMETNIH SMOL Danica Popovi}1, Rajko Bobovnik2, Silvester Bolka2, Miroslav Vukadinovi}1, Vojkan Lazi}1, Rebeka Rudolf3,4 1University of Belgrade, School of Dental Medicine, Dr. Suboti}a 8, 11000 Belgrade, Serbia 2Faculty of Polymer Technology, Ozare 19, 2380 Slovenj Gradec, Slovenia 3Zlatarna Celje d.o.o., Kersnikova ulica 19, 3000 Celje, Slovenia 4University of Maribor, Faculty of Mechanical Engineering, Smetanova 17, 2000 Maribor, Slovenia d.popovic984@gmail.com Prejem rokopisa – received: 2017-02-24; sprejem za objavo – accepted for publication: 2017-03-31 doi:10.17222/mit.2017.025 Wear resistance is one of the most important physical properties of the artificial teeth used in acrylic dentures. The goal of this research was to synthesize a new composite material made of matrix Poly-(methyl methacrylate)-PMMA with different percentages (2 % and 3 % of volume fractions) of zinc-oxide nanoparticles (ZnO NPs) as reinforcing elements, to improve its mechanical properties. The dynamic mechanical behaviour of this composite was studied through the DMA method in comparison to the pure PMMA supported by the characterization of their microstructures. Then the wear resistance was analysed on the samples, which were prepared in the form of teeth. In this context their vertical height loss was measured after 100,000 chewing cycles on a chewing simulator, before and after the artificial thermal ageing. Investigations showed that the PMMA/ZnO NP composites dampened the vibrations better than the pure PMMA, which could be assigned to the homogenous distribution of ZnO NPs in the PMMA matrix. It was found that the mean vertical height loss for the pure PMMA teeth was significantly higher (more than 4 times) compared to composite teeth made with ZnO NPs. Introducing the thermal artificial ageing led to the finding that there was no effect on the height loss by the composite material with 3 % of volume fractions of ZnO NPs. Based on this it was concluded that PMMA/ZnO NPs composites showed improved in-vitro wear resistance compared to acrylic-resin denture teeth, so this new composite material should be preferred when occlusal stability is considered to be of high priority. Keywords: poly-(methyl methacrylate)–PMMA, zinc-oxide nanoparticles, composite, resin teeth Odpornost proti obrabi je ena izmed najbolj pomembnih fizikalnih lastnosti umetnih zob pri akrilnih protezah. Cilj te raziskave je bil sintetiziranje novega kompozitnega materiala, izdelanega iz matrice poli-(metil meta akrilata) – PMMA z razli~nimi volumskimi odstotki (2 % in 3 %) nanodelcev cinkovega oksida (ZnO NPs), uporabljenimi kot oja~itveni element za izbolj{anje mehanskih lastnosti materiala. Dinami~no mehansko obna{anje teh kompozitov smo raziskali z metodo DMA v primerjavi s ~istim PMMA, ter s karakterizacijo njihovih mikrostruktur. Nato je bila analizirana odpornost proti obrabi na vzorcih, ki so bili pripravljeni v obliki zob. Izmerjena je bila njihova navpi~na izguba vi{ine po 100.000 `ve~ilnih ciklih na `ve~ilnem simulatorju, pred in po umetnem toplotnem staranju. Preiskave so pokazale, da so PMMA/ZnO NPs kompoziti bolje bla`ili vibracije kot ~isti PMMA, kar lahko pripi{emo homogeni porazdelitvi ZnO NPs v PMMA matrici. Ugotovili smo, da je bila povpre~na navpi~na izguba vi{ina za ~isti PMMA zob znatno vi{ja (ve~ kot 4×) v primerjavi s kompozitnimi zobmi z ZnO NPs. Umetno toplotno staranje ni imelo u~inka na izgubo navpi~ne vi{ine zoba iz kompozitnega materiala s 3 volumskimi % ZnO NPs. Na podlagi tega smo sklenili, da imajo kompoziti PMMA/ZnO NPs izbolj{ano in vitro odpornost proti obrabi, v primerjavi z zobno protezo iz akrilne smole, zato naj bi imel ta novi kompozitni material prednost pri uporabi, kadar je potrebna okluzalna stabilnost. Klju~ne besede: poli-(metil meta akrilat)-PMMA, nanodelci cinkovega oksida, kompoziti, zobje iz umetnih smol 1 INTRODUCTION Teeth loss in the human population is associated with numerous problems such as chewing difficulties, alteration of speech and facial expressions, which all leads to de-socialization. According to the Oral Health-Healthy People 2010: Objectives for Improving Health 26 % of the US population and 33 % of the Central Europe population aged between 65 years and 74 years are toothless; but some dramatic data and a wide variation in edentulism prevalence among adults aged 50 and above was found in the following countries: 48.3 % in Ireland, Malaysia 56.6 % and Netherlands 65.4 %, which indicates an international problem.1–3 The problems of toothless people in the region of the Balkan peninsula and South-Eastern Europe are usually solved with a complete or partial acrylic denture, which is the easiest and the cheapest solution, because more than 90 % of these patients are in a bad financial situation and incapable of buying dental implants and hybrid bridges above them. With an acrylic denture the wear resistance represents one of the most important physical properties because, in the case that the material has excessive wear, that might cause loss of the vertical dimension of occlusion (VDO) which is associated with decreased occlusal forces, loss of masticatory efficiency, improper tooth relationship, and fatigue of masticatory muscles.4 Recently, a study examined how VDO can Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 871 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 620.3:616.314-77:678.6:67.017 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)871(2017) D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH 871–878 SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH SINTEZA PMMA/ZnO NANODELCEV KOMPOZITOV ZA IZDELAVO ZOB IZ UMETNIH SMOL Danica Popovi}1, Rajko Bobovnik2, Silvester Bolka2, Miroslav Vukadinovi}1, Vojkan Lazi}1, Rebeka Rudolf3,4 1University of Belgrade, School of Dental Medicine, Dr. Suboti}a 8, 11000 Belgrade, Serbia 2Faculty of Polymer Technology, Ozare 19, 2380 Slovenj Gradec, Slovenia 3Zlatarna Celje d.o.o., Kersnikova ulica 19, 3000 Celje, Slovenia 4University of Maribor, Faculty of Mechanical Engineering, Smetanova 17, 2000 Maribor, Slovenia d.popovic984@gmail.com Prejem rokopisa – received: 2017-02-24; sprejem za objavo – accepted for publication: 2017-03-31 doi:10.17222/mit.2017.025 Wear resistance is one of the most important physical properties of the artificial teeth used in acrylic dentures. The goal of this research was to synthesize a new composite material made of matrix Poly-(methyl methacrylate)-PMMA with different percentages (2 % and 3 % of volume fractions) of zinc-oxide nanoparticles (ZnO NPs) as reinforcing elements, to improve its mechanical properties. The dynamic mechanical behaviour of this composite was studied through the DMA method in comparison to the pure PMMA supported by the characterization of their microstructures. Then the wear resistance was analysed on the samples, which were prepared in the form of teeth. In this context their vertical height loss was measured after 100,000 chewing cycles on a chewing simulator, before and after the artificial thermal ageing. Investigations showed that the PMMA/ZnO NP composites dampened the vibrations better than the pure PMMA, which could be assigned to the homogenous distribution of ZnO NPs in the PMMA matrix. It was found that the mean vertical height loss for the pure PMMA teeth was significantly higher (more than 4 times) compared to composite teeth made with ZnO NPs. Introducing the thermal artificial ageing led to the finding that there was no effect on the height loss by the composite material with 3 % of volume fractions of ZnO NPs. Based on this it was concluded that PMMA/ZnO NPs composites showed improved in-vitro wear resistance compared to acrylic-resin denture teeth, so this new composite material should be preferred when occlusal stability is considered to be of high priority. Keywords: poly-(methyl methacrylate)–PMMA, zinc-oxide nanoparticles, composite, resin teeth Odpornost proti obrabi je ena izmed najbolj pomembnih fizikalnih lastnosti umetnih zob pri akrilnih protezah. Cilj te raziskave je bil sintetiziranje novega kompozitnega materiala, izdelanega iz matrice poli-(metil meta akrilata) – PMMA z razli~nimi volumskimi odstotki (2 % in 3 %) nanodelcev cinkovega oksida (ZnO NPs), uporabljenimi kot oja~itveni element za izbolj{anje mehanskih lastnosti materiala. Dinami~no mehansko obna{anje teh kompozitov smo raziskali z metodo DMA v primerjavi s ~istim PMMA, ter s karakterizacijo njihovih mikrostruktur. Nato je bila analizirana odpornost proti obrabi na vzorcih, ki so bili pripravljeni v obliki zob. Izmerjena je bila njihova navpi~na izguba vi{ine po 100.000 `ve~ilnih ciklih na `ve~ilnem simulatorju, pred in po umetnem toplotnem staranju. Preiskave so pokazale, da so PMMA/ZnO NPs kompoziti bolje bla`ili vibracije kot ~isti PMMA, kar lahko pripi{emo homogeni porazdelitvi ZnO NPs v PMMA matrici. Ugotovili smo, da je bila povpre~na navpi~na izguba vi{ina za ~isti PMMA zob znatno vi{ja (ve~ kot 4×) v primerjavi s kompozitnimi zobmi z ZnO NPs. Umetno toplotno staranje ni imelo u~inka na izgubo navpi~ne vi{ine zoba iz kompozitnega materiala s 3 volumskimi % ZnO NPs. Na podlagi tega smo sklenili, da imajo kompoziti PMMA/ZnO NPs izbolj{ano in vitro odpornost proti obrabi, v primerjavi z zobno protezo iz akrilne smole, zato naj bi imel ta novi kompozitni material prednost pri uporabi, kadar je potrebna okluzalna stabilnost. Klju~ne besede: poli-(metil meta akrilat)-PMMA, nanodelci cinkovega oksida, kompoziti, zobje iz umetnih smol 1 INTRODUCTION Teeth loss in the human population is associated with numerous problems such as chewing difficulties, alteration of speech and facial expressions, which all leads to de-socialization. According to the Oral Health-Healthy People 2010: Objectives for Improving Health 26 % of the US population and 33 % of the Central Europe population aged between 65 years and 74 years are toothless; but some dramatic data and a wide variation in edentulism prevalence among adults aged 50 and above was found in the following countries: 48.3 % in Ireland, Malaysia 56.6 % and Netherlands 65.4 %, which indicates an international problem.1–3 The problems of toothless people in the region of the Balkan peninsula and South-Eastern Europe are usually solved with a complete or partial acrylic denture, which is the easiest and the cheapest solution, because more than 90 % of these patients are in a bad financial situation and incapable of buying dental implants and hybrid bridges above them. With an acrylic denture the wear resistance represents one of the most important physical properties because, in the case that the material has excessive wear, that might cause loss of the vertical dimension of occlusion (VDO) which is associated with decreased occlusal forces, loss of masticatory efficiency, improper tooth relationship, and fatigue of masticatory muscles.4 Recently, a study examined how VDO can Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 871 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS UDK 620.3:616.314-77:678.6:67.017 ISSN 1580-2949 Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(5)871(2017) affect brain function in complete denture wearers, and also measured occlusal forces when there is a change in VDO in which electroencephalograms were used.5 Many authors have pointed out the necessity of creating complete new dentures, or relining of the denture base, which is not a simple procedure and presents financial pressure for the patients.6 So, finding the new combi- nation of materials with improved mechanical properties, which include better wear resistance, would be helpful for many reasons for those patients. Many basic researches based on polymer materials represent the challenge. One of these are made of supramolecular polymers as a new high-tech material.7,8 Over the past two decades a great effort has been made by the community of researchers for the synthesis and utilization of nanomaterials in the field of medical materials.9 Using fine nanoparticles as reinforcing elements could improve the functional properties of conventional materials such as PMMA, which has been used for dentures since 1937. The ZnO NPs are small objects that behave as an entire unit regarding their transport and properties. The introduction of ZnO NPs represents a new approach in dentistry by the production of a denture, because they could decrease the level of residual monomer and, besides this, they also have anti- microbial activity. This was shown in recent studies.10,11 Namely, ZnO NPs are environmentally friendly materials in the nanometre size range, and could be synthesized in a wide range of particle shapes and structures.12–14 ZnO in bulk form is a largely inert, white powder compound that has a very broad application, from the chemical industry through cosmetic and medical products. In dentistry they have an important place as a medicament for healing injures after periodontal surgical treatment, in the cementation of temporary, as well as definitive, crowns etc. ZnO is also a biocompatible material that exhibits antimicrobial properties against gram-negative as well as gram-positive bacteria.15 A recent study also investigated, proved and determined the minimal inhibitory concentration of ZnO NPs against Candida Albicans.16 Nanostructured materials with unique and fascinating properties motivate scientists tremendously to explore and understand their formation and growth processes and following the critical volume in several studies we tried to find the right percentage of ZnO NPs in a PMMA matrix which could advance the properties of this new composite, because increasing the percentage of added NPs could downgrade some of them.17,18 According to the presented state-of-the-art study, the aim of this research was: (i) To synthesize a composite PMMA/ZnO NPs material, to test its mechanical properties, and evaluate the developed microstructures; (ii) To investigate the wear resistance of resin teeth in a chewing simulator with two combinations, before and after thermal artificial ageing. This investigation included not only the classical testing samples, but also the real model of the shapes of first upper molars (testing resin teeth) and showed possible practical applications in dentistry. 2 EXPERIMENTAL PART 2.1 Materials for composite PMMA/ZnO NPs Commercially available ZnO nanopowder (purity 99.99 %, density 5.606 g/cm3, insoluble in water, for- mula weight 81.37, melting point 1975 °C) was pur- chased from Interdent (Belgrade, Serbia). The average size of the ZnO NPs was 30 nm and they were in the form of a sphere. The typical shape and morphology of characteristic ZnO NPs are shown in Figure 1, which was made on a transmission electron microscope (TEM). PMMA matrix (polymethylmethacrylate powder with benzoyl-peroxide, pigments and for a liquid-methyl- methacrylate HQ 60, ethylene glycoldimethacrylate, methylmethacrylate composite) was distributed by Galenika (Belgrade, Serbia). This material was used for Biogal® acrylic teeth (Galenika, Belgrade, Serbia) and has already passed through the applied ISO Standards. 2.2 Measurement of ZnO nanoparticles’ size distri- bution and zeta potential in water and MMA Because of the frequent problem with agglomeration among nanoparticles, especially ZnO NPs in water, we used the DLS technique to confirm or deny this fact indirectly and compare them with ZnO NPs in suspen- sion with MMA. Initially, ZnO NPs powder (0.05 g) with water and MMA monomer (15 mL by volume in both cases) was mixed through magnetic stirring for the time duration of 45 min. The size distributions and surface charge of the nanoparticles were determined with a Malvern Zetasizer Nano ZS (Malvern Instruments Ltd., U.K.) by the DLS technique. The suitable parameters for ZnO NPs Absorption: 0.1 and Refractive index: 2.0. were chosen according to 19. D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH 872 Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 1: Micrograph of shape and morphology of characteristic ZnO NPs (TEM, 10 000×) 2.3 Preparation of samples Samples were made in two shapes: Lamina (dimen- sion: 45 mm × 3 mm × 7 mm) for characterization of the new composite and then in the form of the first upper molar for direct potential use in dentistry. The ZnO NPs in the PMMA matrix were mixed with PMMA powder and then free radical chain poly- merization was performed of MMA monomer in bulk. The new composite was synthesized at the selection of the appropriate technological regime, which started in the moment of adding monomer liquid PMMA/ZnO NPs powder for 1 h and 45 min at 100 °C. The procedure for both samples followed by shaping the samples firstly in wax, then duplicated and changed in to the pure PMMA, PMMA filled with 2 % of volume fractions and 3 % of volume fractions of ZnO NPs. The volume % were chosen according to our previous experiences and based on 7,16,18. All samples in the form of the first upper molar were identical in form, plan parallel and highly polished. They were prepared for a wear test in a chewing simulator. Schematic presentations of the composite PMMA/ZnO NPs synthesis and their corresponding testing are shown in Figure 2. A list of samples’ names, their mixing ratio PMMA/ZnO NPs and technical descriptions are given in Table 1. PMMA1 was chosen for a comparison with the new composites. 2.4 Characterization Thermo-mechanical properties were examined using a Perkin Elmer DMA 8000 dynamic mechanical anal- yser. The TT_DMA software, Version 14310, was used for the evaluation of the results. The viscoelastic proper- ties of the samples were analysed by recording the storage modulus (E’), loss modulus (E’’) and loss factor (tan ) as a function of temperature. The height of the peak of the loss factor determined the damping beha- viour – with decreasing peak height the damping beha- viour was increasing. For the analyses of test samples, the DMA instrument was operated in the dual cantilever mode. The viscoelastic analyses were carried out on samples with dimensions of approximately 42 mm × 5 mm × 2 mm. The samples were heated at 2 °C/min from room temperature to 180 °C under an air atmo- sphere. A frequency of 10 Hz and amplitude of 20 μm were used. A flash differential scanning calorimeter (Flash DSC) was used for the identification of the glass transition and reorganisation of the polymers for all the samples. Flash DSC works by ultra heating and cooling rates that induce physical transitions and chemical processes. This charac- terization was very important for dentistry practice because it allowed us to follow the level of polymeri- zation in the residual monomer, which could produce an allergic reaction by the use of the new composite for the resin teeth in removable dentures in the mouths of patients. 2.5 Determination of wear resistance The wear tests of the samples in the shape of upper first molar were performed in a chewing simulator CS-4.2 economy line (SD Mechatronics, Germany). The simulator has two identical sample chambers and two stepper motors which allow computer-controlled vertical and horizontal movements. The masticatory cycle in this research consisted of three phases: contact with a vertical load of 5 kg, horizontal sliding of 0.4 mm, vertical sliding of 0.2 mm and separating the teeth and machine antagonist. In our investigation, two types of tests were made: before and after thermal artificial ageing including all samples from Table 1. The machine was set at 100.000 cycles, which simulated chewing over a period of one year. The loss of substance was measured with an electronic digital calliper – Globathronics GmbH (Ger- many). Accelerated thermal ageing was carried out by immersing the samples in a water bath at a temperature of 70 °C.20,21 D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 873 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Table 1: Composition, mixing ratio and forms of samples for different testing Samples/ abbreviation Composition Mixing ratio Form forcharacterization Form for wear test PMMA1 Pure PMMA 15 mL MMA + 23.4 gPMMA Lamina Tooth model PMMA2 PMMA/2vol%ZnO NPs 15 mL + 22.93 g PMMA+ 0,47 g ZnO NPs Lamina Tooth model PMMA3 PMMA/3vol%ZnO NPs 15 mL + 22.7 g PMMA+ 0,7 g ZnO NPs Lamina Tooth model Figure 2: Schematic presentation of synthesis composite PMMA/ZnO NPs and samples for testing affect brain function in complete denture wearers, and also measured occlusal forces when there is a change in VDO in which electroencephalograms were used.5 Many authors have pointed out the necessity of creating complete new dentures, or relining of the denture base, which is not a simple procedure and presents financial pressure for the patients.6 So, finding the new combi- nation of materials with improved mechanical properties, which include better wear resistance, would be helpful for many reasons for those patients. Many basic researches based on polymer materials represent the challenge. One of these are made of supramolecular polymers as a new high-tech material.7,8 Over the past two decades a great effort has been made by the community of researchers for the synthesis and utilization of nanomaterials in the field of medical materials.9 Using fine nanoparticles as reinforcing elements could improve the functional properties of conventional materials such as PMMA, which has been used for dentures since 1937. The ZnO NPs are small objects that behave as an entire unit regarding their transport and properties. The introduction of ZnO NPs represents a new approach in dentistry by the production of a denture, because they could decrease the level of residual monomer and, besides this, they also have anti- microbial activity. This was shown in recent studies.10,11 Namely, ZnO NPs are environmentally friendly materials in the nanometre size range, and could be synthesized in a wide range of particle shapes and structures.12–14 ZnO in bulk form is a largely inert, white powder compound that has a very broad application, from the chemical industry through cosmetic and medical products. In dentistry they have an important place as a medicament for healing injures after periodontal surgical treatment, in the cementation of temporary, as well as definitive, crowns etc. ZnO is also a biocompatible material that exhibits antimicrobial properties against gram-negative as well as gram-positive bacteria.15 A recent study also investigated, proved and determined the minimal inhibitory concentration of ZnO NPs against Candida Albicans.16 Nanostructured materials with unique and fascinating properties motivate scientists tremendously to explore and understand their formation and growth processes and following the critical volume in several studies we tried to find the right percentage of ZnO NPs in a PMMA matrix which could advance the properties of this new composite, because increasing the percentage of added NPs could downgrade some of them.17,18 According to the presented state-of-the-art study, the aim of this research was: (i) To synthesize a composite PMMA/ZnO NPs material, to test its mechanical properties, and evaluate the developed microstructures; (ii) To investigate the wear resistance of resin teeth in a chewing simulator with two combinations, before and after thermal artificial ageing. This investigation included not only the classical testing samples, but also the real model of the shapes of first upper molars (testing resin teeth) and showed possible practical applications in dentistry. 2 EXPERIMENTAL PART 2.1 Materials for composite PMMA/ZnO NPs Commercially available ZnO nanopowder (purity 99.99 %, density 5.606 g/cm3, insoluble in water, for- mula weight 81.37, melting point 1975 °C) was pur- chased from Interdent (Belgrade, Serbia). The average size of the ZnO NPs was 30 nm and they were in the form of a sphere. The typical shape and morphology of characteristic ZnO NPs are shown in Figure 1, which was made on a transmission electron microscope (TEM). PMMA matrix (polymethylmethacrylate powder with benzoyl-peroxide, pigments and for a liquid-methyl- methacrylate HQ 60, ethylene glycoldimethacrylate, methylmethacrylate composite) was distributed by Galenika (Belgrade, Serbia). This material was used for Biogal® acrylic teeth (Galenika, Belgrade, Serbia) and has already passed through the applied ISO Standards. 2.2 Measurement of ZnO nanoparticles’ size distri- bution and zeta potential in water and MMA Because of the frequent problem with agglomeration among nanoparticles, especially ZnO NPs in water, we used the DLS technique to confirm or deny this fact indirectly and compare them with ZnO NPs in suspen- sion with MMA. Initially, ZnO NPs powder (0.05 g) with water and MMA monomer (15 mL by volume in both cases) was mixed through magnetic stirring for the time duration of 45 min. The size distributions and surface charge of the nanoparticles were determined with a Malvern Zetasizer Nano ZS (Malvern Instruments Ltd., U.K.) by the DLS technique. The suitable parameters for ZnO NPs Absorption: 0.1 and Refractive index: 2.0. were chosen according to 19. D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH 872 Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 1: Micrograph of shape and morphology of characteristic ZnO NPs (TEM, 10 000×) 2.3 Preparation of samples Samples were made in two shapes: Lamina (dimen- sion: 45 mm × 3 mm × 7 mm) for characterization of the new composite and then in the form of the first upper molar for direct potential use in dentistry. The ZnO NPs in the PMMA matrix were mixed with PMMA powder and then free radical chain poly- merization was performed of MMA monomer in bulk. The new composite was synthesized at the selection of the appropriate technological regime, which started in the moment of adding monomer liquid PMMA/ZnO NPs powder for 1 h and 45 min at 100 °C. The procedure for both samples followed by shaping the samples firstly in wax, then duplicated and changed in to the pure PMMA, PMMA filled with 2 % of volume fractions and 3 % of volume fractions of ZnO NPs. The volume % were chosen according to our previous experiences and based on 7,16,18. All samples in the form of the first upper molar were identical in form, plan parallel and highly polished. They were prepared for a wear test in a chewing simulator. Schematic presentations of the composite PMMA/ZnO NPs synthesis and their corresponding testing are shown in Figure 2. A list of samples’ names, their mixing ratio PMMA/ZnO NPs and technical descriptions are given in Table 1. PMMA1 was chosen for a comparison with the new composites. 2.4 Characterization Thermo-mechanical properties were examined using a Perkin Elmer DMA 8000 dynamic mechanical anal- yser. The TT_DMA software, Version 14310, was used for the evaluation of the results. The viscoelastic proper- ties of the samples were analysed by recording the storage modulus (E’), loss modulus (E’’) and loss factor (tan ) as a function of temperature. The height of the peak of the loss factor determined the damping beha- viour – with decreasing peak height the damping beha- viour was increasing. For the analyses of test samples, the DMA instrument was operated in the dual cantilever mode. The viscoelastic analyses were carried out on samples with dimensions of approximately 42 mm × 5 mm × 2 mm. The samples were heated at 2 °C/min from room temperature to 180 °C under an air atmo- sphere. A frequency of 10 Hz and amplitude of 20 μm were used. A flash differential scanning calorimeter (Flash DSC) was used for the identification of the glass transition and reorganisation of the polymers for all the samples. Flash DSC works by ultra heating and cooling rates that induce physical transitions and chemical processes. This charac- terization was very important for dentistry practice because it allowed us to follow the level of polymeri- zation in the residual monomer, which could produce an allergic reaction by the use of the new composite for the resin teeth in removable dentures in the mouths of patients. 2.5 Determination of wear resistance The wear tests of the samples in the shape of upper first molar were performed in a chewing simulator CS-4.2 economy line (SD Mechatronics, Germany). The simulator has two identical sample chambers and two stepper motors which allow computer-controlled vertical and horizontal movements. The masticatory cycle in this research consisted of three phases: contact with a vertical load of 5 kg, horizontal sliding of 0.4 mm, vertical sliding of 0.2 mm and separating the teeth and machine antagonist. In our investigation, two types of tests were made: before and after thermal artificial ageing including all samples from Table 1. The machine was set at 100.000 cycles, which simulated chewing over a period of one year. The loss of substance was measured with an electronic digital calliper – Globathronics GmbH (Ger- many). Accelerated thermal ageing was carried out by immersing the samples in a water bath at a temperature of 70 °C.20,21 D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 873 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Table 1: Composition, mixing ratio and forms of samples for different testing Samples/ abbreviation Composition Mixing ratio Form forcharacterization Form for wear test PMMA1 Pure PMMA 15 mL MMA + 23.4 gPMMA Lamina Tooth model PMMA2 PMMA/2vol%ZnO NPs 15 mL + 22.93 g PMMA+ 0,47 g ZnO NPs Lamina Tooth model PMMA3 PMMA/3vol%ZnO NPs 15 mL + 22.7 g PMMA+ 0,7 g ZnO NPs Lamina Tooth model Figure 2: Schematic presentation of synthesis composite PMMA/ZnO NPs and samples for testing 2.6 Microstructure analysis Microstructural characterizations of all the samples were carried out with scanning electron microscopy (SEM-Sirion 400 NC and Quanta 200 3D). The specimens, without previous preparation, were broken in an atmosphere of N2 to minimise the influence of metallographic preparation regarding changes in the structure of the composite. After breaking, the samples were sprayed with Au (Jeol JSM 8310 appliance), which enables the observation of the non-conductive surfaces of the samples with an electron beam. The samples were positioned into the chamber of the microscope and observations were performed with an accelerating voltage of 15 kV. 2.7 Statistical analysis The statistical analysis included all the parameters involved in the laboratory tests and it was performed in the statistical package SPSS® 17.0. Multivariate linear regression analysis was used to determine the predictors of differences between the analysed groups of samples. The threshold value for accepting the working hypothesis was set at p < 0.05. 3 RESULTS AND DISCUSSION 3.1 Macroscopic view A macroscopic view of both samples’ shapes confirmed the homogenous structure of PMMA and PMMA/ZnO NPs. There was only an acceptable diffe- rence in colour. Adding ZnO NPs to the mixture with PMMA gave a brighter colour as the percentage of nanoparticles grew. This could be solved in the future by the addition of proper pigment, which results in a similar colour as in the teeth or oral mucosa, which are replicated in resin teeth and denture bases. 3.2 Size distribution and zeta-potential by DLS The statement of the homogeneous distribution of ZnO NPs in samples was also confirmed by DLS measurement of the results, which are shown in Table 2. In ZnO NPs with MMA suspension, the average particle size was 2.21±0.10 μm, while 42.1 % of the particles were in the size range of 21.04 nm, whereas in ZnO NPs with water suspension, the average particle size was 4.01±0.4 μm, while 38.1 % of the particles were in the size range of 68.06 nm. Therefore, the average size of the particles was greater in ZnO NPs with water suspension with more presence of larger-sized clusters of particles. Distribution of ZnO NPs in MMA suspension was homogenous, which showed additionally the insolubility of ZnO NPs in water. ZnO NPs with MMA had a sharp and narrow intensity vs. size curve, as com- pared to ZnO NPs with water; thereby, it can be stated that there is a higher degree of agglomeration in the ZnO NPs with water suspension (Figure 3). The present study showed how it is possible to predict, simulate and characterize the composite material, which is aimed for use in dentistry. The DLS technique could confirm indirectly that ZnO NPs in an MMA suspension give a smaller agglomeration, which was also proven through the calculated large hydrodynamic diameters. ZnO NPs which have a zeta-potential between minus 30 mV and plus 30 mV show a tendency for coagulation. Therefore, ZnO NPs suspensions with water were unstable solutions and showed the tendency of agglomerate, while adding MMA in the ZnO NPs did not show much difference in the stability of the suspension as compared with the water. The samples in our work were made by mixing first ZnO NPs nanopowder and PMMA powder matrix in vacuum, at room temperature for 2 h then adding MMA, which reduced the aggregation among ZnO NPs. This may allow us to prepare homogenous nanocomposites with organic matrices without additional surface modification. The chosen average ZnO NPs size showed D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH 874 Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: Number vs. size distribution curve of ZnO NPs with MMA and water Table 2: Composition and the type of mixing used in the preparation of the solvent with the DLS measurements Exp. No. Composition Methodology DLS Measurement Run of measurement Average size range (ìm) Zeta potential (mV) 1 0.05 g ZnONPs+ 15 mL H20 Magnetic stirring (45 min) I 3.50 4.01 ± 0.40 –3.45 –3.38 ± 0.81II 4.08 –2.35 III 4.48 –4.35 2 0.05 g ZnONPs+ 15 mL MMA Magnetic stirring (45 min) I 2.28 2.21 ± 0.10 –3.40 –3.48 ± 0.07II 2.28 –3.57 III 2.06 –3.49 that is of vital importance for the final properties of the new developed composite PMMA/ZnO NPs since it affected the UV-vis absorption and thermal stability. During the synthesis it was found that smaller ZnO NPs showed a higher degree of agglomeration, which is in accordance with the literature.22–24 3.3 Determination of properties The results obtained by the DMA (Figure 4) showed the following results. The difference in the elastic modulus at low temperatures which may be due to a slight decrease in the cross linking of PMMA by adding ZnO NPs. At elevated temperatures, the incidence of chain transfer increased (at termination). With chain transfer, the propagating polymer radical reacts with another molecule by proton abstraction and this leads to branching and cross-linking. Partly, the E modulus increases due to the addition of ZnO NPs, but this increase is minimal. Because of the absence of good interactions between the ZnO NPs and the PMMA matrix, the loss factor increases with the amount of addition of ZnO NPs in the PMMA matrix. ZnO NPs added to the PMMA increased the E modulus and reduced the level of cross-linking. At the same time, the poor interaction between the filler and the matrix results in a less elastic response, which means that the composite of PMMA/ZnO NPs dampen the vibrations better than pure PMMA. With the increasing of the loss factor (tan delta) the damping of the material also increases. In our case, the height of the loss factor peak of the sample PMMA3 was the highest; therefore, the damping of the vibrations of the sample PMMA3 is the highest (Figure 5). Characterization using the DMA method showed that changes in the vibration-damping behaviour of the new composite could mean that at the gingival a more gentle feeling appears, which leads to this material being more friendly to oral mucosa due to vibration damping. This should also be proven through practice. Composite PMMA/ZnO NPs would require longer curing times in comparison with the pure PMMA. Based on this finding it is expected to obtain a better material with the surface modification of ZnO NPs where good interactions between filler and matrix can be formed. From the Flash DSC measurements (Figure 6), the lowest glass-transition temperature was measured in the PMMA3 sample (129.6 °C), followed by PMMA2 (131.0 °C) and the highest glass-transition temperature was in the pure PMMA1 (131.1 °C). This fact is also confirmed by the position of the peaks of the loss factor (tan ), where the lowest glass-transition temperature was measured in the PMMA3 (146.1 °C), followed by PMMA2 (147.2°C) and the highest glass-transition temperature in the pure PMMA1 (147.8 °C). In this context it was found that the degree of cross-linking was reduced by the addition of ZnO NPs. With increasing glass-transition temperature the loss factor decreased, which indicates that the material PMMA3 (tan = 1.2003) has the best vibration-damping behaviour, followed by PMMA2 (tan = 1.1856). The pure PMMA1 (tan = 1.1839) showed the most elastic response. From the height of tan it can be concluded that the 3 % addition of ZnO NPs in PMMA is needed to obtain a significant change in the vibration-damping behaviour of the composite. Despite the small differen- ces measured, as well on DMA (E modulus, tan ) as on Flash DSC (glass-transition temperatures), all the results had the same tendency. D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 875 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: Results of flash DSC analysis for all samples, heating 1.000 °C/s (diagram heat flow (mW) – T (°C)) Figure 4: DMA results- diagram E – modulus (N/mm2) – T (°C) for all samples Figure 5: DMA results-loss factor vs. temperature for all samples 2.6 Microstructure analysis Microstructural characterizations of all the samples were carried out with scanning electron microscopy (SEM-Sirion 400 NC and Quanta 200 3D). The specimens, without previous preparation, were broken in an atmosphere of N2 to minimise the influence of metallographic preparation regarding changes in the structure of the composite. After breaking, the samples were sprayed with Au (Jeol JSM 8310 appliance), which enables the observation of the non-conductive surfaces of the samples with an electron beam. The samples were positioned into the chamber of the microscope and observations were performed with an accelerating voltage of 15 kV. 2.7 Statistical analysis The statistical analysis included all the parameters involved in the laboratory tests and it was performed in the statistical package SPSS® 17.0. Multivariate linear regression analysis was used to determine the predictors of differences between the analysed groups of samples. The threshold value for accepting the working hypothesis was set at p < 0.05. 3 RESULTS AND DISCUSSION 3.1 Macroscopic view A macroscopic view of both samples’ shapes confirmed the homogenous structure of PMMA and PMMA/ZnO NPs. There was only an acceptable diffe- rence in colour. Adding ZnO NPs to the mixture with PMMA gave a brighter colour as the percentage of nanoparticles grew. This could be solved in the future by the addition of proper pigment, which results in a similar colour as in the teeth or oral mucosa, which are replicated in resin teeth and denture bases. 3.2 Size distribution and zeta-potential by DLS The statement of the homogeneous distribution of ZnO NPs in samples was also confirmed by DLS measurement of the results, which are shown in Table 2. In ZnO NPs with MMA suspension, the average particle size was 2.21±0.10 μm, while 42.1 % of the particles were in the size range of 21.04 nm, whereas in ZnO NPs with water suspension, the average particle size was 4.01±0.4 μm, while 38.1 % of the particles were in the size range of 68.06 nm. Therefore, the average size of the particles was greater in ZnO NPs with water suspension with more presence of larger-sized clusters of particles. Distribution of ZnO NPs in MMA suspension was homogenous, which showed additionally the insolubility of ZnO NPs in water. ZnO NPs with MMA had a sharp and narrow intensity vs. size curve, as com- pared to ZnO NPs with water; thereby, it can be stated that there is a higher degree of agglomeration in the ZnO NPs with water suspension (Figure 3). The present study showed how it is possible to predict, simulate and characterize the composite material, which is aimed for use in dentistry. The DLS technique could confirm indirectly that ZnO NPs in an MMA suspension give a smaller agglomeration, which was also proven through the calculated large hydrodynamic diameters. ZnO NPs which have a zeta-potential between minus 30 mV and plus 30 mV show a tendency for coagulation. Therefore, ZnO NPs suspensions with water were unstable solutions and showed the tendency of agglomerate, while adding MMA in the ZnO NPs did not show much difference in the stability of the suspension as compared with the water. The samples in our work were made by mixing first ZnO NPs nanopowder and PMMA powder matrix in vacuum, at room temperature for 2 h then adding MMA, which reduced the aggregation among ZnO NPs. This may allow us to prepare homogenous nanocomposites with organic matrices without additional surface modification. The chosen average ZnO NPs size showed D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH 874 Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 3: Number vs. size distribution curve of ZnO NPs with MMA and water Table 2: Composition and the type of mixing used in the preparation of the solvent with the DLS measurements Exp. No. Composition Methodology DLS Measurement Run of measurement Average size range (ìm) Zeta potential (mV) 1 0.05 g ZnONPs+ 15 mL H20 Magnetic stirring (45 min) I 3.50 4.01 ± 0.40 –3.45 –3.38 ± 0.81II 4.08 –2.35 III 4.48 –4.35 2 0.05 g ZnONPs+ 15 mL MMA Magnetic stirring (45 min) I 2.28 2.21 ± 0.10 –3.40 –3.48 ± 0.07II 2.28 –3.57 III 2.06 –3.49 that is of vital importance for the final properties of the new developed composite PMMA/ZnO NPs since it affected the UV-vis absorption and thermal stability. During the synthesis it was found that smaller ZnO NPs showed a higher degree of agglomeration, which is in accordance with the literature.22–24 3.3 Determination of properties The results obtained by the DMA (Figure 4) showed the following results. The difference in the elastic modulus at low temperatures which may be due to a slight decrease in the cross linking of PMMA by adding ZnO NPs. At elevated temperatures, the incidence of chain transfer increased (at termination). With chain transfer, the propagating polymer radical reacts with another molecule by proton abstraction and this leads to branching and cross-linking. Partly, the E modulus increases due to the addition of ZnO NPs, but this increase is minimal. Because of the absence of good interactions between the ZnO NPs and the PMMA matrix, the loss factor increases with the amount of addition of ZnO NPs in the PMMA matrix. ZnO NPs added to the PMMA increased the E modulus and reduced the level of cross-linking. At the same time, the poor interaction between the filler and the matrix results in a less elastic response, which means that the composite of PMMA/ZnO NPs dampen the vibrations better than pure PMMA. With the increasing of the loss factor (tan delta) the damping of the material also increases. In our case, the height of the loss factor peak of the sample PMMA3 was the highest; therefore, the damping of the vibrations of the sample PMMA3 is the highest (Figure 5). Characterization using the DMA method showed that changes in the vibration-damping behaviour of the new composite could mean that at the gingival a more gentle feeling appears, which leads to this material being more friendly to oral mucosa due to vibration damping. This should also be proven through practice. Composite PMMA/ZnO NPs would require longer curing times in comparison with the pure PMMA. Based on this finding it is expected to obtain a better material with the surface modification of ZnO NPs where good interactions between filler and matrix can be formed. From the Flash DSC measurements (Figure 6), the lowest glass-transition temperature was measured in the PMMA3 sample (129.6 °C), followed by PMMA2 (131.0 °C) and the highest glass-transition temperature was in the pure PMMA1 (131.1 °C). This fact is also confirmed by the position of the peaks of the loss factor (tan ), where the lowest glass-transition temperature was measured in the PMMA3 (146.1 °C), followed by PMMA2 (147.2°C) and the highest glass-transition temperature in the pure PMMA1 (147.8 °C). In this context it was found that the degree of cross-linking was reduced by the addition of ZnO NPs. With increasing glass-transition temperature the loss factor decreased, which indicates that the material PMMA3 (tan = 1.2003) has the best vibration-damping behaviour, followed by PMMA2 (tan = 1.1856). The pure PMMA1 (tan = 1.1839) showed the most elastic response. From the height of tan it can be concluded that the 3 % addition of ZnO NPs in PMMA is needed to obtain a significant change in the vibration-damping behaviour of the composite. Despite the small differen- ces measured, as well on DMA (E modulus, tan ) as on Flash DSC (glass-transition temperatures), all the results had the same tendency. D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 875 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 6: Results of flash DSC analysis for all samples, heating 1.000 °C/s (diagram heat flow (mW) – T (°C)) Figure 4: DMA results- diagram E – modulus (N/mm2) – T (°C) for all samples Figure 5: DMA results-loss factor vs. temperature for all samples The difference in the E modulus at low temperatures may have been due to the slight decrease in the cross- linking of the PMMA by adding ZnO NPs. The degree of cross linking by adding ZnO NPs decreased, as shown by the measurements on the Flash DSC and the level of the peak of tan . This results, at temperatures up to about 45 °C, in a minimum E modulus in the PMMA3 and a maximum E modulus in the PMMA2 (the differen- ce between the glassy transition of PMMA2 and pure PMMA1 was minimal). Partly, the E modulus increased due to the addition of ZnO NPs, but this increase was minimal. Because of the absence of good interactions between the ZnO NPs and the PMMA matrix, the loss factor increased with the amount of addition of ZnO NPs in the PMMA matrix. 3.4 Wear resistance The mean vertical dimension of each tooth sample was measured (from the top of the field where the wear test was performed to the bottom) before the artificial chewing action and after. All the samples were subjected to artificial thermal ageing (the samples were in warm water at 70 °C for one month) and, after that, to mecha- nical ageing on the chewing simulator. The results of mean height loss of the control (PMMA1) and tested group (PMMA2, PMMA3) are shown in Figure 7. Recently, a study showed that the median vertical wear of polymer denture teeth, made by different materials, has been reported to be above 0.2 mm after 2 years of observation and, in over 50 % of these, variability of wear. This could be attributed to specific patient factors such as biting force, nutrition habits and other unknown factors. Gender differences were also found in the spatial and temporal parameters of masticatory movement path and rhythm.25–27 On the other hand, we used the same material and improved these mechanical properties by adding a different percentage of ZnO NPs; we also checked the height loss of the material only after one year of mechanical ageing on a chewing simulator. The tested PMMA2/3 samples showed a higher wear resistance than the tooth sample PMMA1 by 4 times. Artificial thermal ageing had an effect on the pure PMMA1 and PMMA2, but there was no effect in the group with samples made of PMMA3. There was no statistically significant difference in the loss of height between samples made of PMMA2 with PMMA3. Occlusal wear values of the samples made of pure PMMA1 and PMMA2 after thermal artificial ageing were 2-times as big compared to the samples made from PMMA3. Wear resistance of restorative materials under clinical conditions is a rather complicated phenomenon compared to other mechanical and physical properties of materials.25 Furthermore, the cause of thermal stability and its effect on the wear resistance of both PMMA2 and PMMA3 samples could be linked significantly. Long- term polymerization at 70 °C, after polymerization at the usual 100 °C for 1 h and 45 min can be recommended, because of its positive effect on the wear resistance of PMMA1 and PMMA2 after artificial ageing (70 °C, one month). PMMA3 had lower height loss before, as well as after, thermal artificial ageing compared to the other samples. 3.5 Microstructure The microstructure of the fracture surface of the newly developed PMMA2 is shown in micrographs (Figure 8), where the visible fracture is brittle. Detailed observation by higher magnification revealed that ZnO NPs (coloured in white) are distributed homogeneously through the PMMA matrix with some small evidence of ZnO NPs’ agglomeration (about 1 μm) and signs of de-polymerized dark fields around them. This led to the deteriorated final properties of the composite. In the future, it is necessary to avoid the de-polymerization pro- cess by the adequate preparation of ZnO NPs’ surface D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH 876 Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 8: Micrographs of characteristic surface fracture in PMMA2 Figure 7: Loss of height of resin teeth after wear test on chewing simulator before and after thermal artificial aging modification.23 On the other hand, from the literature it is known that homogeneously dispersed ZnO NPs in the PMMA matrix can be explained by the theory of the different kinds of integration by grafting copolymer chains.28,29 4 CONCLUSIONS Based on the methodology applied in this study, and by considering the obtained results, the following con- clusions can be drawn: • The newly developed composite of PMMA/ZnO NPs has a better vibrations damping effect than pure PMMA1. This could lead to the gingival’s more gentle feeling of appearance. • Composite resin denture PMMA teeth reinforced by the ZnO NPs showed better wear resistance by about 4 times compared to the pure PMMA1. • The microstructure of PMMA/ZnO NPs consists of a PMMA matrix and homogenously distributed ZnO NPs. • Combinations of different characterization techniques in the designing of polymer composites reinforced by nanoparticles with in vitro chewing simulation enable the determination of functional composite’s behavior in dentistry. Conflicts of interests There are no conflicts of interests to declare. Acknowledgment This study was supported by the research project "Development of PMMA composite enriched/enhanced with nanoparticles (Ag, ZnO) and biocompatibility testing", number of project 451-03-2802-IP Type 1/144 and by the Infrastructure Programmes I0-0046 and I0-0029 financed by the Slovenian Agency ARRS. Thanks to the Ministry of Education, Science and Sport, Republic of Slovenia (Programme MARTINA, OP20. 00369), which enabled the research with co-financing. Special acknowledgements go to Mohammed Shariq for work on DLS. Note: The responsible translator for the English language is mag. Shelagh Hedges, Faculty of Mechanical Engineer- ing, University of Maribor, Slovenia. Abbreviations: PMMA – Poly-(Methyl-Methacrylate)(powder) MMA – Methyl-Methacrylate(liquid) ZnO NPs – Zinc-Oxide Nanoparticles DMA – Dynamic mechanical analysis PMMA2 – PMMA + 2 vol. %ZnO NPs PMMA3 – PMMA + 3 vol. %ZnO NPs Flash DSC – Flash differential scanning calorimeter DLS – Dynamic light scattering TEM – Transmission Electron Microscope 5 REFERENCES 1 Healthy people, 2010, vol. 2, section 21, Oral Health, www.healthy- people.gov/Publications, 21–18 to 21–19, 25.11.2008 2 F. Müller, M. Naharro, G. E. Carlsson, What are prevalence and incidence of tooth loss in the adult and elderly population in Europe?, Clin. Oral Implants Res., 18 (2007) 3, 2–14, doi:10.1111/ j.1600-0501.2007.01459.x 3 D. A. Felton, Edentulism and comorbid factors, J. Prosthodont., 18 (2009) 2, 86–88, doi:10.1111/j.1532-849X.2009.00437.x 4 J. Zeng, Y. Sato, C. Ohkubo, T. Hosoi, In vitro wear resistance of three types of composite resin denture teeth, J. Prosthet. Dent., 94 (2005) 5, 453–457, doi:10.1016/j.prosdent.2005.08.010 5 M. Risa, Y. Yoshikazu, M. Masakazu , O. Chikahiro, The influence of vertical dimension of occlusion changes on the electroencephalo- grams of complete denture wearers, J. Prosthodont. Res., 58 (2014) 2, 121–126, doi:10.1016/j.jpor.2014.01.003 6 A. L. Machado, E.T. Giampaolo, C. E. Vergani, A. C. Pavarina, D. Salles, J. H. Jorge, Weight loss and changes in surface roughness of denture base and reline materials after simulated tooth brushing in vitro, Gerodontology, 29 (2012) 2, 121–127, doi:10.1111/j.1741- 2358.2010.00422.x 7 M. @igon, G. Ambro`i~, Supramolecular polymers, Mater. Tehnol., 37 (2003) 5, 231–236 8 B. M. Geilich, T. J. Webster, Reduced adhesion of Staphylococcus aureus to ZnO/PVC nanocomposites, Int. J. Nanomed., 8 (2013), 1177–1184, doi:10.2147/IJN.S42010 9 I. Bilecka, M. Niederberger, New developments in the nonaqueous and/or non-hydrolytic sol–gel synthesis of inorganic nanoparticles, Electrochim. Acta, 55 (2010) 26, 7717–7725, doi:10.1016/j.electacta. 2009.12.066 10 K. Memarzadeh, M. Vargas, J. Huang, J. Fan, R. P. Allaker, Nano metallic-oxides as antimicrobials for implant coatings, Key Eng. Mater., 493–494 (2012) 489–494, doi:10.4028/www.scientific.net/ KEM.493–494.489 11 M. Bitenc, Z. Crnjak Orel, Synthesis and characterization of crystalline hexagonal bipods of zinc oxide, Mater. Res. Bull., 44 (2009) 2, 381–387, doi:10.1016/j.materresbull.2008.05.005 12 Y. Li, J. Z. Zhang, Hydrogen generation from photoelectron chemical water splitting based on nanomaterials, Laser Photonics Rev., 4 (2010) 4, 517–528, doi:10.1002/lpor.200910025 13 C. Buzea, I. Pacheco, K. Robbie , Nanomaterials and nanoparticles: Sources and toxicity, Biointerphases, 2 (2007) 4, 17–71, doi:10.1116/ 1.2815690 14 G. R. Patzke, Y. Zhou, R. Kontic, F. Conrad, Oxide nanomaterials: synthetic developments, mechanistic studies, and technological innovations, Angew. Chem. Int. Edit., 50 (2011) 4, 826–859, doi:10.1002/anie.201000235 15 J. Sawai, Quantitative evaluation of antibacterial activities of metallic oxide powders (ZnO, MgO and CaO) by conductimetric assay, J. Microbiol. Meth., 54 (2003) 2, 177–182 16 M. Cierech, J. Wojnarowicz, D. Szmigiel, B. B¹czkowski, A. M. Grudniak, K. I. Wolska, Preparation and characterization of ZnO-PMMA resin nanocomposites for denture bases, Acta Bioeng. Biomech., 18 (2016) 2, 31–41 17 P. Patil, G. Gaikwad, D. R. Patil, J. Naik, Synthesis of 1-D ZnOnano- rods and polypyrrole/1-D ZnO nanocomposites for photocatalysis and gas sensor applications, Bull. Mater. Sci., 39 (2015) 3, 655–665, doi:10.1007/s12034-016-1208-9 18 M. Demir, M. Memesa, P. Castignolles, G. Wegner, PMMA/Zinc Oxide Nanocomposites Prepared by In-Situ Bulk Polymerization, Macromol. Rapid Comm., 27 (2006) 10, 763–770, doi:10.1002/marc. 200500870 D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 877 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS The difference in the E modulus at low temperatures may have been due to the slight decrease in the cross- linking of the PMMA by adding ZnO NPs. The degree of cross linking by adding ZnO NPs decreased, as shown by the measurements on the Flash DSC and the level of the peak of tan . This results, at temperatures up to about 45 °C, in a minimum E modulus in the PMMA3 and a maximum E modulus in the PMMA2 (the differen- ce between the glassy transition of PMMA2 and pure PMMA1 was minimal). Partly, the E modulus increased due to the addition of ZnO NPs, but this increase was minimal. Because of the absence of good interactions between the ZnO NPs and the PMMA matrix, the loss factor increased with the amount of addition of ZnO NPs in the PMMA matrix. 3.4 Wear resistance The mean vertical dimension of each tooth sample was measured (from the top of the field where the wear test was performed to the bottom) before the artificial chewing action and after. All the samples were subjected to artificial thermal ageing (the samples were in warm water at 70 °C for one month) and, after that, to mecha- nical ageing on the chewing simulator. The results of mean height loss of the control (PMMA1) and tested group (PMMA2, PMMA3) are shown in Figure 7. Recently, a study showed that the median vertical wear of polymer denture teeth, made by different materials, has been reported to be above 0.2 mm after 2 years of observation and, in over 50 % of these, variability of wear. This could be attributed to specific patient factors such as biting force, nutrition habits and other unknown factors. Gender differences were also found in the spatial and temporal parameters of masticatory movement path and rhythm.25–27 On the other hand, we used the same material and improved these mechanical properties by adding a different percentage of ZnO NPs; we also checked the height loss of the material only after one year of mechanical ageing on a chewing simulator. The tested PMMA2/3 samples showed a higher wear resistance than the tooth sample PMMA1 by 4 times. Artificial thermal ageing had an effect on the pure PMMA1 and PMMA2, but there was no effect in the group with samples made of PMMA3. There was no statistically significant difference in the loss of height between samples made of PMMA2 with PMMA3. Occlusal wear values of the samples made of pure PMMA1 and PMMA2 after thermal artificial ageing were 2-times as big compared to the samples made from PMMA3. Wear resistance of restorative materials under clinical conditions is a rather complicated phenomenon compared to other mechanical and physical properties of materials.25 Furthermore, the cause of thermal stability and its effect on the wear resistance of both PMMA2 and PMMA3 samples could be linked significantly. Long- term polymerization at 70 °C, after polymerization at the usual 100 °C for 1 h and 45 min can be recommended, because of its positive effect on the wear resistance of PMMA1 and PMMA2 after artificial ageing (70 °C, one month). PMMA3 had lower height loss before, as well as after, thermal artificial ageing compared to the other samples. 3.5 Microstructure The microstructure of the fracture surface of the newly developed PMMA2 is shown in micrographs (Figure 8), where the visible fracture is brittle. Detailed observation by higher magnification revealed that ZnO NPs (coloured in white) are distributed homogeneously through the PMMA matrix with some small evidence of ZnO NPs’ agglomeration (about 1 μm) and signs of de-polymerized dark fields around them. This led to the deteriorated final properties of the composite. In the future, it is necessary to avoid the de-polymerization pro- cess by the adequate preparation of ZnO NPs’ surface D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH 876 Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Figure 8: Micrographs of characteristic surface fracture in PMMA2 Figure 7: Loss of height of resin teeth after wear test on chewing simulator before and after thermal artificial aging modification.23 On the other hand, from the literature it is known that homogeneously dispersed ZnO NPs in the PMMA matrix can be explained by the theory of the different kinds of integration by grafting copolymer chains.28,29 4 CONCLUSIONS Based on the methodology applied in this study, and by considering the obtained results, the following con- clusions can be drawn: • The newly developed composite of PMMA/ZnO NPs has a better vibrations damping effect than pure PMMA1. This could lead to the gingival’s more gentle feeling of appearance. • Composite resin denture PMMA teeth reinforced by the ZnO NPs showed better wear resistance by about 4 times compared to the pure PMMA1. • The microstructure of PMMA/ZnO NPs consists of a PMMA matrix and homogenously distributed ZnO NPs. • Combinations of different characterization techniques in the designing of polymer composites reinforced by nanoparticles with in vitro chewing simulation enable the determination of functional composite’s behavior in dentistry. Conflicts of interests There are no conflicts of interests to declare. Acknowledgment This study was supported by the research project "Development of PMMA composite enriched/enhanced with nanoparticles (Ag, ZnO) and biocompatibility testing", number of project 451-03-2802-IP Type 1/144 and by the Infrastructure Programmes I0-0046 and I0-0029 financed by the Slovenian Agency ARRS. Thanks to the Ministry of Education, Science and Sport, Republic of Slovenia (Programme MARTINA, OP20. 00369), which enabled the research with co-financing. Special acknowledgements go to Mohammed Shariq for work on DLS. Note: The responsible translator for the English language is mag. Shelagh Hedges, Faculty of Mechanical Engineer- ing, University of Maribor, Slovenia. Abbreviations: PMMA – Poly-(Methyl-Methacrylate)(powder) MMA – Methyl-Methacrylate(liquid) ZnO NPs – Zinc-Oxide Nanoparticles DMA – Dynamic mechanical analysis PMMA2 – PMMA + 2 vol. %ZnO NPs PMMA3 – PMMA + 3 vol. %ZnO NPs Flash DSC – Flash differential scanning calorimeter DLS – Dynamic light scattering TEM – Transmission Electron Microscope 5 REFERENCES 1 Healthy people, 2010, vol. 2, section 21, Oral Health, www.healthy- people.gov/Publications, 21–18 to 21–19, 25.11.2008 2 F. Müller, M. Naharro, G. E. Carlsson, What are prevalence and incidence of tooth loss in the adult and elderly population in Europe?, Clin. Oral Implants Res., 18 (2007) 3, 2–14, doi:10.1111/ j.1600-0501.2007.01459.x 3 D. A. Felton, Edentulism and comorbid factors, J. Prosthodont., 18 (2009) 2, 86–88, doi:10.1111/j.1532-849X.2009.00437.x 4 J. Zeng, Y. Sato, C. Ohkubo, T. Hosoi, In vitro wear resistance of three types of composite resin denture teeth, J. Prosthet. Dent., 94 (2005) 5, 453–457, doi:10.1016/j.prosdent.2005.08.010 5 M. Risa, Y. Yoshikazu, M. Masakazu , O. Chikahiro, The influence of vertical dimension of occlusion changes on the electroencephalo- grams of complete denture wearers, J. Prosthodont. Res., 58 (2014) 2, 121–126, doi:10.1016/j.jpor.2014.01.003 6 A. L. Machado, E.T. Giampaolo, C. E. Vergani, A. C. Pavarina, D. Salles, J. H. Jorge, Weight loss and changes in surface roughness of denture base and reline materials after simulated tooth brushing in vitro, Gerodontology, 29 (2012) 2, 121–127, doi:10.1111/j.1741- 2358.2010.00422.x 7 M. @igon, G. Ambro`i~, Supramolecular polymers, Mater. Tehnol., 37 (2003) 5, 231–236 8 B. M. Geilich, T. J. Webster, Reduced adhesion of Staphylococcus aureus to ZnO/PVC nanocomposites, Int. J. Nanomed., 8 (2013), 1177–1184, doi:10.2147/IJN.S42010 9 I. Bilecka, M. Niederberger, New developments in the nonaqueous and/or non-hydrolytic sol–gel synthesis of inorganic nanoparticles, Electrochim. Acta, 55 (2010) 26, 7717–7725, doi:10.1016/j.electacta. 2009.12.066 10 K. Memarzadeh, M. Vargas, J. Huang, J. Fan, R. P. Allaker, Nano metallic-oxides as antimicrobials for implant coatings, Key Eng. Mater., 493–494 (2012) 489–494, doi:10.4028/www.scientific.net/ KEM.493–494.489 11 M. Bitenc, Z. Crnjak Orel, Synthesis and characterization of crystalline hexagonal bipods of zinc oxide, Mater. Res. Bull., 44 (2009) 2, 381–387, doi:10.1016/j.materresbull.2008.05.005 12 Y. Li, J. Z. Zhang, Hydrogen generation from photoelectron chemical water splitting based on nanomaterials, Laser Photonics Rev., 4 (2010) 4, 517–528, doi:10.1002/lpor.200910025 13 C. Buzea, I. Pacheco, K. Robbie , Nanomaterials and nanoparticles: Sources and toxicity, Biointerphases, 2 (2007) 4, 17–71, doi:10.1116/ 1.2815690 14 G. R. Patzke, Y. Zhou, R. Kontic, F. Conrad, Oxide nanomaterials: synthetic developments, mechanistic studies, and technological innovations, Angew. Chem. Int. Edit., 50 (2011) 4, 826–859, doi:10.1002/anie.201000235 15 J. Sawai, Quantitative evaluation of antibacterial activities of metallic oxide powders (ZnO, MgO and CaO) by conductimetric assay, J. Microbiol. Meth., 54 (2003) 2, 177–182 16 M. Cierech, J. Wojnarowicz, D. Szmigiel, B. B¹czkowski, A. M. Grudniak, K. I. Wolska, Preparation and characterization of ZnO-PMMA resin nanocomposites for denture bases, Acta Bioeng. Biomech., 18 (2016) 2, 31–41 17 P. Patil, G. Gaikwad, D. R. Patil, J. Naik, Synthesis of 1-D ZnOnano- rods and polypyrrole/1-D ZnO nanocomposites for photocatalysis and gas sensor applications, Bull. Mater. Sci., 39 (2015) 3, 655–665, doi:10.1007/s12034-016-1208-9 18 M. 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Jovanovic, Influence of microwave heating on the polymerization kinetics and application properties of the PMMA dental materials, J. Appl. Polym. Sci., 119 (2011) 6, 3598–3606, doi:10.1002/app.33041 22 C. Feldmann, Polyol mediated synthesis of oxide particle suspen- sions and their application, Scripta Mater., 44 (2001) 8–9, 2193–2196, doi:10.1016/S1359-6462(01)00902-2 23 R. Y. Hong, J. Z. Qian, J. X. Cao, Synthesis and characterization of PMMA grafted ZnO nanoparticles, Powder Technol., 163 (2006), 160–168, doi:10.1016/j.powtec.2006.01.015 24 A. An`lovar, A. Crnjak, M. @igon, Poly(methyl methacrylate) com- posites prepared by in situ polymerization using organophillicnano- to-submicrometer zinc oxide particles, Eur. Polym. J., 46 (2010) 6, 1216–1224, doi:10.1016/j.eurpolymj.2010.03.010 25 M. Ghazal, B. Yang, K. Ludwig, M. Kern, Two-body wear of resin and ceramic denture teeth in comparison to human enamel, Dent. Mater., 24 (2008) 4, 502–507, doi:10.1016/j.dental.2007.04.012 26 S. D. Heintze, G. Zellweger, S. Sbicego, V. Rousson, C. Muñoz- Viveros, T. Stober, Wear of two denture teeth materials in vivo- 2-year results, Dent. Mater., 29 (2013) 9, 191–204, doi:10.1016/ j.dental.2013.04.012 27 T. Kyoko, S. Hiroshi, Gender differences in masticatory movement path and rhythm in dentate adults, J. Prosthodont. Res., 58 (2014) 4, 237–242, doi:10.1016/j.jpor.2014.06.001 28 E. Thang, G. Cheng, X. Ma, Preparation of nano-ZnO/PMMA composite particles via grafting of the copolymer onto the surface of zinc oxide nanoparticle, Powder Technol., 161 (2006) 3, 209–214, doi:10.1016/j.powtec.2005.10.007 29 A. Anzlovar, Z. Crnjak Orel, M. Zigon, Sub micrometer and nano-ZnO as filler in PMMA materials, Mater. Tehnol., 45 (2011) 3, 269–274 D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH 878 Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS ERRATUM In Materiali in tehnologije/Materials and Technology 50 (2016) 5, 797-804, doi: 10.17222/mit.2015.307 in the article entitled: Measurement of bio impedance on an isolated rat sciatic nerve elicited with specific current stimulating pulses – Meritev bioimpedance na izoliranem `ivcu ischiadicusu podgane vzbujene s posebnimi tokovnimi stimulacijskimi impulzi written by Janez Rozman1,3, Monika C. @u`ek2, Robert Frange`2, Samo Ribari~3 1Center for Implantable Technology and Sensors, ITIS d. o. o., Lepi pot 11, 1000 Ljubljana, Slovenia 2Institute of Physiology, Pharmacology and Toxicology, Veterinary Faculty, University of Ljubljana, Gerbi~eva 60, 1000 Ljubljana, Slovenia 3 Institute of Pathophysiology, Zalo{ka 4, Medical Faculty, University of Ljubljana, Republic of Slovenia two additional authors are missing. The correct authorship of this article is: Robert Brajkovi~1, Dejan Kri`aj1, Janez Rozman2,3, Monika C. @u`ek4, Robert Frange`4, Samo Ribari~3 1Faculty of Electrical Engineering, University of Ljubljana, Tr`a{ka 25, 1000 Ljubljana, 2Center for Implantable Technology and Sensors, ITIS d. o. o. Ljubljana, Lepi pot 11, 1000 Ljubljana, 3Institute of Pathophysiology, Medical Faculty, University of Ljubljana, Vrazov trg 2, 1000 Ljubljana, 4Institute of Physiology, Pharmacology and Toxicology, Veterinary Faculty, University of Ljubljana, Gerbi~eva 60, 1000 Ljubljana, Republic of Slovenia MIT Editorial MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS 19 M. Roman, Particle Size and Zeta Potential of ZnO, APCBEE Pro- cedia 9 (2014) 13–17, doi:10.1016/j.apcbee.2014.01.003 20 A. Mahomed, D. W. L. Hukins, S. N. Kukureka, Effect of accelerated aging on the viscoelastic properties of Elast-Eon™: A polyurethane with soft poly(dimethylsiloxane) and poly(hexamethylene oxide) segments, Mat. Sci. Eng. C, 30 (2010) 8, 1298–1303, doi:10.1016/ j.msec.2010.07.014 21 P. Spasojevic, B. Adnadjevic, S. Velickovic, J. Jovanovic, Influence of microwave heating on the polymerization kinetics and application properties of the PMMA dental materials, J. Appl. Polym. Sci., 119 (2011) 6, 3598–3606, doi:10.1002/app.33041 22 C. Feldmann, Polyol mediated synthesis of oxide particle suspen- sions and their application, Scripta Mater., 44 (2001) 8–9, 2193–2196, doi:10.1016/S1359-6462(01)00902-2 23 R. Y. Hong, J. Z. Qian, J. X. Cao, Synthesis and characterization of PMMA grafted ZnO nanoparticles, Powder Technol., 163 (2006), 160–168, doi:10.1016/j.powtec.2006.01.015 24 A. An`lovar, A. Crnjak, M. @igon, Poly(methyl methacrylate) com- posites prepared by in situ polymerization using organophillicnano- to-submicrometer zinc oxide particles, Eur. Polym. J., 46 (2010) 6, 1216–1224, doi:10.1016/j.eurpolymj.2010.03.010 25 M. Ghazal, B. Yang, K. Ludwig, M. Kern, Two-body wear of resin and ceramic denture teeth in comparison to human enamel, Dent. Mater., 24 (2008) 4, 502–507, doi:10.1016/j.dental.2007.04.012 26 S. D. Heintze, G. Zellweger, S. Sbicego, V. Rousson, C. Muñoz- Viveros, T. Stober, Wear of two denture teeth materials in vivo- 2-year results, Dent. Mater., 29 (2013) 9, 191–204, doi:10.1016/ j.dental.2013.04.012 27 T. Kyoko, S. Hiroshi, Gender differences in masticatory movement path and rhythm in dentate adults, J. Prosthodont. Res., 58 (2014) 4, 237–242, doi:10.1016/j.jpor.2014.06.001 28 E. Thang, G. Cheng, X. Ma, Preparation of nano-ZnO/PMMA composite particles via grafting of the copolymer onto the surface of zinc oxide nanoparticle, Powder Technol., 161 (2006) 3, 209–214, doi:10.1016/j.powtec.2005.10.007 29 A. Anzlovar, Z. Crnjak Orel, M. Zigon, Sub micrometer and nano-ZnO as filler in PMMA materials, Mater. Tehnol., 45 (2011) 3, 269–274 D. POPOVI] et al.: SYNTHESIS OF PMMA/ZnO NANOPARTICLES COMPOSITE USED FOR RESIN TEETH 878 Materiali in tehnologije / Materials and technology 51 (2017) 5, 871–878 MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS ERRATUM In Materiali in tehnologije/Materials and Technology 50 (2016) 5, 797-804, doi: 10.17222/mit.2015.307 in the article entitled: Measurement of bio impedance on an isolated rat sciatic nerve elicited with specific current stimulating pulses – Meritev bioimpedance na izoliranem `ivcu ischiadicusu podgane vzbujene s posebnimi tokovnimi stimulacijskimi impulzi written by Janez Rozman1,3, Monika C. @u`ek2, Robert Frange`2, Samo Ribari~3 1Center for Implantable Technology and Sensors, ITIS d. o. o., Lepi pot 11, 1000 Ljubljana, Slovenia 2Institute of Physiology, Pharmacology and Toxicology, Veterinary Faculty, University of Ljubljana, Gerbi~eva 60, 1000 Ljubljana, Slovenia 3 Institute of Pathophysiology, Zalo{ka 4, Medical Faculty, University of Ljubljana, Republic of Slovenia two additional authors are missing. The correct authorship of this article is: Robert Brajkovi~1, Dejan Kri`aj1, Janez Rozman2,3, Monika C. @u`ek4, Robert Frange`4, Samo Ribari~3 1Faculty of Electrical Engineering, University of Ljubljana, Tr`a{ka 25, 1000 Ljubljana, 2Center for Implantable Technology and Sensors, ITIS d. o. o. Ljubljana, Lepi pot 11, 1000 Ljubljana, 3Institute of Pathophysiology, Medical Faculty, University of Ljubljana, Vrazov trg 2, 1000 Ljubljana, 4Institute of Physiology, Pharmacology and Toxicology, Veterinary Faculty, University of Ljubljana, Gerbi~eva 60, 1000 Ljubljana, Republic of Slovenia MIT Editorial MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS Instrumentalia, d.o.o. Lesko kova cesta 9 E SI – 1000 Ljubljana Tel: + 386 1 524 0196 Fax: + 386 1 524 0198 GSM: + 386 51 385 007 š Instrumentalia, d.o.o. Lesko kova cesta 9 E SI – 1000 Ljubljana Tel: + 386 1 524 0196 Fax: + 386 1 524 0198 GSM: + 386 51 385 007 š TALUM, d.d., KIDRI^EVO Tovarni{ka ulica 10 2325 Kidri~evo, Slovenia Telephone: +386 2 799 51 00 Telefax: +386 2 799 51 03 INSTRUCTIONS FOR AUTHORS 1 SUBMISSION OF MANUSCRIPT An electronic version (using Microsoft Word or similar) of the manuscript, complete with abstract, keywords, figures and tables must be submitted to the MIT Editorial Office (mit@imt.si) as an e-mail attachment. All contributions must be written in Slovene or English, following the MIT template. The manuscript must be accompanied by the MIT checklist, data and cover letter. Authors must provide the details of two potential reviewers who can make an independent assessment of the quality of the manuscript. The purpose of the reviewing procedure is to have the manuscripts evaluated objectively, and so suggested reviewers with local affiliations and who may be too closely associated with the authors, will be disregarded. Authors must provide details of two potential referees who are not associated with the specific study, nor with the general research activities of the authors. After the review process the manuscript will be returned to the author, with the decision of the Editor and, if appropriate, the reviews of the referees. If the paper is accepted the author will be asked to amend the manuscript in accordance with the referees' comments and return it to the MIT Editorial Office. Authors must state that the results in their manuscript have not been published before, nor are they included in a paper that has been submitted to another journal. Upon acceptance of an article, the copyright is transferred to the publisher. This transfer will ensure the widest possible dissemination of the article. 1.1 Preparation of the manuscript The paper must be as short as possible and should not exceed 4-6 printed pages (approx. 20.000 characters). Papers presented at conferences must be restricted to 2-4 printed pages (approx. 10.000 characters). Title page Papers must have a concise but informative title, which should not exceed one line. The words from the title must be suitable for indexing and searching. The title must be followed by the name(s) of the author(s) and by the name and address of the institution(s) where the work was carried out. An e-mail address for the contact author must be supplied. Abstract Papers must include an abstract, which should provide an overview of the content and conclusions of the paper and highlight the new information they contain. The abstract must be understandable in isolation and written in the past tense, because it refers to work that was already done. The length of the abstract should not exceed 250 words. Keywords The author must supply 3–4 keywords that describe the content of the article and are suitable for indexing and searching. 1.2 Symbols, equations (units of measurement) Units of measurement must comply with the Law of Units of Measurement and Measures (Official Gazette of the Republic of Slovenia 2001/26), i.e., international SI units. Equations must be marked on the right-hand side of the text with numbers in round brackets. Symbols in text should be written as text, not as equation. 1.3 Tables Tables must be clearly referred to in the text using Arabic numerals. Each table must have a title which makes the general meaning understandable without reference to the text. 1.4 Figures Figures are reduced to a single-column width (7.9 cm) except in special cases (max. printed size ˜16 cm). The lettering used on a figure should be chosen so that after reduction the height of numbers and capital letters falls within the range (1.2–2.4) mm. Lines and arrows should also be of sufficient thickness so as to remain clear after the reduction process. Figures must be saved in any supported format, e.g. JPG, TIFF. Figures and figures captions must be inserted in the text. Illustrations can be printed in colour when they are judged by the Editor to be essential to the presentation. Further information concerning colour illustrations and the costs to the author can be obtained from the publisher. Colour figures can be included, but with an additional page charge of 80. Maximum number of figures and tables in paper is 10, for papers presented at conferences must be restricted to 8. 1.5 References The references must be collected at the end of the article, and numbered in the order of their appearance in the text. Each reference must be complete, the use of ibid., idem., et al., etc. is not permitted. TALUM, d.d., KIDRI^EVO Tovarni{ka ulica 10 2325 Kidri~evo, Slovenia Telephone: +386 2 799 51 00 Telefax: +386 2 799 51 03 INSTRUCTIONS FOR AUTHORS 1 SUBMISSION OF MANUSCRIPT An electronic version (using Microsoft Word or similar) of the manuscript, complete with abstract, keywords, figures and tables must be submitted to the MIT Editorial Office (mit@imt.si) as an e-mail attachment. All contributions must be written in Slovene or English, following the MIT template. The manuscript must be accompanied by the MIT checklist, data and cover letter. Authors must provide the details of two potential reviewers who can make an independent assessment of the quality of the manuscript. The purpose of the reviewing procedure is to have the manuscripts evaluated objectively, and so suggested reviewers with local affiliations and who may be too closely associated with the authors, will be disregarded. Authors must provide details of two potential referees who are not associated with the specific study, nor with the general research activities of the authors. After the review process the manuscript will be returned to the author, with the decision of the Editor and, if appropriate, the reviews of the referees. If the paper is accepted the author will be asked to amend the manuscript in accordance with the referees' comments and return it to the MIT Editorial Office. Authors must state that the results in their manuscript have not been published before, nor are they included in a paper that has been submitted to another journal. Upon acceptance of an article, the copyright is transferred to the publisher. This transfer will ensure the widest possible dissemination of the article. 1.1 Preparation of the manuscript The paper must be as short as possible and should not exceed 4-6 printed pages (approx. 20.000 characters). Papers presented at conferences must be restricted to 2-4 printed pages (approx. 10.000 characters). Title page Papers must have a concise but informative title, which should not exceed one line. The words from the title must be suitable for indexing and searching. The title must be followed by the name(s) of the author(s) and by the name and address of the institution(s) where the work was carried out. An e-mail address for the contact author must be supplied. Abstract Papers must include an abstract, which should provide an overview of the content and conclusions of the paper and highlight the new information they contain. The abstract must be understandable in isolation and written in the past tense, because it refers to work that was already done. The length of the abstract should not exceed 250 words. Keywords The author must supply 3–4 keywords that describe the content of the article and are suitable for indexing and searching. 1.2 Symbols, equations (units of measurement) Units of measurement must comply with the Law of Units of Measurement and Measures (Official Gazette of the Republic of Slovenia 2001/26), i.e., international SI units. Equations must be marked on the right-hand side of the text with numbers in round brackets. Symbols in text should be written as text, not as equation. 1.3 Tables Tables must be clearly referred to in the text using Arabic numerals. Each table must have a title which makes the general meaning understandable without reference to the text. 1.4 Figures Figures are reduced to a single-column width (7.9 cm) except in special cases (max. printed size ˜16 cm). The lettering used on a figure should be chosen so that after reduction the height of numbers and capital letters falls within the range (1.2–2.4) mm. Lines and arrows should also be of sufficient thickness so as to remain clear after the reduction process. Figures must be saved in any supported format, e.g. JPG, TIFF. Figures and figures captions must be inserted in the text. Illustrations can be printed in colour when they are judged by the Editor to be essential to the presentation. Further information concerning colour illustrations and the costs to the author can be obtained from the publisher. Colour figures can be included, but with an additional page charge of 80. Maximum number of figures and tables in paper is 10, for papers presented at conferences must be restricted to 8. 1.5 References The references must be collected at the end of the article, and numbered in the order of their appearance in the text. Each reference must be complete, the use of ibid., idem., et al., etc. is not permitted. References to unpublished or not readily accessible reports must be avoided. References must be cited in English. All references should have a DOI number, if it exists for the given reference. Authors should avoid self-citing in the references as far as possible. In the list of references, monographs, articles in journals, journals, contributions to conference pro- ceedings, patent documents, electronic monographs, articles and other electronic documents must be cited in accordance with the following examples: 1. Monographs H. Ibach, H. Luth, Solid State physics, 2nd ed., Springer, Berlin 1991, 245 2. Articles in journals T. Mauder, J. Stetina, Improvement of the casting of special steel with a wide solid liquid interface, Mater. Tehnol., 50 (2015) 1–2, doi:10.17222/mit.2014.122 3. Contributions to conference proceedings, sym- posiums or conferences I. Rak, M. Kocak, V. Gliha, N. Gubeljak: Fracture behaviour of over-matched high strength steel welds containing soft root layers, Proc. of the 2nd Inter. Symp. on Mis-Matching of Interfaces and Welds, Reinsford, 1997, 627–641 4. Contributions in electronic form/online • Articles: M. P. Wnuk: Principles of fracture mechanics for space applications, http://mit.imt.si/ Revija/izvodi/kzt996/wnuk.pdf, 30.01.2000 • Other: http://www.imt.si/, 15.03.2016 5. Standards ISO 15787:2001(E) Technical product documen- tation, Heat treated ferrous part, Presentation and indication ISO Committee, Geneve SPECIAL NOTICE TO AUTHORS In the past few years we have received an increasing number of articles that include computer simulations, with 3D colour images as a result. As the journal charges 80 per page for colour printing, some authors have decided to convert their colour images to grey-scale images. We believe that this lessens both the impact and the clarity of the results. Since the Materials and Technology journal consistently strives for better quality presentation and a higher impact factor, we will be rejecting articles of this type. 2 ARTICLE PROOFS Authors will receive a set of proofs. They are requested to return the proofs with any corrections within two days. In the case of a delay the Editor will postpone publication of the article until the article proof is received. 3 COPYRIGHT In addition to the paper, authors must also enclose a written statement that the paper is original work and has not been published in this form anywhere else and that it is not under consideration for publication elsewhere. On publication, copyright will pass to the publisher. The Journal of Materials and Technology must be stated as the source in all later publications. The Editorial Board of the Journal of Materials and Technology: • decide whether to accept a paper for publication; • obtain professional reviewers for papers and decide on any proposals to shorten or extend them; • obtain correct terminology and edit language. The e-files of papers will be kept in the archives of the Materials and Technology journal. 4 PUBLICATION FEE Authors will be asked to pay a publication fee for each article before its publishing in MIT journal. After the article has been accepted for publishing, this fee needs to be paid. The fee is  300 for regular articles, and  150 for articles, presented at ICMT annual conference in Portoro`. The additional cost for coloured printing for one page is  80. These fees do include value-added tax (VAT). MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY Lepi pot 11, 1000 Ljubljana, Slovenia Phone: +386 1 4701 860, +386 1 4701 857, Fax: +386 01 4701 939, e-mail: mit@imt.si Editor-in-chief: dr. Paul John McGuiness ----------------------------------------------------------------------------------------------------------------- MIT TEMPLATE TITLE (The fewest possible words that adequately describe the contents of the paper in English) NASLOV (v slovenskem jeziku, for Slovenian authors only) Author`s Name Author`s Surname1, Author`s Name Author`s Surname2, Author`s Name Author`s Surname3* 1Author`s Affiliation, Address, City, State 2Author`s Affiliation, Address, City, State 3Author`s Affiliation, Address, City, State *Corresponding author`s e-mail Abstract (in English) The abstract must be viewed as a miniversion of the paper and should not exceed 250 words. It must be written in `past tense` , because it refers to work already done. It must state the principal objectives and scope of the investigation, describe the methodology employed, summarize the results and state the principal conclusions. It must not give information or conclusions that are not inculded in the paper and it must not cite references to the literature. Keywords (in English) Up to 3-4 words (abbreviations of the expressions are not allowed, only the common used ones: SEM, XRD, XPS, etc.) References to unpublished or not readily accessible reports must be avoided. References must be cited in English. All references should have a DOI number, if it exists for the given reference. Authors should avoid self-citing in the references as far as possible. In the list of references, monographs, articles in journals, journals, contributions to conference pro- ceedings, patent documents, electronic monographs, articles and other electronic documents must be cited in accordance with the following examples: 1. Monographs H. Ibach, H. Luth, Solid State physics, 2nd ed., Springer, Berlin 1991, 245 2. Articles in journals T. Mauder, J. Stetina, Improvement of the casting of special steel with a wide solid liquid interface, Mater. Tehnol., 50 (2015) 1–2, doi:10.17222/mit.2014.122 3. Contributions to conference proceedings, sym- posiums or conferences I. Rak, M. Kocak, V. Gliha, N. Gubeljak: Fracture behaviour of over-matched high strength steel welds containing soft root layers, Proc. of the 2nd Inter. Symp. on Mis-Matching of Interfaces and Welds, Reinsford, 1997, 627–641 4. Contributions in electronic form/online • Articles: M. P. Wnuk: Principles of fracture mechanics for space applications, http://mit.imt.si/ Revija/izvodi/kzt996/wnuk.pdf, 30.01.2000 • Other: http://www.imt.si/, 15.03.2016 5. Standards ISO 15787:2001(E) Technical product documen- tation, Heat treated ferrous part, Presentation and indication ISO Committee, Geneve SPECIAL NOTICE TO AUTHORS In the past few years we have received an increasing number of articles that include computer simulations, with 3D colour images as a result. As the journal charges 80 per page for colour printing, some authors have decided to convert their colour images to grey-scale images. We believe that this lessens both the impact and the clarity of the results. Since the Materials and Technology journal consistently strives for better quality presentation and a higher impact factor, we will be rejecting articles of this type. 2 ARTICLE PROOFS Authors will receive a set of proofs. They are requested to return the proofs with any corrections within two days. In the case of a delay the Editor will postpone publication of the article until the article proof is received. 3 COPYRIGHT In addition to the paper, authors must also enclose a written statement that the paper is original work and has not been published in this form anywhere else and that it is not under consideration for publication elsewhere. On publication, copyright will pass to the publisher. The Journal of Materials and Technology must be stated as the source in all later publications. The Editorial Board of the Journal of Materials and Technology: • decide whether to accept a paper for publication; • obtain professional reviewers for papers and decide on any proposals to shorten or extend them; • obtain correct terminology and edit language. The e-files of papers will be kept in the archives of the Materials and Technology journal. 4 PUBLICATION FEE Authors will be asked to pay a publication fee for each article before its publishing in MIT journal. After the article has been accepted for publishing, this fee needs to be paid. The fee is  300 for regular articles, and  150 for articles, presented at ICMT annual conference in Portoro`. The additional cost for coloured printing for one page is  80. These fees do include value-added tax (VAT). MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY Lepi pot 11, 1000 Ljubljana, Slovenia Phone: +386 1 4701 860, +386 1 4701 857, Fax: +386 01 4701 939, e-mail: mit@imt.si Editor-in-chief: dr. Paul John McGuiness ----------------------------------------------------------------------------------------------------------------- MIT TEMPLATE TITLE (The fewest possible words that adequately describe the contents of the paper in English) NASLOV (v slovenskem jeziku, for Slovenian authors only) Author`s Name Author`s Surname1, Author`s Name Author`s Surname2, Author`s Name Author`s Surname3* 1Author`s Affiliation, Address, City, State 2Author`s Affiliation, Address, City, State 3Author`s Affiliation, Address, City, State *Corresponding author`s e-mail Abstract (in English) The abstract must be viewed as a miniversion of the paper and should not exceed 250 words. It must be written in `past tense` , because it refers to work already done. It must state the principal objectives and scope of the investigation, describe the methodology employed, summarize the results and state the principal conclusions. It must not give information or conclusions that are not inculded in the paper and it must not cite references to the literature. Keywords (in English) Up to 3-4 words (abbreviations of the expressions are not allowed, only the common used ones: SEM, XRD, XPS, etc.) MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY Lepi pot 11, 1000 Ljubljana, Slovenia Phone: +386 1 4701 860, +386 1 4701 857, Fax: +386 01 4701 939, e-mail: mit@imt.si Editor-in-chief: dr. Paul John McGuiness ----------------------------------------------------------------------------------------------------------------- Povzetek (v slovenskem jeziku, for Slovenian authors only) Do 250 besed Ključne besede (v slovenskem jeziku, for Slovenian authors only) Od 3 do 4 besede (okrajšave izrazov niso dovoljene, razen splošno sprejetih: SEM, XRD, XPS, itd.) 1 INTRODUCTION The purpose of the introduction is to supply sufficient background information to allow the reader to understand and evaluate the results of the present study. It should state briefly and clearly the purpose in writing paper, present as clearly as possible the nature and scope of the problem investigated, review recent literature to orient the reader, state the method of the investigation, and if necessary, the reasons for the choice of a particular method. 2 EXPERIMENTAL PART The experimental part must give full details of the experimental aparatus and the methods used in obtaining results. It must include accurate details relating to the equipment and materials used, incuding quantities, temperatures, times etc. It must provide sufficient information for a collegaue in the same field to reproduce the experiment. The data must be presented clearly and concisely, using graphs and tables. Figures and Tables must be written bold and not abbreviated throughout text. Figures and Tables must also follow the numerical order of appearance in the text. Once mentioned, tehy can be repeated. Equations must be written with Equation Editor and must be marked on the right-hand side of the text with numbers in round brackets. Throughout the text they must not be abbreviated.Physical quantities in the text have to be written as a plain text. Equation 1 (1) Equation 2 (2) 3 RESULTS Results section should provide an overall picture of the experiments without repeating any of the details in the experimental section. The data should be presented clearly and concisely, using graphs and tables. Figures and Tables should be written bold and not abbreviated throughout text. MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY Lepi pot 11, 1000 Ljubljana, Slovenia Phone: +386 1 4701 860, +386 1 4701 857, Fax: +386 01 4701 939, e-mail: mit@imt.si Editor-in-chief: dr. Paul John McGuiness ----------------------------------------------------------------------------------------------------------------- Figures and Tables must also follow the numerical order of appearance in the text. Once mentioned they can be repeated. 4 DISCUSSION The purpose of the discussion is to present the principles, relationships, and generalisations shown by the results. The discussion must make clear the significance of the results and compare the findings with previously published work. It must discuss the results not simply repeat them. 5 CONCLUSIONS The conclusion section should be short. It must present one or more conclusions which have to be drawn from the results and the subsequent discussion. Conclusions must be concise and easily understood by the reader. Acknowledgment 6 REFERENCES All references need a DOI number if it exists for the given reference. In the list of references monographs, articles in journals, journals, contributions to conference proceedings, patent documents, electronic monographs, articles and other electronic documents must be cited in accordance with the following examples: 1. Monographs: H. Ibach, H. Luth, Solid State physics, 2nd ed., Springer, Berlin 1991, 245 2. Articles in journals: T. Mauder, J. Stetina, Improvement of the casting of special steel with wide solid liquid interface, Mater. Tehnol., 50 (2015) 1, doi:10.17222/mit.2014.122 3. Contributions to conference proceedings, symposiums: I. Rak, M. Kocak, V. Gliha, N. Gubeljak, Fracture behaviour of over-matched high strength steel welds containing soft root layers, Proc. of the 2nd Inter. Symp. on Mis- Matching of Interfaces and Welds, Reinsford 1997, 627–641 4. Contributions in electronic form/online: Articles: M. P. Wnuk: Principles of fracture mechanics for space applications, http://mit.imt.si/Revija/izvodi/kzt996/wnuk.pdf, 30.01.2000 Other: http://www.imt.si, 15.03.2016 5. Standards: ISO 15787:2001(E) Technical product documentation, Heat treated ferrous part, Presentation and indication ISO Committee, Geneve MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY Lepi pot 11, 1000 Ljubljana, Slovenia Phone: +386 1 4701 860, +386 1 4701 857, Fax: +386 01 4701 939, e-mail: mit@imt.si Editor-in-chief: dr. Paul John McGuiness ----------------------------------------------------------------------------------------------------------------- Povzetek (v slovenskem jeziku, for Slovenian authors only) Do 250 besed Ključne besede (v slovenskem jeziku, for Slovenian authors only) Od 3 do 4 besede (okrajšave izrazov niso dovoljene, razen splošno sprejetih: SEM, XRD, XPS, itd.) 1 INTRODUCTION The purpose of the introduction is to supply sufficient background information to allow the reader to understand and evaluate the results of the present study. It should state briefly and clearly the purpose in writing paper, present as clearly as possible the nature and scope of the problem investigated, review recent literature to orient the reader, state the method of the investigation, and if necessary, the reasons for the choice of a particular method. 2 EXPERIMENTAL PART The experimental part must give full details of the experimental aparatus and the methods used in obtaining results. It must include accurate details relating to the equipment and materials used, incuding quantities, temperatures, times etc. It must provide sufficient information for a collegaue in the same field to reproduce the experiment. The data must be presented clearly and concisely, using graphs and tables. Figures and Tables must be written bold and not abbreviated throughout text. Figures and Tables must also follow the numerical order of appearance in the text. Once mentioned, tehy can be repeated. Equations must be written with Equation Editor and must be marked on the right-hand side of the text with numbers in round brackets. Throughout the text they must not be abbreviated.Physical quantities in the text have to be written as a plain text. Equation 1 (1) Equation 2 (2) 3 RESULTS Results section should provide an overall picture of the experiments without repeating any of the details in the experimental section. The data should be presented clearly and concisely, using graphs and tables. Figures and Tables should be written bold and not abbreviated throughout text. MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY Lepi pot 11, 1000 Ljubljana, Slovenia Phone: +386 1 4701 860, +386 1 4701 857, Fax: +386 01 4701 939, e-mail: mit@imt.si Editor-in-chief: dr. Paul John McGuiness ----------------------------------------------------------------------------------------------------------------- Figures and Tables must also follow the numerical order of appearance in the text. Once mentioned they can be repeated. 4 DISCUSSION The purpose of the discussion is to present the principles, relationships, and generalisations shown by the results. The discussion must make clear the significance of the results and compare the findings with previously published work. It must discuss the results not simply repeat them. 5 CONCLUSIONS The conclusion section should be short. It must present one or more conclusions which have to be drawn from the results and the subsequent discussion. Conclusions must be concise and easily understood by the reader. Acknowledgment 6 REFERENCES All references need a DOI number if it exists for the given reference. In the list of references monographs, articles in journals, journals, contributions to conference proceedings, patent documents, electronic monographs, articles and other electronic documents must be cited in accordance with the following examples: 1. Monographs: H. Ibach, H. Luth, Solid State physics, 2nd ed., Springer, Berlin 1991, 245 2. Articles in journals: T. Mauder, J. Stetina, Improvement of the casting of special steel with wide solid liquid interface, Mater. Tehnol., 50 (2015) 1, doi:10.17222/mit.2014.122 3. Contributions to conference proceedings, symposiums: I. Rak, M. Kocak, V. Gliha, N. Gubeljak, Fracture behaviour of over-matched high strength steel welds containing soft root layers, Proc. of the 2nd Inter. Symp. on Mis- Matching of Interfaces and Welds, Reinsford 1997, 627–641 4. Contributions in electronic form/online: Articles: M. P. Wnuk: Principles of fracture mechanics for space applications, http://mit.imt.si/Revija/izvodi/kzt996/wnuk.pdf, 30.01.2000 Other: http://www.imt.si, 15.03.2016 5. Standards: ISO 15787:2001(E) Technical product documentation, Heat treated ferrous part, Presentation and indication ISO Committee, Geneve MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY Lepi pot 11, 1000 Ljubljana, Slovenia Phone: +386 1 4701 860, +386 1 4701 857, Fax: +386 01 4701 939, e-mail: mit@imt.si Editor-in-chief: dr. Paul John McGuiness ----------------------------------------------------------------------------------------------------------------- MIT CHECKLIST PLEASE READ CAREFULLY THE STATEMENTS WRITTEN BELOW AND FULFILL IT WITH MARKING THE CHECKBOX. THIS DOCUMENTATION IS REQUIRED TO PROCESS YOUR SUBMITTED MANUSCRIPT, ENTITLED: __________________ FOR PUBLICATION IN MATERIALS AND TECHNOLOGY JOURNAL. Have you included a cover letter that makes clear the novelty of the investigation? Have you supplied the names of at least two potential referees? Is the manuscript written in accordance with the Instructions for Authors and with the MIT template? Does the number of figures and tables in your paper not exceed the required number according with Instructions for Authors? Are the figures and captions positioned in the text? Do you accept the terms of the MIT journal policy about the publication fee? Signature:________________________ SPECIAL NOTICE: From 1st October 2016 the journal Materials and Technology will be introducing a publication fee. All fees listed below, were agreed by the Editorial Board of Materials and Technology during its meeting on 15th June 2016: Publication fee -regular €300, 00 Publication fee – ICM&T conference paper €150, 00 Additional charge for colour printing € 80, 00 /page MIT journal subrscription (Slovenia) €42, 00 MIT journal subscription (abroad) €85, 00 MIT (Materials and Technology journal) Editorial Office Editor-in-Chief: dr. Paul John McGuiness Ljubljana, September 2016 OBVESTILO! Od 1.10.2016 dalje objava člankov v reviji Materiali in tehnologije (MIT) ni več brezplačna. Objavo članka Uredništvo revije MIT zaračunava na podlagi sklepa iz seje Uredniškega odbora revije Materiali in tehnologije/Materials and Technology, dne 15. junija 2016, po spodnjem ceniku: Objava prispevka 300, 00 € Objava konferečnega prispevka (ICM&T) 150, 00 € Doplačilo za barvni tisk 80, 00 €/stran Naročnina na revijo MIT (Slovenija) 42, 00 € Naročnina na revijo MIT (tujina) 85, 00 € Uredništvo revije Materiali in tehnologije (MIT) Glavni in odgovorni urednik: dr. Paul John McGuiness Ljubljana, september 2016 MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY Lepi pot 11, 1000 Ljubljana, Slovenia Phone: +386 1 4701 860, +386 1 4701 857, Fax: +386 01 4701 939, e-mail: mit@imt.si Editor-in-chief: dr. Paul John McGuiness ----------------------------------------------------------------------------------------------------------------- MIT CHECKLIST PLEASE READ CAREFULLY THE STATEMENTS WRITTEN BELOW AND FULFILL IT WITH MARKING THE CHECKBOX. THIS DOCUMENTATION IS REQUIRED TO PROCESS YOUR SUBMITTED MANUSCRIPT, ENTITLED: __________________ FOR PUBLICATION IN MATERIALS AND TECHNOLOGY JOURNAL. Have you included a cover letter that makes clear the novelty of the investigation? Have you supplied the names of at least two potential referees? Is the manuscript written in accordance with the Instructions for Authors and with the MIT template? Does the number of figures and tables in your paper not exceed the required number according with Instructions for Authors? Are the figures and captions positioned in the text? Do you accept the terms of the MIT journal policy about the publication fee? Signature:________________________ SPECIAL NOTICE: From 1st October 2016 the journal Materials and Technology will be introducing a publication fee. All fees listed below, were agreed by the Editorial Board of Materials and Technology during its meeting on 15th June 2016: Publication fee -regular €300, 00 Publication fee – ICM&T conference paper €150, 00 Additional charge for colour printing € 80, 00 /page MIT journal subrscription (Slovenia) €42, 00 MIT journal subscription (abroad) €85, 00 MIT (Materials and Technology journal) Editorial Office Editor-in-Chief: dr. Paul John McGuiness Ljubljana, September 2016 OBVESTILO! Od 1.10.2016 dalje objava člankov v reviji Materiali in tehnologije (MIT) ni več brezplačna. Objavo članka Uredništvo revije MIT zaračunava na podlagi sklepa iz seje Uredniškega odbora revije Materiali in tehnologije/Materials and Technology, dne 15. junija 2016, po spodnjem ceniku: Objava prispevka 300, 00 € Objava konferečnega prispevka (ICM&T) 150, 00 € Doplačilo za barvni tisk 80, 00 €/stran Naročnina na revijo MIT (Slovenija) 42, 00 € Naročnina na revijo MIT (tujina) 85, 00 € Uredništvo revije Materiali in tehnologije (MIT) Glavni in odgovorni urednik: dr. Paul John McGuiness Ljubljana, september 2016 E L E C T R O N I C A C C E S S http ://mit. imt. s i ISSN 1580-2949 M IN M ATER. TEH N O L. LETN IK VO LU M E [TEV. N O . STR. P. LJU BLJA N A SLO VEN IJA SEP.-O KT. SEP.-O C T. 51 5 707-879 ISSN: 1580-2949 UDK: 669+666+678+53 2017 5