VSEBINA – CONTENTS
^estitka/Congratulations
F. Vodopivec . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3
PREGLEDNI ^LANKI – REVIEW ARTICLES
Polyhydroxyalkanoates: biodegradable polymers and plastics from renewable resources
Polihidroksialkanoati: biorazgradljivi polimeri in plastike iz obnovljivih virov
M. Koller, A. Salerno, A. Muhr, A. Reiterer, G. Braunegg . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5
Challenges and advantages of recycling wrought aluminium alloys from lower grades of metallurgically clean scrap
Recikliranje gnetnih aluminijevih zlitin iz nizkocenovnih vrst metalur{ko ~istega odpadnega aluminija
V. Kevorkijan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13
IZVIRNI ZNANSTVENI ^LANKI – ORIGINAL SCIENTIFIC ARTICLES
Graphite-flake carbon-black-reinforced polystyrene-matrix composite films deposited on glass-fiber woven fabrics as plane
heaters
Kompozit polistirena, oja~an z grafitnimi luskami in sajami, nanesen na tkanino iz steklenih vlaken za plo{~ate grelnike
M. Erol, E. Çelik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25
Postopki priprave za proizvodnjo trdnih goriv iz nenevarnih odpadkov
The processes of preparation for the production of solid fuels from non-hazardous waste
J. Ekart, N. Samec, F. Kokalj, B. Polanec . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29
Problems associated with a robot laser cell used for hardening
Problematika robotskega laserskega kaljenja
M. Babi~, M. Milfelner, I. Beli~, P. Kokol . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37
A burnishing process based on the optimal depth of workpiece penetration
Postopek gladilnega valjanja na osnovi optimalne globine prodiranja v obdelovanec
Dj. Vukelic, D. Miljanic, S. Randjelovic, I. Budak, D. Dzunic, M. Eric, M. Pantic . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43
Calculation of the lubricant layer for a coarse surface of a band and rolls
Izra~un sloja maziva na grobi povr{ini traku in valjev
D. ]ur~ija, F. Vodopivec, I. Mamuzi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53
Structural and magnetic properties of cerium-doped yttrium-iron garnet thin films prepared on different substrates using the
sol-gel process
Strukturne in magnetne lastnosti s cerijem dopirane itrij-`elezove garnetne tanke plasti, izdelane s sol-gel postopkom na razli~nih podlagah
Y. Öztürk, M. Erol, E. Çelik, Ö. Mermer, G. Kahraman, I. Avgýn. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 59
Structural stability and electronic properties of AlCu3, AlCu2Zr and AlZr3
Stabilnost strukture in elektronske lastnosti AlCu3, AlCu2Zr in AlZr3
R. Cheng, X. Y. Wu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 65
Synthesis of a Fe3O4/PAA-based magnetic fluid for Faraday-rotation measurements
Sinteza magnetne teko~ine na osnovi Fe3O4/PAA za meritve Faradayeve rotacije
S. Küçükdermenci, D. Kutluay, Ý. Avgýn . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 71
A new approach to evaluating the chemical micro-heterogeneity of a continuously cast steel slab
Nov na~in ocenjevanja kemijske mikroheterogenosti kontinuirno ulitih slabov
J. Dobrovská, F. Kavi~ka, V. Dobrovská, K. Stránský, H. Francová . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 79
Heat-flux computation from measured-temperature histories during hot rolling
Ra~unanje toplotnega toka na osnovi preteklih meritev temperature med vro~im valjanjem
J. Ondrouskova, M. Pohanka, B. Vervaet . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 85
ISSN 1580-2949
UDK 669+666+678+53 MTAEC9, 47(1)1–141(2013)
MATER.
TEHNOL.
LETNIK
VOLUME 47
[TEV.
NO. 1
STR.
P. 1–141
LJUBLJANA
SLOVENIJA
JAN.–FEB.
2013
Thermal defunctionalization of an oxygen-plasma-treated polyethersulfone
Termi~na defunkcionalizacija polimera polietersulfona, obdelanega v kisikovi plazmi
M. Mozeti~, A. Vesel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 89
Diagnostics of the microstructural changes in sub-zero-processed Vanadis 6 P/M ledeburitic tool steel
Diagnostika sprememb mikrostrukture pri kriogeni obdelavi P/M ledeburitnega orodnega jekla Vanadis 6
J. Sobotová, P. Jur~i, J. Adámek, P. Salabová, O. Prikner, B. [u{tar{i~, D. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 93
STROKOVNI ^LANKI – PROFESSIONAL ARTICLES
Utilization of geopolymerization for obtaining construction materials based on red mud
Uporaba geopolimerizacije za pridobivanje gradbenega materiala na osnovi rde~ega blata
M. Vuk~evi}, D. Turovi}, M. Krgovi}, I. Bo{kovi}, M. Ivanovi}, R. Zejak . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 99
Analysis of process parameters for a surface-grinding process based on the Taguchi method
Analiza procesnih parametrov postopka bru{enja povr{ine s Taguchijevo metodo
M. K. Külekcý . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 105
Characterization of an Cu-Cr-Zr alloy synthesized with the powder-metallurgy technique
Karakterizacija zlitine Cu-Cr-Zr, sintetizirane s tehniko pra{ne metalurgije
M. Ipek. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111
Cutting-tool recycling process with the zinc-melt method for obtaining thermal-spray feedstock powder (WC-Co)
Postopek recikliranja orodij za rezanje z metodo taljenja v Zn za pridobivanje prahu WC-Co za termi~no nabrizgavanje
E. Altuncu, F. Ustel, A. Turk, S. Ozturk, G. Erdogan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 115
Performance testing of an optical ground wire composite
Preizku{anje zmogljivosti kompozitnega podzemnega opti~nega kabla
S. Karabay, E. A. Güven, A. T. Ertürk . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 119
Investigation of the cooling process with nanofluids according to ISO 9950 and ASTM D 6482 standards
Preiskava postopka ohlajanja v nanoteko~ini po standardih ISO 9950 in ASTM D 6482
J. @upan, D. Landek, T. Filetin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 125
Analysis of the effects of metamaterials on the radio-frequency electromagnetic fields in the human head and hand
Analiza u~inkov metamaterialov na radiofrekven~no elektromagnetno polje v ~love{ki glavi in roki
M. R. I. Faruque, M. T. Islam, N. A. N. Mohamed, B. Yatim, M. A. M. Ali . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 129
Possibilities for desulphurization of an alloy steel in a VOD device while using chemical heating
Mo`nost raz`vepljanja legiranega jekla v VOD-napravi z izkori{~anjem kemijskega ogrevanja
K. Michalek, L. ^amek, K. Gryc, T. Huczala, V. Troszok . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 135
DOGODKI – EVENTS
Promocija {tudija in raziskav materialov
Promotion of study and research of materials
M. Torkar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 141
^estitka
Mednarodno zdru`enje za toplotno obdelavo in
in`enirstvo povr{in podeljuje ~astno ~lanstvo IFHTSE
posameznikom, ki prispevajo izreden, globalno priznan
in pomemben prispevek k razvoju toplotne obdelave in
in`enirstva povr{in. Izr. prof. dr. Vojteh Leskov{ek je bil
izvoljen za ~astnega ~lana IFHTSE na 20. kongresu, ki je
bil 22. oktobra 2012 v Pekingu na Kitajskem. Priznanje
temelji na njegovem dolgoro~nem prispevku pri {tudiju
toplotne obdelave in in`enirstva povr{in z mo~nim
poudarkom na lomni `ilavosti, vakuumski toplotni
obdelavi orodnih in hitroreznih jekel z uporabo kriogene
tehnike, kot tudi termomehanske obdelave.
Slovenija ima spo{tljivo tradicijo v proizvodnji jekel,
ki dosegajo optimalne uporabne lastnosti po primerni
toplotni obdelavi, od jekel za mo~no obremenjene dele
strojev in dele motorjev za vozila do malo legiranih in
hitroreznih jekel. Razli~na orodja, od livarskih kokil do
orodij za utopno kovanje, so klju~nega pomena za
kvaliteto in konkuren~nost industrijskih podjetij, ki
ve~ino svoje proizvodnje prodajo v razvitih zahodno-
evropskih de`elah.
Po ustanovitvi prvega Laboratorija za vakuumsko
toplotno obdelavo leta 1984 na Metalur{kem in{titutu v
Ljubljani, sedaj In{titutu za kovinske materiale in tehno-
Materiali in tehnologije / Materials and technology 47 (2013) 1, 3–4 3
Predsednik IFHTSE prof. XU Kevel izro~a priznanje Fellow recognition izr. prof. dr. Vojtehu Leskov{ku
The President of IFHTSE, Prof. XU Kevel, presenting the Fellow recognition to Prof. Dr. V. Leskov{ek
Congratulations
The International Federation of Heat Treatment and
Surface Engineering awards the honour of Fellow of
IFHTSE to individuals who have made outstanding,
globally recognised and significant contributions to the
development of heat treatment and surface engineering.
Prof. Dr. Vojteh Leskov{ek was elected as a Fellow of
IFHTSE at the 20th Congress in Beijing, China on the
22nd of October 2012. The recognition is based on his
long-term contribution to the study of heat treatment and
surface engineering with a strong emphasis on fracture
toughness, the vacuum heat treatment of tool and
high-speed steels, and the use of sub-zero as well as
thermochemical treatments.
Slovenia has a respected tradition in the production
of steels that obtain optimal properties for use after a
proper thermal treatment, from steels for heavily loaded
parts of machinery and motors for parts of vehicles as
well as low-alloyed and high-speed steels. Different
tools, from casting moulds to forging dies, are essential
for the quality and competitiveness of industrial
companies marketing their products, mostly with exports
to developed Western European countries.
After establishing the first laboratory for vacuum
thermal treatment at the Metallurgical Institute of Ljub-
ljana, now the Institute of Metals and Technology, in
logije, je izr. prof. dr. Leskov{ek prevzel te`ko nalogo, da
bi pokazal mo`nosti nove tehnologije in jo prenesel v
industrijo, ki je uporabljala konvencionalne postopke in
opremo za toplotno obdelavo. Z upo{tevanjem mo`nosti
razpolo`ljive vakuumske pe~i se je obrnil k proizva-
jalcem in uporabnikom livarskih orodij, kova{kih orodij
in orodij za {tancanje. Po uspehu v manj{ih podjetjih se
je tehnologija raz{irila tudi v ve~ja podjetja in izvoznike,
ve~ vakuumskih naprav za toplotno obdelavo je nado-
mestilo konvencionalne naprave za toplotno obdelavo v
ve~jih industrijskih podjetjih. Uspeh na podro~ju
vakuumske toplotne obdelave je bil dobra podlaga za
podporo ve~jih industrijskih podjetij in{titutu pri nabavi
naprave za tehnologijo obdelave povr{in v plazmi v letu
1993. Tudi uporaba te tehnologije v industrijskih pod-
jetjih v Sloveniji je v porastu.
Poleg uvajanja in raz{irjanja teh dveh novih tehno-
logij so pomembni tudi rezultati izr. prof. dr. Leskov{ka
na podro~ju raziskav postopkov toplotne obdelave.
Opazen dose`ek je bilo uvajanje metode za dolo~anje
lomne `ilavosti trdih in krhkih jekel, posebno {e hitro-
reznih in mo~no legiranih orodnih jekel. Rotacijska
simetrija vzorcev omogo~a, da se izognemo nenatan~no-
sti, povezani z nesimetri~no hitrostjo ohlajanja vzorcev,
in izra~un kriti~ne velikosti napake, ki vpliva na
zdr`ljivost industrijskih orodij. Pomemben uspeh je tudi
razvoj ena~be, ki povezuje lomno `ilavost hitroreznih in
ledeburitnih jekel s trdoto po Rockwellu, vsebnostjo
zaostalega avstenita, z dele`em in razdaljo med karbidi,
kot tudi z modulom elasti~nosti. [tevilni ~lanki s citira-
njem, uporabnost in znanstveni dose`ki so raz{irili ime
izr. prof. dr. Vojteha Leskov{ka v skupnostih za toplotno
obdelavo in orodna jekla ter podprli njegovo izvolitev za
~astnega ~lana IFHTSE. Njegov uspeh je dober zgled za
mlade, da s trdim delom lahko tudi znanstvenik iz
majhne dr`ave dose`e mednaroden ugled in priznanje.
Priznanje in izvolitev izr. prof. dr. Vojteha Leskov{ka
za ~astnega ~lana IFHTSE odlikuje tudi In{titut za
kovinske materiale in tehnologije, Ljubljana.
Iskrene ~estitke izr. prof. dr. Vojtehu Leskov{ku v
imenu kolegov in sodelavcev in{tituta ter industrijskih
podjetij.
prof. dr. Franc Vodopivec
4 Materiali in tehnologije / Materials and technology 47 (2013) 1, 3–4
1984, Prof. Dr. Leskov{ek took over the difficult task of
demonstrating the possibilities of the new technology
and its expansion in industrial companies using con-
ventional processes and equipment for thermal
treatment. Considering the possibilities of the available
vacuum furnaces, he turned to producers and users of
casting moulds, forging and stamping dies. Thanks to
successes in smaller companies, the technology
expanded to larger companies and exporters and several
vacuum-treatment facilities substituted conventional heat
treatment in larger industrial companies. The success in
vacuum heat treatment was a good reference for the
support of several industrial companies for the institute
to acquire the equipment for plasma surface treatment
technology in 1993. The use of this technology in
industrial companies in Slovenia is also expanding.
Besides the introduction and expansion of two new
technologies of remarkable importance are the results of
Prof. Dr. Leskov{ek in the field of the research of
heat-treatment processes. A notable achievement was the
introduction of a method for determining the fracture
toughness of hard and brittle steels, especially
high-speed and high-alloyed tool steels. The rotational
symmetry of the specimens makes it possible to avoid
inaccuracies related to the uneven cooling rate of
specimens and the calculation of a critical defect size
that affects the useful life of industrial tools. Another
remarkable success is the development of an equation
that relates the fracture toughness of high-speed and
ledeburitic tool steels to Rockwell hardness, the content
of residual austenite, the fraction of carbide particles and
spacing as well as the modulus of elasticity. A number of
articles with the quoted and other original applied and
scientific findings have spread the name of Prof. Dr.
Vojteh Leskov{ek in the heat-treatment and tool-steels
communities and supported his election to Fellow of
IFHTSE. His success is a good example for youth, that
with hard and properly focused work, scientists from
small countries can also acquire an international
reputation and recognition.
The recognition and election of Prof. Dr. Vojteh
Leskov{ek as Fellow of IFHTSE also honours the
Institute of Metals and Technology, Ljubljana, Slovenia.
We would like to offer our sincere congratulation to
Prof. Dr. Vojteh Leskov{ek on behalf of his colleagues
and collaborators from the institute and industrial
companies.
Prof. Dr. Franc Vodopivec
M. KOLLER et al.: POLYHYDROXYALKANOATES: BIODEGRADABLE POLYMERS AND PLASTICS ...
POLYHYDROXYALKANOATES: BIODEGRADABLE
POLYMERS AND PLASTICS FROM RENEWABLE
RESOURCES
POLIHIDROKSIALKANOATI: BIORAZGRADLJIVI POLIMERI IN
PLASTIKE IZ OBNOVLJIVIH VIROV
Martin Koller1,2, Anna Salerno1, Alexander Muhr1, Angelika Reiterer1,
Gerhart Braunegg2
1Graz University of Technology, Institute of Biotechnology and Biochemical Engineering, Petersgasse 12, 8010 Graz, Austria
2ARENA – Association for resource efficient and sustainable technologies, Inffeldgasse 23, 8010 Graz, Austria
martin.koller@tugraz.at
Prejem rokopisa – received: 2012-07-24; sprejem za objavo – accepted for publication: 2012-08-28
Polyhydroxyalkanoates (PHAs) attract considerable attention as sustainable "green plastics" with a real potential to replace their
petrol-based competitors in some applications in the not-too-distant future. To reach this goal PHAs must be able to compete
with the established petrol-based plastics in both technical and economic terms. The current PHA production is based on prized
substrates of high nutritional value such as sucrose, starch or vegetable oils. An alternative, carbon-rich industrial waste can be
used as a suitable feedstock. This would contribute to making PHAs economically competitive and would avoid the conflict
with human nutrition or animal feeding.
Consequently, the decision about the location of the PHA-production facilities depends on the preferable in-house availability of
such waste streams. The issue of competitive, large-scale PHA production in Europe was the topic of the ANIMPOL and the
WHEYPOL projects. Both intended to develop novel processes for the transformation of abundant, locally available, renewable
wastes. In the ANIMPOL case, waste lipids from slaughterhouses are converted to glycerol and a mixture of saturated and
unsaturated fatty acid esters (FAEs), better known as biodiesel. The production of saturated FAEs is 50 000 t per year in Europe,
decreasing the biodiesel performance as an engine fuel. However, within the project it was demonstrated that they can be
efficiently metabolized to PHAs. In the WHEYPOL project attention was focused on 1.4 × 108 t per year of whey from dairies.
This waste is of limited use and causes environmental concern. However, lactose, the main carbohydrate found in whey, can be
used as a substrate in the WHEYPOL bioprocesses. These strategies demonstrate the feasibility of making "green plastics"
competitive by integrating their manufacturing directly into the existing production lines, where the convertible waste streams
accrue.
Keywords: animal waste, biodiesel, biopolymers, lipids, polyhydroxyalkanoates, whey
Polihidroksialkanoati (PHA) vzbujajo posebno pozornost kot trajnostna "zelena plastika" z realno mo`nostjo, da nadomestijo
svoje tekmece, ki temeljijo na nafti, za dolo~ene namene uporabe v bli`nji prihodnosti. Za dosego tega cilja morajo biti PHA
sposobni tekmovati tako po tehni~nih kot po ekonomskih vidikih z uveljavljenimi plastikami na osnovi nafte. Sedanja
proizvodnja PHA temelji na dragih podlagah z veliko hranilno vrednostjo, kot so saharoza, {krob ali rastlinska olja. Kot
alternativna surovina se lahko uporabi z ogljikom bogate industrijske odpadke. To bi prispevalo k bolj ekonomi~ni izdelavi PHA
in bi prepre~ilo konflikt s prehrano ljudi in krmo `ivali.
Posledi~no je odlo~itev o lokaciji izdelave PHA odvisna od razpolo`ljivosti in virov takih odpadkov. Vpra{anje konkuren~ne
proizvodnje PHA v velikih koli~inah v Evropi je bila tema projektov ANIMPOL in WHEYPOL. Oba sta imela namen razviti
nove postopke predelave obilnih in lokalno razpolo`ljivih obnovljivih odpadkov. V primeru ANIMPOL-a so bili odpadni lipidi
iz klavnic pretvorjeni v glicerol in me{anico estrov (FAE) nasi~enih in nenasi~enih ma{~obnih kislin, bolj poznani kot biodizel.
Proizvodnja nasi~ene FAE v Evropi je 50 000 t na leto in zmanj{uje zmogljivost biodizla kot pogonskega goriva. Projekt je
pokazal, da jih lahko pretvorimo v PHA. Pri projektu WHEYPOL je bila pozornost usmerjena na 1,4 × 108 t na leto sirotke iz
mlekarn. Ta odpadek ima omejeno uporabo in povzro~a okoljske skrbi. Vendar pa se laktoza, glavni ogljikovodik v sirotki, lahko
uporabi kot osnova pri WHEYPOL-bioprocesih. Te strategije ka`ejo, da je mogo~a proizvodnja konkuren~ne "zelene plastike" z
vklju~itvijo njihove izdelave neposredno v proizvodne linije, kjer se vredni odpadki pojavljajo.
Klju~ne besede: `ivalski odpadki, biodizel, biopolimeri, lipidi, polihidroksialkanoati, sirotka
1 INTRODUCTION
1.1 "Green plastics" today
Currently, we are experiencing a remarkably dynamic
biopolymer market with an impressive increase in the
volume and range of the products: in 2010, the market
value reached a magnitude of 1010 US dollars with an
obvious upward trend. Only from 2008 to 2015, the
global production of biopolymers is estimated to
increase from 180 kt to 1710 kt (Plastic Additives and
Compounding 2008). This positive development is
unfortunately associated with an emergence of various
plastics that are labeled with the stylish attribute "green
plastic" by the manufactures, but display properties that
do not match the strict definitions by which they could
be classified as biobased, biodegradable, compostable or
biocompatible. On the other hand, polyhydroxyal-
kanoates (PHAs) are natural, thermoplastic, aliphatic
biopolyesters that fully comply with these requirements.1
Among all bio-based plastics, they are unique by being
entirely produced and degraded by living cells.2 The
spectrum of their potential applications ranges from
simple packaging to high-quality materials for niche
uses.
Materiali in tehnologije / Materials and technology 47 (2013) 1, 5–12 5
UDK 577.11:577.115 ISSN 1580-2949
Review article/Pregledni ~lanek MTAEC9, 47(1)5(2013)
1.2 General aspects of PHAs
PHAs are synthesized and stored as intracellular
granules by a wide range of prokaryotic genera using
various renewable feedstocks. Predominately, the
conversion of hexoses, pentoses, starch, sucrose, lactose,
maltose, lipids, alcohols, or methane is reported in
literature.3–5 Figure 1 provides a scanning transmission
electron microscopy (STEM) picture of Cupriavidus-
necator cells cultivated on carbohydrates in a continuous
process. PHA inclusions are well visible as bright,
refractive granules amounting to the mass fraction 52 %
of the cell mass.
Principally, a high intracellular energy charge,
characterized by an elevated pool of acetyl-CoA,
NAD(P)H, ATP and others, promotes a PHA formation.
Such conditions result from a sufficient availability of an
external carbon source and a restricted supply of the
growth-essential substrates such as nitrogen, phosphate,
dissolved oxygen, or certain micro-components.3,6,7
PHAs carry out important biological tasks, mainly acting
as intracellular-energy and carbon reserves that can be
re-utilized under the conditions of carbon starvation.8
They are crucial for the regulation of the intracellular
energy flow, routing the carbon compounds to the
metabolic pathways and acting as protective factors
against the environmental pressures like osmotic shock,
UV irradiation, desiccation, heat and oxidative stress.9
1.3 Material properties and downstream processing
PHAs are polyesters consisting mainly of enantio-
merically pure R-configured 3-hydroxyalkanoate (3HA)
monomers; the general chemical structure is shown in
Figure 2.
Depending on the microbial production strain, the
nutrient supply and the process parameters during a
biosynthesis, PHA-polymer chains contain 100 to 100
000 3HA units. Among all known PHAs, the
homopolyester of R-3-hydroxybutyrate (3HB)-poly
([R]-3-hydroxybutyrate) (PHB) is most intensely studied
and understood. Unfortunately, PHB exhibits a rather
high crystallinity that hampers its processability. The low
difference between the decomposition temperature
(around 270 °C) and the high melting point (around 180
°C) provides a "window of processability" that is too
narrow for many processing techniques, e.g., a melt
extrusion or blowing of plastic films. This can be
overcome by interrupting the crystalline PHB matrix
with an incorporation of additional PHA building blocks,
mainly 3-hydroxyvalerate (3HV) or an achiral building
block of 4-hydroxybutyrate (4HB). The resulting
co-polyesters display enhanced material properties in a
broader range of applications. The exact material
properties are strongly dependent on the monomer
composition of co-polyesters. This composition can be
triggered during a PHA biosynthesis by feeding the
precursor substrates to obtain the desired building
blocks. For example, 3HV building blocks are produced
by many PHA-accumulating strains supplying odd-num-
bered fatty acids as precursors that lead to 3HV in the
synthesized PHAs.8
After a production of the PHA-rich cells, efficient
downstream processing is needed. This includes sepa-
rating the biomass from the liquid phase, a PHA
recovery from the biomass, and product refining. A PHA
recovery from the cells may include a solvent or a
supercritical fluid extraction, a chemical or enzymatic
digestion of the cell wall, using high cell fragility of
special microbial species, or a mechanical cell disrup-
tion.10 The choice of the best method mainly depends on
the required product purity that will satisfy the envisaged
application. The recovered biopolyesters can be pro-
cessed with the conventional equipment well established
in the polymer industry. This technological path leads to
proper biodegradable and biobased substitutes for a
variety of "classical" petrol-based plastics such as ther-
moplastics, elastomers, and even latex rubbers.11–14
M. KOLLER et al.: POLYHYDROXYALKANOATES: BIODEGRADABLE POLYMERS AND PLASTICS ...
6 Materiali in tehnologije / Materials and technology 47 (2013) 1, 5–12
Figure 2: General chemical structure of PHAs. The asterisk (*)
indicates the chiral center, n: number of building blocks in the
polyester chain (for natural PHAs: n = 100 to 10000), m: number of
carbon atoms in the monomer backbone; in the case of 3HAs, m = 1,
R: side chain of the building blocks; in the case of 3HB, R = -CH3.
Slika 2: Splo{na kemijska struktura PHA. Zvezdica (*) ozna~uje
kiralni center, n: {tevilo gradnikov v poliestrski verigi (pri naravni
PHA: n = 100 do 10000), m: {tevilo ogljikovih atomov v monomerni
hrbtenici; v primeru 3HAs, m = 1, R: stranska veriga gradnikov; v
primeru 3HB, R = -CH3.
Figure 1: STEM picture of Cupriavidus-necator cells harboring PHA
biopolymers (magnification 65.000x)
Slika 1: STEM-posnetek celic Cupriavidus necator, ki zadr`ujejo
PHA-biopolimere (pove~ava 65000-kratna)
2 CHALLENGES IN THE PRODUCTION OF
PHAs AND THE ATTEMPTS TO OVERCOME
THEM
2.1 Global PHA production
Despite the enormous efforts during the last decades
dedicated to the improvements in the PHA production,
characterization and processing on the laboratory and
pilot scale, their ultimate market penetration has not yet
taken off. However, serious activities undertaken in the
past by ICI-Marlborough (UK), Chemie Linz (Austria),
and, more recently, by Monsanto, Metabolix (both USA),
PHBISA (Brazil), Tepha (USA) and Tianan or Shandong
(PR China) constitute laudable exceptions. Due to an
obvious gap between the often highly enthusiastic public
announcements made by the manufacturers citing their
capacities, and the actual PHA production, it is
challenging to collect reliable data on the quantities of
PHAs that are currently produced on the (semi)industrial
scale. Realistically, it can be estimated that today PHAs
cover about 5 % of the world’s biolastic market, and less
than 0.05 % of the overall global plastic production.
2.2 Economics of the PHA production: decisive factors
The major part, up to a half, of the entire PHA
production costs is attributed to the cost of carbon
substrates. On the one hand, from the ethical point of
view, the explosive nutritional situation of the mankind
in many global regions strongly discourages the
utilization of various renewable feedstocks that can be
used for food, for the production of chemicals, plastics
or fuels; starch or edible oils are the prime examples. On
the other hand, there are a number of diverse industrial
carbon-rich wastes that do not interfere with the nutrition
chain, and, at the same time, cause disposal problems
and costs. Their utilization as carbon substrates for the
microbial processes is a viable strategy for overcoming
this ethical conflict and can be considered as the most
promising approach in making PHAs economically
competitive. Waste renewable resources like ligno-
cellulosics, carbohydrates, waste lipids or alcohols, are
mainly produced in agriculture and connected industrial
branches.8,12,15,16
In addition to introducing the substrates, costs have
to be reduced by optimizing the downstream processing
required for the PHA recovery after a cell harvest. As
intracellular products, PHAs have to be separated from
the surrounding non-PHA cell mass (NPCM, also known
as the residual biomass) that mainly consists of proteins,
lipids, nucleic acids and polysaccharides. Here, a high
usage of hazardous solvents and enormous energy
demand still constitute the state-of-the-art, somewhat
contradicting the patterns of sustainability and economic
feasibility.3,17
Apart from the selection of raw materials and the
improvements in downstream processing, enhanced
productivity reached through the design of an optimal
engineering set-up is indispensable for the final break-
through of PHAs on the market. Batch and fed-batch
discontinuous fermentation modes are currently most
commonly used techniques for the microbial PHA
production.18–20 In contrast, the continuous production
mode is a well-known tool for achieving high produc-
tivities, lower production costs and a constant product
quality in biotechnological processes. Therefore, an
increasing research is focusing on investigating and
assessing the potential of continuous PHA-production
processes.21–23 Recently, based on the kinetic conside-
rations regarding biomass growth and PHA accumu-
lation, the continuous production of PHB in a five-stage
bioreactor cascade (5-CSTR) was investigated.23 This
5-CSTR acts as a device-related engineering substitute
for a tubular bioreactor that is theoretically the best
match for the process-engineering requirements for an
efficient PHA production from a kinetic point of view.
As the main results, the authors report on a high PHB
productivity of 1.85 g/(L h), and a constant and superior
product quality concerning the mechanical properties
and molecular masses.23
3 CURRENT ROUTES IN PHA RESEARCH
Current research in the field of PHAs focuses on
several key topics. The application of growth additives
that shorten the time for the production of a catalytically
active biomass is a pre-requisite for enhancing the entire
volumetric productivity of the process. Inexpensive
growth additives can be found in agriculture, e.g., pro-
teinaceous materials or side streams from the cultivation
of green grass land, and were already tested successfully
on the laboratory scale.24
Efforts in the field of genetic engineering mainly
intend to increase the volumetric PHA productivity and
achieve higher molar masses of the biopolymers. This
can be accomplished with a knock-out of enzymes
responsible for the intracellular PHA degradation. In
addition, metabolic bottle necks that can limit a fast and
complete substrate conversion can be overcome by
genetic modifications. Enhancement of the microbial
oxygen uptake may be improved by inserting the genes
encoding catalase or peroxidases.25
In the field of downstream processing, together with
lower energy consumption, environmentally safe and
efficient solvents are investigated in order to enhance the
recovery of PHAs from the cells. This goes in parallel
with an examination of the novel biological lysis
methods and enhanced strategies for a mechanical cell
disruption. For an efficient polymer recovery, an increase
in both the intracellular polymer content as well as the
PHA granule size is beneficial. These factors are deter-
mined during the PHA biosynthesis.26 In addition, the
remaining NPCM has to be converted in a sustainable,
value-adding way. Currently, the research in this direc-
tion is devoted to the anaerobic digestion of NPCM in
M. KOLLER et al.: POLYHYDROXYALKANOATES: BIODEGRADABLE POLYMERS AND PLASTICS ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 5–12 7
biogas plants, its application in agriculture as a "green
fertilizer", or to the chemical or enzymatic digestion of
NPCM into a nutrient-rich cocktail to be used for
subsequent microbial cultivations. Downstream process-
ing can also be facilitated by genetic modification;
boosted nuclease excretion after a cell disruption results
in decreased amounts of nucleic acids in the medium,
leading to a lower viscosity that facilitates the separation
of the PHA granules from the surrounding liquid phase.
The technological drawbacks of the bio-production
itself can be handled by an application of the robust,
microbial production strains that remain genetically
stable for a long time under continuous cultivation
conditions. At the same time, they should resist the
contamination with the microbial competitors that, by
endangering the whole production set-ups, represent a
high risk.8 Using extremophilic species, like the halo-
philic archaeon Haloferax mediterranei, minimizes the
normally indispensable, energy-demanding sterilization
precautions for the PHA production set-ups.27,28
In future, continuous PHA production should not
only aim to increase the volumetric productivity, but also
to open the door for tailor-made material properties by
fine-tuning the polyester composition. This can be
accomplished with the formation of block-copolymers,
where a sequential arrangement of softer and harder
polymer segments can result in well-adjusted novel poly-
meric materials. Here, a multistage bioreactor cascade
for the PHA production is presented with potentially
adequate process engineering equipment.23
During the last decades, the preparation of nano-
composites and composites involving natural fibers has
become one of the key segments in the biopolymer
science. To enhance material properties, PHAs can be
processed together with a variety of compatible matters,
resulting in the creation of novel materials. For this
purpose, polymeric materials like poly(vinyl alcohol),
poly(lactic acid), poly(-caprolactone), synthetic ana-
logues of PHAs (e.g., atactic PHB), inorganic (e.g., clays
or calcium carbonate) and organic (bagasse, wheat flour,
fruit peels, fruit fibers, saw dust and straw) fillers of agri-
cultural origin were already tested.29,30 Nanocomposites
have a potential to improve special polymer properties,
such as gas permeability and thermal and mechanical
characteristics. Here, only rather small amounts of a
filler (mainly modified clay) are needed to improve the
material performance.29 Natural-fiber composites often
display excellent mechanical properties, and decrease the
density of the final product. In most cases, the fibers
from agricultural residues are used as fillers; their
application additionally enhances the biodegradability
and reduces the production cost of the final product.30
4 CASE STUDY 1: FROM SURPLUS WHEY TO
PHA BIOPOLYESTERS
Whey is a surplus product in the cheese industry.
Casein precipitates from milk enzymatically or by
acidification. This transformation results in curd cheese
(mainly caseins) and full-fat whey that undergoes a
degreasing and concentration procedure. From the
resulting whey concentrate, whey retentate (protein
fraction) is removed; the remaining whey permeate
(carbohydrate fraction) contains about 80 % of the
lactose originally included in the milk.
In 2008, 1.6 × 108 t of whey were produced accord-
ing to the estimates from OECD and FAO. The produc-
tion has increased continually in the past and this trend is
expected to continue in the future. The highest quantities
of whey are available in the northern hemisphere with
roughly 4 × 107 t in the USA, 2.2 × 105 t in Canada, and
5 × 107 t in the EU-25.31
Due to its quantities, whey is not only an excess raw
material, but it also causes disposal problems because of
its high biochemical oxygen demand.18,32 It accrues in the
volumes almost equal to the processed milk, which
means that the processing of one million liters of milk
into cheese, almost one million liters of whey has to be
treated. This exemplary quantity contains up to 50 t of
lactose as the main carbon ingredient in whey. Just in the
northern Italy, where a number of large dairies are
located, about 1 million liters of whey have to be
disposed off daily. This is often accomplished with a
direct release of untreated whey into the aquatic environ-
ments.
Hence, it was reasonable to design an integral
process for the PHA production from the surplus whey
including the following steps:24
• Set-up of the optimized upstream technology for the
feedstock whey
• Set-up of the optimized PHA-formation conditions
with the selected microbial strains using the whey
lactose as a carbon source
• Set-up of the biotechnological strategy for triggering
the polymer composition as a tool for fine-tuning
PHA properties
• Optimization of the downstream strategy for a PHA
recovery
• Detailed product characterization
• Economic and ecological appraisal of the entire
process and its steps
• Evaluation of the applications of the products
It is obvious that the discussed points can only be
achieved, if there are intense interactions between the
experts in different scientific fields: microbiology (a
search for the ideal microbial production strain), bio-
technology (the fermentation technology as the core
part), chemical engineering (tailor-made downstream
processing, design of industrial plant), polymer science
(an on-going characterization of the products), and
life-cycle assessment (LCA).
Within the FP5 Growth Program of the EU, the
WHEYPOL project realized these ideas between 2001
and 2004. The demanded interactions between different
research areas were put into practice with the formation
M. KOLLER et al.: POLYHYDROXYALKANOATES: BIODEGRADABLE POLYMERS AND PLASTICS ...
8 Materiali in tehnologije / Materials and technology 47 (2013) 1, 5–12
of a network consisting of seven scientific groups and
three industrial partners each specialized in a certain
field needed for the process development.
In a study that was accomplished within the WHEY-
POL project, Koller and colleagues27 compare the
potential of three prokaryotic wild-type strains for the
utilization of whey as carbon feedstock for the PHA
production. In this work the archaeon Haloferax
mediterranei and the eubacterial strains Pseudomonas
hydrogenovora and Hydrogenophaga pseudoflava were
investigated in bioreactors on the laboratory scale.
Among these organisms, Haloferax mediterranei
turned out to be the most promising candidate for an
eventual industrial-scale PHA production using whey.
This is due to the strain’s high robustness and stability;
the risk of microbial contamination during the cultivation
is restricted to the absolute minimum, thus a lot of
energy can be saved with lower sterility demands.
Additionally, the strain produces a P(3HB-co-8 %-3HV)
co-polyester directly from the 3HV-unrelated carbon-
source whey, avoiding the costs for odd-numbered acids
as the 3HV-related precursors.33 The strain grew well on
hydrolyzed whey permeate with the maximum specific
growth rate μmax of 0.11 1/h. PHAs were accumulated at
the maximum specific production rate of 0.08 g/(g h).
The conversion yield for whey sugar to PHA amounted
to 0.33 g/g. After a further optimization of the pro-
duction conditions, the productivity of this strain on
hydrolyzed whey permeate was increased to 0.09 g/(L h)
(the specific rate of 0.15 g/(g h); 16.8 g/L of biomass
containing 73 % of PHAs were finally obtained. By
co-feeding the precursors for 3HV and 4HB production
(pentanoic acid and -butyrolactone, respectively),
together with whey permeate as the main carbon source,
a high value P-(3HB-co-21.8 %-3HV-co-5.1 %-4HB)
terpolyester was completed by Haloferax mediterranei.
A detailed characterization of thermal properties and
molecular-mass distribution of the materials was
accomplished; the results for polymer characterization
indicate that the materials might be of special interest for
application in the medical field. The partial conversion of
whey sugars to 3HV units and the excellent polymer
characteristics (low melting temperature, high molecular
masses within a narrow distribution), together with a
simple recovery of PHAs from the cells, make the strain
especially interesting. The estimated production price
amounted to  2.82 per kg of PHAs. For a further
process improvement, the recycling of the highly saline
side streams should be tested and optimized.27,28
P. hydrogenovora features the disadvantages of low
final polymer contents, low productivity and product
yields due to a redirection of the carbon flux towards
unwanted by-products such as various organic acids.
Using this organism, the final PHB homopolyester con-
tent amounted to the mass fraction 12 % (qp: 2.9 mg/(g
h)). By co-feeding the 3HV-releated precursor penta-
noate, the strain accumulated 12 % of poly-3(HB-co-
21%-HV) (qp: 2.0 mg/(g h)).28,34
The research showed that H. pseudoflava produces
PHAs of a rather good quality (high molar masses of Mw
= 859 MDa and low polydispersities of 3.3) from whey
at acceptable specific production rates and yields, but is
not competitive with H. mediterranei in terms of strain
stability and robustness. In detail, using this strain, 40 %
of poly-3(HB-co-5 %-HV) in the cells with an addition
of pentanoic acid (qp: 12.5 mg/(g h)) was obtained.
Without the pentanoic acid, the strain accumulated 30 %
of the homopolyester PHB in the cells (qp: 16.0
mg/(g h)).28
5 CASE STUDY 2: SLAUGHTERING WASTE
FOR THE PHA PRODUCTION
Animal lipids from slaughtering and animal-
processing industries in Europe amount to more than
500000 t per year. It is estimated that using this waste
stream for biodiesel production through transesteri-
fication results in 50000 t of saturated biodiesel (SFAE)
fraction. SFAE causes problems in fuel mixtures due to
its precipitation at a low temperature, which may result
in filter plugging. Using SFAE as a carbon feed stock,
35000 t of PHAs can theoretically be produced annually
(a yield of 0.7 g of PHAs per g of SFAE). Considering
the entire amount of biodiesel that is currently produced
in the EU-25 (about 22 Mt), approximately 2 Mt of
glycerol phase (CGP) are available as the major
by-product of the conversion of lipids to biodiesel. This
is in considerable excess over the quantities of the
glycerol needed for its various classical applications. If
applied in the production of microbial PHA-accu-
mulating biomass, one can expect more than 0.4 g of
biomass or PHAs per gram of glycerol. Regarding the
entire quantity of glycerol from the animal-lipid trans-
esterification, one can estimate that more than 20000 t of
PHA-rich biomass can be produced annually. These
amounts are shown in Figure 3.
M. KOLLER et al.: POLYHYDROXYALKANOATES: BIODEGRADABLE POLYMERS AND PLASTICS ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 5–12 9
Figure 3: Amounts of available animal-waste lipids, and the potential
PHA production
Slika 3: Koli~ina razpolo`ljivih `ivalskih lipidov in mogo~a proiz-
vodnja PHA iz njih
In 2010, the project "Biotechnological conversion of
carbon-containing wastes for eco-efficient production of
high-added value products", acronym ANIMPOL, was
launched and funded through the FP7 of the EU.
Within this project, a sound industrial process for the
conversion of lipid-rich animal waste from the meat
processing industry into biodiesel is being developed
using innovative methods. Saturated biodiesel fractions
have a negative effect on biodiesel fuel properties, are
separated and, subsequently, used as feedstock for the
biotechnological production of PHAs. The remaining
unsaturated biodiesel fraction performs as an excellent
2nd generation biofuel. The significance of the project is
obvious considering the high quantities of animal-
derived waste in Europe. The project brings together
partners managing waste resources from slaughter-
houses, the rendering industry, biodiesel production, as
well as the polymer-processing industry.
Experimentally, the most promising results for the
PHA production from the animal-waste-derived SFAE
are obtained by using Cupriavidus necator, a well-known
wild-type strain from the Burkholderiaceae family. Fast
microbial growth, satisfying cell densities, and high
volumetric productivities of PHA homo- and co-poly-
esters, together with promising material characteristics,
were obtained.
Figure 4 illustrates the fermentation pattern of the
laboratory bioreactor-scale production of PHB with C.
necator using SFAE as the main carbon source. This
fermentation was accomplished by using SFAE as the
sole carbon source for both the biomass growth and the
PHA accumulation in the fed-batch cultivation mode (a
repeated supply of SFAE according to its conversion by
the cells). The maximum of 28.0 g/L of PHAs was
obtained, corresponding to the PHA share in a cell’s dry
mass that is 80.3 %. The specific growth rate of the
production strain amounted to 0.17 L/h. Especially the
yield of the biomass production from SFAE of more than
0.6 g of CDM per g of SFAE is exceptionally high if
compared to the well-known PHA substrates like sugars,
where the theoretical yield does not exceed 0.48 g/g.
This is due to the metabolic background of fatty acid
catabolism of the cells. Considering the early stage of the
process development, the high volumetric PHA
productivity of 0.94 g/(L h) for the entire process can
also be considered very promising if compared to the
available data for the industrial PHA production from
expensive substrates. Taking into account only the phase
of the predominant PHA accumulation after nitrogen
limitation, volumetric productivity was as high as 1.36
g/(L h). Regarding the specific volumetric productivity
during nitrogen-limited conditions, the value of 0.19 g/(g
h) was calculated. This is significantly higher than the
values reported before for the utilization of whey lactose.
During this phase, residual biomass (NPCM) concen-
tration remained constant at about 7 g/L (see Figure 4).
In addition, the produced PHA was a poly-(3HB-co-0.84
%-3HV) copolyester; here, odd-numbered fatty acids in
SFAE acted as 3HV-related precursor substrates. The
composition of the polyester during the fermentation and
the concentration of 3HV in the fermentation broth are
illustrated in Figure 5. It is well visible that, due to the
identical composition of the added SFAE during the
entire process, the share of 3HV in PHA (about 0.8 %) is
constant during the phase of nitrogen limitation.
6 CONCLUSIONS
The study presents strategies for upgrading industrial
wastes like surplus whey and residues from the
slaughtering and biodiesel industries to feedstocks for
biopolymer production. This approach can be regarded
as a promising route for making the entire
PHA-biopolymer-production process economically
competitive. The selection of the discussed raw materials
M. KOLLER et al.: POLYHYDROXYALKANOATES: BIODEGRADABLE POLYMERS AND PLASTICS ...
10 Materiali in tehnologije / Materials and technology 47 (2013) 1, 5–12
Figure 5: Poly-(3HB-co-0.84 %-3HV) production with C. nector on
SFAE: composition of the polyester (small bars) and 3HV
concentration (high bars) during the process
Slika 5: Proizvodnja poly-(3HB-co-0,84 %-3HV) s C. nector na
SFAE: sestava poliestra (mali stolpci) in koncentracija 3HV (veliki
stolpci) med postopkom
Figure 4: Fermentation pattern for the poly-(3HB-co-0.84 %-3HV)
production with C. nector on SFAE. The arrow at t = 7 h indicates the
limitation of the nitrogen supply. Parallel lines at the time axis and the
time curve for residual biomass indicate the end of the microbial
growth after the nitrogen depletion, provoking the redirection of the
intracellular carbon flow towards the PHA accumulation.
Slika 4: Model fermentacije za poli-(3HB-co-0,84 %-3HV) proiz-
vodnjo s C. nector na SFAE. Pu{~ica pri t = 7 h ka`e omejitev dobave
du{ika. Vzporedne linije na ~asovni krivulji za preostalo biomaso
ka`ejo na konec mikrobne rasti po iz~rpanju du{ika, kar spodbudi
preusmeritev toka medceli~nega ogljika proti kopi~enju PHA.
is especially important to provide European industry
with an advantage, on the one hand, to handle
tremendous waste streams through a value-added
conversion, and, on the other hand, to break new ground
in developing and commercializing bio-polymeric
materials.
Uniting the potential enhancements of each process
step, one can make substantial progress towards an
environmentally benign and cost-efficient technology.
The development of efficient biopolymer-production
processes needs a close cooperation between the experts
from industry and different scientific fields; hence, a
multidisciplinary approach is required. Decision makers
from companies, chemical engineers, microbiologists,
enzymologists, polymer scientists, genetic engineers, and
the experts involved in LCA and the Cleaner Production
studies have to join their respective expertise. Such an
approach has a potential to close the existing gaps
between the success achieved on the laboratory scale
that, in most cases, affects only singular aspects of the
biopolymer production, and the final success of
competitive bio-plastics on the market.
Regardless of the selected microbial production
strain, the facilities for the PHA production from whey
and animal-derived waste should be integrated into the
existing process lines of large dairies or biodiesel
companies, where the raw material actually accrues. This
can be considered as a viable strategy for minimizing the
production costs by taking profit through the synergistic
effects.
Due to the environmental considerations and
expected shortage of fossil reserves, the increasing
global demand for bio-plastics in the future is generally
undisputed. Available data about the laboratory-scale
developments shows a significant progress in combining
the environmental benefit of the future-oriented
bio-polyesters with the economic viability of their
production. This development will hopefully facilitate
the decisions of responsible policy-makers from various
renewable waste-generating industries and from the
polymer industry to break new ground in a sustainable
production. If successfully implemented, these seminal
biopolymers will be decoupled from the manufacturing,
where biopolyester and biofuel syntheses compete with
human nutrition, frequently lacking consideration of
environmental requirements and acceptable human
working conditions.
Acknowledgement
The authors gratefully acknowledge the support of
the European Commission in the projects WHEYPOL
(GRD2-2000-30385) and ANIMPOL (Contract No:
245084). Furthermore, the authors are thankful for the
STEM picture of C. necator provided by Dr. Elisabeth
Ingoli}, FELMI-ZFE Graz.
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M. KOLLER et al.: POLYHYDROXYALKANOATES: BIODEGRADABLE POLYMERS AND PLASTICS ...
12 Materiali in tehnologije / Materials and technology 47 (2013) 1, 5–12
V. KEVORKIJAN: CHALLENGES AND ADVANTAGES OF RECYCLING WROUGHT ALUMINIUM ALLOYS ...
CHALLENGES AND ADVANTAGES OF RECYCLING
WROUGHT ALUMINIUM ALLOYS FROM LOWER
GRADES OF METALLURGICALLY CLEAN SCRAP
RECIKLIRANJE GNETNIH ALUMINIJEVIH ZLITIN IZ
NIZKOCENOVNIH VRST METALUR[KO ^ISTEGA ODPADNEGA
ALUMINIJA
Varu`an Kevorkijan
Impol R&R, d. o. o., Partizanska 38, 2310 Slovenska Bistrica, Slovenia
varuzan.kevorkijan@impol.si
Prejem rokopisa – received: 2012-09-05; sprejem za objavo – accepted for publication: 2012-09-25
In the recycling of wrought aluminium alloys from lower grades of scrap (metallurgically clean but highly contaminated with
non-metallic impurities) the following two tasks were identified as the most demanding: (i) achieving the required final
chemical composition of an alloy with a minimal addition of primary aluminium and alloying elements; and (ii) keeping the
level of impurities (inclusions, hydrogen, trace elements and alkali metals) in the molten metal below the critical level. Because
of the lack of chemically based refining processes for reducing the concentration of alloying and trace elements in the molten
aluminium, once the concentrations of these constituents in the melt exceed the corresponding concentration limits, the only
practical solution for their reduction would be an appropriate dilution with primary metal. To avoid such a costly correction,
carefully predicting and ensuring the chemical composition of the batch in the pre-melting stage of casting should be applied.
Fortunately, some of the impurities, like hydrogen and alkali metals, as well as various (mostly exogeneous) inclusions, could be
successfully reduced by employing existing refining procedures.
In this work, (i) the state-of-the-art technologies, including some emerging technical topics such as the evolution of wrought
alloys toward scrap-intensive compositions, monitoring of the content of organics in the incoming scrap and the quality of
molten metal achieved by different smelting and refining technologies, and (ii) the relevant economic advantages of the
recycling of wrought aluminium alloys from the lower grades of scrap are reported. By analyzing the market prices of various
grades of scrap and the total cost of their recycling, the cost of aluminium ingots made from recycled aluminium was modelled
as a function of aluminium and the alloying-element content in the incoming scrap. Furthermore, scrap mixtures for producing
aluminium wrought alloys of standard quality from lower grades of scrap and with a significant new added value were
illustrated.
Keywords: wrought aluminium alloys, recycling, low grades of aluminium scrap, quality of recycled metal, economic benefits
Pridobivanje recikliranega aluminija standardne kakovosti iz nizkocenovnih virov bo v prihodnje odlo~ilno vplivalo na
konkuren~nost in uspe{nost evropske aluminijske industrije. Pri recikliranju gnetnih aluminijevih zlitin iz nizkocenovnih vrst
odpadnega aluminija (metalur{ko ~istih, vendar onesna`enih z nekovinskimi ne~isto~ami) sta posebej zahtevni naslednji dve
tehnolo{ki nalogi: (i) zagotavljanje `elene kemijske sestave aluminijeve zlitine ob minimalnem dodatku primarnega aluminija in
legirnih elementov in (ii) ohranjanje nivoja ne~isto~ (vklju~kov, vodika, elementov v sledovih in alkalijskih kovin) v talini v
mejah dovoljenega. Ker koncentracijo legirnih elementov in elementov v sledovih v talini tehnolo{ko ni mogo~e spreminjati s
kemijskimi postopki rafinacije, je, br` ko njihova koncentracija prese`e dovoljeno mejo, edina re{itev red~enje z dodatkom
primarnega aluminija. Tovrstnemu dragemu na~inu zagotavljanja predpisane kemijske sestave taline se lahko izognemo le z
doseganjem `elene kemijske sestave pred taljenjem, tj. na stopnji na~rtovanja vhodne zmesi. Preostale ne~isto~e, kot so npr.
vodik in alkalijske kovine ter nekateri vklju~ki (predvsem primarni), lahko uspe{no obvladujemo `e s sedanjimi postopki
rafinacije taline, ki jih tu opisujemo.
V tem delu tudi opisujemo (i) sodobne tehnolo{ke postopke na~rtovanja recikliranju prijaznih sestav gnetnih aluminijevih zlitin
s pove~anim dele`em odpadnega aluminija, spremljanja koncentracije organskih ne~isto~ v vhodnem odpadnem aluminiju ter
dolo~anja kakovosti taline, pridobljene z razli~nimi postopki taljenja in rafinacije, in ugotavljamo (ii) ekonomske prednosti
proizvodnje gnetnih aluminijevih zlitin z recikliranjem metalur{ko ~istega odpadnega aluminija ni`jega cenovnega razreda.
Izhajajo~ iz sestave in cen razli~nih vrst odpadnega aluminija ter stro{kov njihovega recikliranja nam je uspelo razviti
funkcionalni model, ki dolo~a ekonomi~nost proizvodnje ingotov iz recikliranega aluminija na osnovi vsebnosti aluminija in
legirnih elementov v vhodnem materialu. Model smo aplicirali na razli~ne vrste nizkocenovnega odpadnega aluminija in
pokazali, da proizvodnja gnetnih zlitin standardne kakovosti iz tovrstnih virov zagotavlja najvi{jo mogo~o dodano vrednost.
Klju~ne besede: gnetne aluminijeve zlitine, recikliranje, nizkocenovne vrste odpadnega aluminija, kakovost recikliranih zlitin,
ekonomske prednosti
1 INTRODUCTION
In recent years recycling of low-grade scrap has
become an increasingly important issue of the metal
supply for both casting and wrought alloys. Looking to
the future, the production of recycled aluminium of
standard quality from the cheapest sources will play an
increasingly significant role in the growth of the
European aluminium industry. Despite the economic
slowdown, the consumption of primary aluminium in the
EU is expected to increase (to 8 Mt by 2012), while the
European production of primary aluminium is expected
to decrease (gradually down to 2.86 Mt by 2012). The
gap between the expected production of primary
aluminium and its consumption (of about 5.24 Mt) will
be covered by the imports and recycling inside the EU.
Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23 13
UDK 669.715:620.28 ISSN 1580-2949
Review article/Pregledni ~lanek MTAEC9, 47(1)13(2013)
At the same time, the continuous increase in the relative
proportion of recycled- vs. primary-aluminium sources
will be driven by the pressure to improve business results
and striving for individual profit maximisation. Addi-
tional very important benefits of recycling aluminium
from low-grade scrap are: (i) spreading the risk of a
potential shortage of raw materials by diversifying the
supply sources of aluminium away from exclusively
primary metal and clean scrap suppliers; and (ii) an
improvement in the logistics – ensuring an appropriate
and cost-effective supply from different scrap sources.
Other advantages of recycling low grades of aluminium
scrap are the additional energy savings and a higher
compositional flexibility in combination with clean
grades of scrap and dross.
For several decades, a kind of belief existed in the
aluminium industry that the standard quality of wrought
alloys could be achieved only by combining sufficient
amounts of primary aluminium, internal scrap and only
clean, well-sorted external scrap. Consequently, ingots
made from primary aluminium, internal scrap and clean
industrial, or external, old scrap (a single wrought alloy
with the mass fraction less than 2 % of non-metallic
impurities) were obligatory in the mass production of
wrought alloys as the only source of the new and
recycled aluminium capable of assuring the standard
quality of end products.
Scrap for the production of wrought alloys should be
sorted with a strict control of the concentration of
alloying elements in order to achieve the prescribed com-
positional tolerances1–12. An additional problem is caused
by a very limited ability of wrought alloys to tolerate the
elements not normally present in their composition. In
other words, well-defined wrought scrap of a proper
composition can be effectively remelted into a wrought
alloy of the same composition, but it is very demanding
to achieve a new wrought composition with direct reuse,
without an addition of the primary metal and alloying
elements. An addition of primary aluminium is necessary
to dilute impurities (the elements not normally present in
a wrought alloy) to an acceptable level, while the alloy-
ing elements are added, if necessary, for the correction of
their concentration. Thus, most of the external scrap
inside the EU (above 60 %) is preferably applied in the
production of casting alloys and only the remainder is
dedicated to remelting.
Although non-metallic impurities can also signifi-
cantly influence the quality of the molten metal, it is not
obligatory that the scrap for wrought alloys should be
clean, without any organic and other non-metallic impu-
rities, if these could be effectively removed before or
during the recycling procedure. Some of the advanced
melting furnaces, such as various rotary or multi-cham-
ber units, allow direct melting of highly contaminated
scrap (e.g., painted and lacquered scrap) with the thermal
de-coating and the consequent recycling.
Moreover, in the internal technical documentation for
the production of wrought alloys, more or less empirical
compositions were often established, resulting in the
production mixtures with the prescribed amounts of
primary aluminium, internal, industrial external and old
external scrap. However, it is important to note that such
empirical compositions are usually adapted to the
common availability of various raw materials and in
many cases these are below the real potential of a
possible replacement of primary aluminium with scrap
without influencing the standard quality of the final
products. The problem is that in many cases these empi-
rical compositions are also approved by the customers,
becoming, in that way, a contractual obligation of a
producer of the alloys.
On the other hand, it is well known that for
aluminium alloys (especially wrought alloys) a practical
"compositional-tolerance limit" exists and a fairly
complete knowledge of these tolerance limits for all the
elements is needed, especially in the recycling operations
where unexpected and unusual impurities can creep in
inadvertently, and even normal impurities may tend to
accumulate and build up to a disastrous degree. In most
cases, the influence of these tolerance limits for various
elements and various combinations of the elements on
the properties (and particularly on the selected proper-
ties) of wrought alloys is not well investigated. Because
of that, customers often require more narrow composi-
tional tolerances than necessary, creating unnecessary
losses for themselves and the casting house. Customers
lose an important part of the competitiveness of their
products in downstream business activities by paying
more for non-optimal tolerance limits and, at the same
time, a casting house loses the added value by producing
alloys from more expensive inputs.
As an example, the average new added value created
by producing wrought alloys from external contaminated
scrap is about 7 % of LME. For the aluminium dross in
the form of pressed skulls the new added value is
significantly higher and can reach approximately one
third of LME (considering pressed skulls as internal
scrap).
2 DIFFICULTIES IN RECYCLING THE
EXISTING WROUGHT ALUMINIUM ALLOYS
As already mentioned, the main difficulty in the
production of wrought aluminium alloys from scrap is to
achieve the proper chemical composition of a melt with a
minimum addition of primary aluminium and alloying
elements. Technically, the problem is in the missing
technology (an economically acceptable, chemically
based refining process) for reducing the concentration of
the critical allying elements, such as copper, iron, man-
ganese, silicon and zinc, in a melt batch produced from
various sorts of scrap. Once the concentration of these
critical elements in a melt is above the concentration
V. KEVORKIJAN: CHALLENGES AND ADVANTAGES OF RECYCLING WROUGHT ALUMINIUM ALLOYS ...
14 Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23
limit for a particular wrought alloy, the only practical
solution would be their dilution by primary metal.
Another technical solution is to avoid an incorrect melt
composition by carefully predicting and assuring the
chemical composition of the batch in a pre-melting stage
of casting. In principle, there is also the third solution: to
convince the customers to accept the so-called "recycl-
ing-friendly wrought alloys" – in other words, the alloys
with broad compositional-tolerance limits and, conse-
quently, to some extent, a different quality-to-cost ratio.
This could be an important future trend in developing
new wrought alloys, working hand in hand with custo-
mers in the implementation of their requirements for
scrap-friendly compositions1–4, but for the existing alloys
and existing customer demands such an approach has
definitely quite a limited potential and it is also too risky.
In most of the current plants, the predominant mode
of recycling – a more accurate scrap blending or a very
strict melt dilution – is decided on the basis of the
margin between recycled metal and primary aluminium.
However, it is important to note that this margin – the
difference between the price of primary aluminium
(which is determined globally) and the price of recycled
metal (which is calculated locally) – is affected by
internal and external circumstances. Among internal
factors the most important are: (i) permanent and stable
sources of new and old scrap, concentrated sufficiently
in one area to justify the cost of collecting; (ii) a
scrap-collecting and sorting infrastructure including
devices for removing impurities and delivery to a
recycling plan; (iii) a method of recycling that is
economically competitive with the production of primary
aluminium and (iv), a market willing to accept the
composition and the quality of wrought alloys made
from scrap.
More expensive, clean and sorted scrap (mainly new
or industrial scrap) contains a minimum concentration of
critical elements, while in the old scrap of a lower cost it
becomes more critical.13 In typical municipal old scrap,
which is a cost-effective source of aluminium, the
minimal concentration of critical elements (silicon, iron,
copper, manganese, zinc, magnesium) is typically too
high for direct remelting into wrought compositions
without a dilution by primary aluminium (Table 1).
On the other hand, in the new scrap resulting from
collecting and/or treating the metal that forms during the
production of aluminium products before these are sold
to the final users, the right alloy composition is assured
in advance; however, the cost of such scrap is signifi-
cantly higher and its availability is usually limited to
closed production loops.
3 ECONOMICS OF RECYCLING CONTAMI-
NATED WROUGHT ALUMINIUM SCRAP
The economics of recycling wrought aluminium
alloys from scrap is specific and differs from the
economics of recycling cast alloys, since cast alloys have
higher compositional-tolerance limits for impurities and
can absorb a wider variety of scrap. During the produc-
tion of cast alloys from scraps of various compositions,
refiners are able to add alloying elements and remove
certain unwanted elements after the melting process.
Cast alloys tend to have higher alloy content than
wrought alloys and because of that they are difficult to
recycle into anything other than cast alloys, since the
removal of most alloying elements from the molten
aluminium would be impractical. On the other hand,
wrought scrap cannot be used for the production of new
wrought alloys unless separated by alloys or alloy groups
and/or diluted with an addition of primary metal.
In this regard, a possible way of improving the
recycling of wrought aluminium alloys is the use of new
and old scrap with higher amounts of organic impurities.
Most aluminium scrap mixtures currently used for the
production of wrought aluminium alloys from low-grade
scrap have an organic-impurity content in the mass
fraction lower than 8 %. The most common organic
impurities are oils, polymers such as polyester and
epoxy, rubber, lacquers, paints, etc. In some heavily
contaminated aluminium scrap, the organic-impurity
level exceeds 18–20 %, while in the clean industrial
scrap the non-aluminium impurity level is usually less
than 2 %. The main reasons to start using contaminated
instead of clean scrap for the production of wrought
aluminium alloys are the improved added value (a
net-profit surplus achieved per weight unit of aluminium
or aluminium alloys of the standard quality recycled
from the low-quality aluminium scrap) and better
logistics (scrap sourcing, availability on the market and
improved supply flexibility). The key advantage is the
fact that the cost of contaminated scrap is significantly
lower than the cost of clean scrap of the same pre-sorting
quality (e.g., a single alloy or a single-series grade), in
that way providing an opportunity for producing alumi-
nium wrought alloys of the standard quality and with
V. KEVORKIJAN: CHALLENGES AND ADVANTAGES OF RECYCLING WROUGHT ALUMINIUM ALLOYS ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23 15
Table 1: Typical concentrations of the main alloying elements in the
municipal old scrap13
Tabela 1: Obi~ajne koncentracije najpogostej{ih legirnih elementov v
komunalnem odpadnem aluminiju13
Element Concentration (%)
Fe 0.60–1.00
Si 0.30–9.00
Cu 0.25–4.00
Mn 0.60–1.50
Zn 0.25–3.00
Mg 0.20–2.00
Cr 0.05–0.30
Ni 0.04–0.30
Pb 0.02–0.25
Sn 0.02–0.30
Bi 0.02–0.30
Ti 0.05–0.25
improved competitiveness. The prerequisite for this lies
in an appropriate performance of the entire recycling
process, from the scrap sourcing and purchasing strategy
to the complete recovery of all the by-products, in order
to achieve the standard quality of the recycled aluminium
and the proper economy. In practice, irrespective of the
fact that the recycling of contaminated scrap is more
demanding and costly than the remelting of clean scrap,
the total cost per weight unit of recycled aluminium or
aluminium alloy of the standard quality produced from
contaminated scrap is lower than the cost of remelted
aluminium or an Al-alloy counterpart produced from
clean scrap that is nowadays much in demand for the
production of wrought alloys, Figures 1 and 2.
As a rule, clean scrap (e.g., the scrap with a mini-
mum of 98 % of Al) represents a costly raw material for
the production of wrought alloys. Its market price is
close to the one theoretically expected, calculated
according to the aluminium content and the cost of
recovery. Therefore, the usage of clean scrap in the pro-
duction of wrought alloys provides only limited possibi-
lities for creating a new added value or, in other words,
for lowering of the cost of the input. Typically, the
market prices for clean scrap of a single wrought alloy
vary slightly below or above the price of the counterpart
ingots, depending on their market availability14,15 (Table
2).
On the contrary, the market price of contaminated
scrap (the scrap with, e.g., 80 % of Al and 20 % of non-
metallic, mostly organic impurities) ranges significantly
(10–25 %) below the theoretically expected price based
on the aluminium content and the cost of recovery14,15
(Table 2).
Thus, taking into account the cost of recycling and all
the related costs, the total cost of producing recycled
aluminium alloy fabricated from clean scrap is usually
close to the cost of melting the same alloy from primary
aluminium (including the cost of the appropriate alloying
elements). In contrast to that, by using less-clean scrap
(the scrap contaminated with organic impurities) and
applying the proper recycling technology, a higher net
V. KEVORKIJAN: CHALLENGES AND ADVANTAGES OF RECYCLING WROUGHT ALUMINIUM ALLOYS ...
16 Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23
Figure 2: Cost of aluminium ingots made from recycled aluminium as
a function of aluminium and alloying-element content in the incoming
scrap
Slika 2: Spreminjanje cene aluminijevih ingotov, izdelanih iz recikli-
ranega aluminija, v odvisnosti od vsebnosti aluminija in legirnih
elementov v vhodnem odpadnem aluminiju
Figure 1: Structure of the cost of Al ingots made from recycled
aluminium
Slika 1: Struktura cene Al-ingotov, izdelanih iz recikliranega alumi-
nija
Table 2: Composition and average price* of selected scrap types14,15
Tabela 2: Struktura in povpre~na cena* izbranih vrst odpadnega aluminija14,15
Scrap description Aluminiumcontent (%)
Oxides
(%)
Foreign material
(%)
Average price
(% LME)
One single wrought alloy 97.2 1.0 1.8 95–99
Two or more wrought alloys of the same series 97.2 0.8 2.0 80–85
Used beverage cans 94 0.8 5.2 55–60
End of profiles with a thermal bridge (one single
wrought alloy) 78 3.8 18.2 55–70
Turnings, one single alloy 95.3 3.7 1.0 80–85
Mixed turnings, two or more alloys 84.0 3.3 12.8 75–80
Packaging (coated) 71.5 3.8 24.7 55–65
Packaging (de-coated) 86.1 12.9 1.0 92–95
Dross (one single wrought alloy) 55.7 44.3 - 15–45
*The reported values are only indicative
*Predstavljene vrednosti cen so le informativnega zna~aja
added value (typically between 5 % and 10 %) can be
achieved.
It is important to note that a significant part of the
new added value is gained by successful buying of less
clean grades of scrap. Hence, it is necessary to
understand the local new and old scrap markets and
organize cost-effective buying from the nearest scrap
suppliers or though collecting new scrap.
The second part of the new added value is achieved
in the process of scrap separation, where the optimum
level and method (e.g., hand sorting or automatic screen-
ing) of separation should be selected following the
compromise between the degree of compositional
separation and the cost of achieving it, also taking into
consideration that a lower level of compositional
separation leads, during the final melting, to a higher
consumption of primary aluminium for diluting impuri-
ties. Of all the sorting technologies, hand sorting remains
the most common method of recovering aluminium.
Because a load of mixed scrap (even new, industrial one)
often includes a limited number of alloys, hand sorting
often makes it possible to produce single-alloy scrap
products. Even if this is not possible, hand sorting can
help meet specifications for the other scrap grades by
removing impurities. A dealer’s experience and know-
ledge of his suppliers is often useful in hand sorting,
because the appearance of a piece of new scrap – the
shape of punching, the type of a scrapped part – will be
sufficient to identify the alloy.
Finally, the remaining part of the new added value
depends on the competitiveness of the selected remelting
technology, which should be able to provide the highest
metal yield, the standard quality of the molten metal and
an operation in accordance with the standard environ-
mental regulations.
4 MELTING TECHNOLOGIES FOR
CONTAMINATED SCRAP
Generally, regarding the level of organic impurities,
two scrap-melting approaches are practiced: with or
without melting additives16. Without melting additives it
is possible to melt clean scrap, preferably containing less
than 2–3 % of organic impurities and contaminated scrap
(with less than about 10 % of organic impurities) by
applying twin or multi-chamber melting furnaces. The
scrap with a higher amount of organic impurities should
be melted with an addition of melting additives (usually
a NaCl and KCl salt mixture) in a drum rotary furnace
with a fixed axis, which is the universal furnace for
melting all kinds of highly contaminated scrap, including
aluminium dross and pressed skulls. However, a more
advanced and economic way of recycling aluminium
from dross and pressed skulls is with a tilting rotary
furnace, in which recycling can be performed with a
significantly lower amount of added salts. In addition,
tilting rotary furnaces are often used in recycling cast
alloys, while only a limited numbers of such devices
have been installed until now for the recycling of
wrought alloys. It is important to note that salts provide
the best quality of a molten metal. The salt mixture
covers the aluminium to prevent further oxidation, strips
away the oxide layer from the molten metal, promotes
the coalescence of metallic droplets and dissolves or
suspends other impurities attached to the metal.
Therefore, the use of salt is imperative for achieving the
maximum quality of recycled aluminium, especially
when highly contaminated scrap and scrap with a large
specific surface area are melted.
However, salts are costly additives in the production
and result in a significant amount of the salt cake
by-product, whose processing introduces an extra cost of
salt recovery and the deposition of the non-metallic
residue on commercial or industrial landfills. The
melting strategy for recycling wrought alloys from scrap
contaminated with organic impurities depends on several
factors, among which the maximum level of organic
impurities in the batches prepared for melting is one of
the most important. For melting contaminated scrap salt
free, various possibilities exist. The most advanced and
integrated device for direct melting of contaminated
scrap without melting additives is the multi-chamber
furnace with a tower. The alternative is to melt conta-
minated scrap in a twin-chamber furnace. However, in
this case the organic impurities should be reduced in
advance to some acceptable level. Salt-free remelting
devices (e.g., three-chamber melting furnace, twin-cham-
ber furnace with a tower, etc.) are suitable for contami-
nated scrap having less than 10 % of the total organic
impurities. This could be achieved mechanically by
shredding or by thermal de-coating. The practical
alternative is lowering the amount of organic impurities
by mixing the contaminated scrap with a sufficient
amount of clean scrap.
In any case, the melting technology chosen will
determine the allowed level of organic impurities in the
scrap, as well as the eventual necessity for melting
additives. Currently, there is no single universal melting
device flexible enough for all the grades of scrap
(regarding the content of organic and non-metallic
impurities, as well as the scrap specific surface area),
operating without melting additives. For example, the
double-pass rotary drum furnace is the only furnace that
is suitable for all kinds of scrap. However, it operates
with the highest salt factor. On the other hand, salt-free
devices are limited by the amount of organic impurities,
which can also reduce productivity. The same problem
exists in rotary furnaces, where the highest productivity
is achieved with a well controlled amount of organic
impurities.
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Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23 17
5 EVOLUTION OF WROUGHT ALLOYS
TOWARD SCRAP INTENSIVE COMPOSITIONS
First of all, it is important to note that customers do
not buy a wrought alloy composition but wrought pro-
perties. This fact, which is crucial in a negotiation about
the optimum wrought-alloy composition, should be well
recognized by both parties involved in an order nego-
tiation – not only by customers, but also by producers of
wrought alloys, and vice-versa.
Unfortunately, in the existing, standardized wrought
aluminium alloys the tolerance limits for all the
constituents of the alloys were well defined before scrap
recycling became the key issue in the added-value
engineering along the aluminium production chain.
Thus, when ordering these traditional alloys, customers
are more or less obliged to request the standard compo-
sition and properties.
The common limitation of the existing wrought
alloys is that they are not compositionally tolerant
enough to be produced by direct mixing and melting of
scrap batches without sorting the mixed scraps to the
desired level. Therefore, traditional wrought aluminium
alloys offer only limited opportunities for a direct reuse
of the recycled wrought alloy scrap without tight compo-
sitional corrections (the so-called "sweetening") by
primary metal and alloying elements.
In the current wrought alloys the real operational
dilemma of how well to sort17 depends on the extent of
primary aluminium, which is to be substituted by a
recycled grade without influencing the quality of the
wrought alloy. It is absolutely clear that the amount of
primary aluminium, which can be effectively replaced in
a particular wrought alloy by a recycled metal, depends
on the level of compositional separation of the scrap. In
other words, more precisely compositionally separated
scrap has a higher potential for replacing primary alumi-
nium without affecting the quality of the final alloy.
Theoretically, by repurposing the completely sorted
scrap (one single wrought alloy), a zero consumption of
primary aluminium can be achieved in the production of
wrought alloys. However, in order to avoid the economic
inefficiencies occurring when such high-value scrap is
repurposed into compositionally more tolerant wrought
alloys, it is always necessary, in practice, to measure the
net economic benefits of such a replacement, taking into
consideration the cost of separation and the market value
of the selected wrought alloy.
On the other hand, compositionally less-separated
scrap grades with the alloys inside the same series, two
wrought alloys of different series or even a mixture of
various wrought alloys, will require, during the melting,
an additional consumption of primary aluminium for
diluting the impurities influencing the final economic
benefit of such a substitution.
In any case, it is important to note that the "de facto"
role of primary aluminium is the dilution of the impurity
level (not sufficiently reduced through compositional
separation of scrap) and not the provision of a sometimes
mystic, necessary amount of "virgin metal", which is,
according to some opinions, obligatory for achieving the
standard quality of wrought alloys.
Finally, the question of economy arises again. By
applying the state-of-the-art scrap-separation technology,
from a technical point of view, it is possible to achieve
compositionally well separated grades of scrap suitable
for direct melting to the appropriate wrought alloys. The
problem is that this is still not economically reasonable
due to the high cost of scrap separation to a level of
impurities acceptable for the existing wrought alloys.
For that very reason, the technique of creating a new
added value through scrap recycling should lead toward
a formulation of new, recycling-friendly wrought alloys
if, finally, this would be acceptable for the end product
customers.
There are several fundamental questions to be
answered concerning future developments of new
wrought alloys designed to provide wider compositional
tolerances of the existing or other alloying elements and,
hence, better opportunities for scrap consumption. The
most important one is whether these new alloys can
possibly be composed without a critical loss of appli-
cation properties or, in other words, still provide the
valuable and desired combination of wrought properties
for customers. Significant efforts, scientifically, techno-
logically and financially, will be necessary for achieving
this goal and implementing it in the industrial usage.
Another important question concerns a possible long-
term reduction in the number of wrought aluminium
alloys by establishing a limited number of universal
wrought compositions, making the refining of alloys
easier. Although a unification of wrought compositions
was proposed several times in the past, the actual deve-
lopment is progressing toward a further diversity of the
alloys and highly tuned properties.
Irrespective of whether a new generation of
recycling-friendly wrought alloys will be developed or
the existing ones unified, it is important to note that
newly tailored wrought alloys will require a fulfilment of
the following two, hardly compatible, demands: (i) com-
positions with relatively broad specification limits on the
major alloying elements and more tolerant limits on
impurities; (ii) no significant restrictions on performance
characteristics for final applications. A complete deve-
lopment and implementation of such alloys is, obviously,
not an easy metallurgical task and will remain, most
probably, the challenge for future decades.
6 DETERMINING THE CONTENT OF
ORGANICS IN ALUMINIUM SCRAP
The suitability of aluminium scrap with organics as
an appropriate source of aluminium in terms of value and
quality for the production of wrought aluminium alloys
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18 Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23
of the standard quality depends on its metallurgical
composition and content of organic and other impurities
(humidity, non-metals such as oxides, non-oxides, etc.)1.
The metallurgical composition of incoming scrap,
influenced by the mix of the included alloys, is routinely
determined in recycling plants by standard optical
emission spectroscopy. However, a similarly fast and
cost-effective method of analysing the amount of
organics and other impurities in the representative
samples of incoming scrap is still under development2,3.
In recycling plants specialising in recycling of low
grades of aluminium scrap, timely and accurate infor-
mation about the amount of organics and other
impurities in the incoming scrap is an absolute pre-
requisite in achieving a higher net added value. The goal
is, considering the amount of organics and other impu-
rities, to recycle each grade of scrap with the optimum
recycling procedure from both the economic and the
metallurgical aspects.
From an economic point of view, the content of
organics and other impurities (of non-aluminium phases)
defines the cost of scrap and it is thus absolutely
indispensable information in commercial issues related
to scrap buying, as well as the added-value engineering
of the entire process of recycling1.
From the metallurgical aspect, the amount of orga-
nics and other impurities determines the key techno-
logical parameters of recycling and the process of
refining the molten metal for achieving the standard
quality of end products. Firstly, the amount of organics
determines the exothermic/endothermic behaviour of the
incoming scrap and the volume of gaseous products
liberated during the early stages of recycling. In addition,
the concentration of organics and other impurities
favours the formation of inclusions, influencing the
quality of the molten metal4.
Aluminium scrap, especially the grades with orga-
nics, represents a highly exothermic input (e.g., scrap
with approximately 6 % of organic impurities is exo-
thermic enough for self-melting). During the melting of
such an input, a huge amount of energy and gaseous
products are generated simultaneously, influencing the
productivity, the cost of recycling, as well as the quality
of the molten aluminium obtained – in particular its
suitability for the production of wrought aluminium
alloys for highly demanding end products.
Inclusions appearing as various solid particles in
molten aluminium or aluminium alloy can be classified
into two main groups: (i) indigenous or in-situ inclu-
sions, and (ii) exogenous inclusions. The growth of
indigenous inclusions is caused by chemical reactions
taking place in the melt due to the existing chemical
composition and the applied processing parameters
(temperature, time, atmosphere, etc.). In contrast with
that, exogenous inclusions already exist as a separate
phase in the system before the melting and are intro-
duced to the melt by raw materials, alloying elements,
additives, refractory materials and the furnace atmo-
sphere4.
Exogenous inclusions can be effectively removed by
filtering. On the other hand, the presence of indigenous
inclusions may be prevented effectively only by master-
ing the overall reactivity in the system – by selecting and
maintaining the proper chemical composition of the melt
and the processing parameters.
Regarding the overall reactivity in a system, it is
particularly important to note that the level of impurities,
and the resultant chemical reactivity in remelted alumi-
nium scrap with organics (painted scrap), is quite
different from that in primary metal or remelted clean
scrap. The organic component of paint can be volatilized
during decoating and pyrolysis, but paint often contains
inorganic compounds (fillers and pigments) that do not
respond to thermal processing or become converted to
oxides or other (usually binary) compounds, remaining
on the surface of pyrolyzed scrap as exogenous inclu-
sions. Many of these exogenous inclusions may also
react with molten aluminium and/or alloying elements
dissolved in the melt, creating indigenous as well as new
exogenous inclusions.
Note that fillers and pigments can react either with
molten aluminium or/and in parallel with some of the
alloying elements dissolved in the melt. Hence, in real
scrap mixtures of a very complex chemical impurity
composition originating from organics, it is very import-
ant to take into account all the thermodynamically
possible chemical reactions enabling the formation of
inclusions. The aim of such a consideration is a careful
prescription of the tolerance limits for all the recycled
alloy constituents (incoming scrap), including impurities
originating from organics.
Due to an enormous growth of the secondary
aluminium industry, the development of an industrial
method of determining the amount of organics and other
impurities in incoming scrap has become a highly
important issue. This is particularly the case in recycling
plants and casting houses where low grades of scrap are
also used for the production of wrought aluminium
alloys. Thus, the purpose of this paper is to present an
industrial method of determining the amount of organics
and other impurities in representative samples of
incoming scrap.
6.1 Batch and continuous procedures of a TG analysis
Industrial TG/DTA analyses can be run either as a
batch or a continuous process, Figures 3 and 4.
The batch TG/DTA device is schematically presented
in Figure 5.
A significantly higher productivity of a scrap analysis
can be achieved with the TG/DTA unit presented in
Figure 6, which operates continuously.
The accuracy of measuring the humidity and the
organics content in the batch method (in an inert
V. KEVORKIJAN: CHALLENGES AND ADVANTAGES OF RECYCLING WROUGHT ALUMINIUM ALLOYS ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23 19
atmosphere) depends on the sensitivity of the balance
employed. In our case this was ±50 g, corresponding to
the maximum relative error of about 1 %. The lowest
relative error was obtained in the samples with minimum
weight losses (the minimum content of organics). For
example, the relative error of measuring the organic
content in a sample with an initial mass of 50 kg and
V. KEVORKIJAN: CHALLENGES AND ADVANTAGES OF RECYCLING WROUGHT ALUMINIUM ALLOYS ...
20 Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23
Figure 6: Industrial TG/DTA device capable of operating in both
continuous and batch modes
Slika 6: Shema industrijske opreme za TG/DTA-analizo odpadnega
aluminija, ki omogo~a merjenje po {ar`nem ali zveznem postopku
Table 3: Comparison of experimental results of the TG/DTA analysis
for aluminium scrap obtained in the batch and continuous modes
Tabela 3: Primerjava eksperimentalnih rezultatov TG/DTA-analize
odpadnega aluminija, pridobljenih pri {ar`nem in zveznem na~inu
merjenja
Scrap lot 1 Sample A,batch mode
Sample B,
continuous
mode
Humidity (%) 0.7 ± 0.02 0.8 ± 0.02
Content of organics (%) 15.6 ± 0.4 14.1 ± 0.4
Content of non-organics
including aluminium (%) 84.6 ± 2 85.8 ± 2
Recycled aluminium* (%) 73.6 ± 2 74.7 ± 2
* The estimated recycling efficiency was 0.88
Figure 5: Industrial TG/DTA device operating in the batch mode
Slika 5: Shema industrijske opreme za {ar`ni postopek TG/DTA-ana-
lize
Figure 7: TG/DTA curve obtained in the batch-process mode
Slika 7: TG/DTA-krivulja pri {ar`nem postopku merjenja
Figure 4: Histogram of mass changes of a representative scrap sample
during TG/DTA performed in the continuous mode
Slika 4: Histogram spreminjanja mase reprezentativnega vzorca
odpadnega aluminija med zvezno TG/DTA-analizo
Figure 3: Histogram of mass changes of a representative scrap sample
during TG/DTA performed in the batch mode
Slika 3: Histogram spreminjanja mase reprezentativnega vzorca
odpadnega aluminija pri {ar`ni TG/DTA-analizi
approximately 10 kg of organics was about 0.25 %,
while in a sample of the same initial mass but having
80 % of organics, the relative error of measurement was
1 %. However, in both cases the relative error of measu-
rement is within the demands of industrial users.
The accuracy of measuring humidity and organics
content with the continuous method (in an oxidizing
atmosphere of argon with 1 % of oxygen) in the same
way depends on the sensitivity of the balance applied
and, in addition, on the time and temperature of the
thermal degradation of organics, and is usually lower
than in the batch mode. The most important prerequisite
in achieving the highest accuracy of measurement in the
continuous mode, similar to that achieved in the batch
mode, is an efficient minimization of oxidation of alumi-
nium during the thermal degradation of organics in the
scrap sample. Due to the fact that the molar mass of
aluminium oxide is almost twice as high as the molar
mass of the stoichiometrically equivalent amount of
aluminium, any oxidation of aluminium during the
TG/DTA measurement is detected as an increase in the
mass of the remaining scrap sample and, correspond-
ingly, interpreted as a lower content of organics.
Two basic approaches are practiced for minimizing
aluminium oxidation during the TG/DTA measurement
performed in the continuous mode: (i) prolonged expo-
sure of the scrap sample at a lower temperature (480 °C
to 520 °C), or rapid heating of the scrap at a temperature
just below the melting point (560 °C to 620 °C). The
right combination of the temperature and the time
strongly depends on the scrap morphology (a thick or
thin gauge), the kind of organics (soluble oil, mineral oil,
paint, plastic or lacquer) and the percentage of the
organic phase. On that account, for each particular lot of
incoming scrap, in which the organics content is to be
analysed by an industrial TG/DTA measurement made in
the continuous mode, the proper time and temperature of
a complete organics removal with a minimum aluminium
oxidation should be defined in advance. The best way of
selecting the proper temperature and time is based on the
results (humidity, content of organics, content of non-
organics including aluminium) obtained on a represen-
tative sample in the batch mode as the reference values.
Accordingly, the parameters of the continuous mode
(temperature and time) should be selected to reproduce
the results at the same accuracy level as determined in
the batch mode.
It is important to note that the industrial TG/DTA
device developed for working in the continuous mode
(Figure 5) can quite easily also operate in the batch
mode. This can be done by heating the sample in the
bottom chamber starting at room temperature in an
atmosphere of pure argon. The decisive advantage of
such an industrial TG/DTA device is in its ability to
operate in both modes. Thus, the optimisation of the
processing parameters (temperature and time) for
operating in the continuous mode with the same level of
accuracy as in the batch mode gained a reference
counterpart and become an end-user-friendly and routine
operation, easily applicable to a wide spectrum of
incoming aluminium scraps. Following this methodo-
logy, the batch measurement should be completed first,
irrespective of the scrap morphology, the kind of orga-
nics and the percentage of organic phase, thus providing
the complete reference values of a scrap analysis. After
that, in the second step, the main processing parameters
(temperature and time) of the continuous mode should be
tuned in order to assure, at the same accuracy level,
comparable results for a scrap analysis.
A comparison of the experimental results (Figures 7
and 8) obtained with the TG/DTA analysis performed in
the batch and continuous mode (Table 3) clearly
confirms that the parameters of the continuous mode
applied in this work (temperature: 560 °C, holding time:
60 s) were correctly selected, resulting in comparable
values of humidity, content of organics and content of
non-organics including aluminium.
7 QUALITY OF THE MOLTEN METAL
The quality of the molten metal is one of the critical
issues, particularly if low-grade scrap becomes the
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Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23 21
Figure 8: Example of a TG curve generated in the continuous mode
Slika 8: Primer TG-krivulje, pridobljene pri zveznem postopku
merjenja
Table 4: Common impurities in primary and recycled molten alumi-
nium18
Tabela 4: Najpogostej{e ne~isto~e v talini na osnovi primarnega in
sekundarnega aluminija18
Impurity
Concentration in
primary
aluminium melt
Concentration in recycled
aluminium melt
Hydrogen 0.1–0.3 ìg/g 0.4–0.6 ìg/g
Inclusions
(PoDFA scale)
>1 mm2/kg
(Al4C3)
0.5–5.0 mm2/kg
(Al2O3, MgO, MgAl2O4,
Al4C3, TiB2)
Sodium 30–150 ìg/g <10 ìg/g
Calcium 2–5 ìg/g 5–40 ìg/g
Lithium 0–20 ìg/g <1 ìg/g
dominant raw material for the production of wrought
alloys of the standard quality.
As already discussed, various scrap-melting techno-
logies influence the quality of the resulting metal
through the concentration of the most common impuri-
ties in the molten aluminium, such as hydrogen, reactive
metals and inclusions.
Throughout almost the whole of the 20th century, the
aluminium produced by remelting scrap was treated by
customers as less valuable than primary aluminium
produced by electrolysis, mostly due to the concerns
over the purity of the recycled metal compared to that of
primary aluminium18,19 (Table 4). However, the develop-
ment of the refining technology (in-line degassing and
filtration) and analytical methods for measuring the
impurity levels in the past 20 years eliminated this
stigma completely, providing the same quality of the
refined molten aluminium, irrespective of its fabrication
pre-history.
8 CHALLENGES FOR THE FUTURE
The most important reasons for the increasing
demands for lower-grade scrap consumption in
wrought-alloy production are in seeking individual profit
maximisation, a shortage of clean scrap and both a
shortage and the high price of primary aluminium.
The increased consumption of lower grades of scrap
(contaminated external scrap) in the production of
wrought aluminium alloys makes the achieving of the
standard quality of the end products more challenging.
Thus, scrap pre-sorting from alloy to alloy, or at least in
a series of alloys, proper mixing of various scraps to pro-
vide the required chemical composition of the raw
material before melting with a minimum consumption of
ingots and alloying elements, advanced melting techno-
logy for achieving a high yield and the required environ-
mental standards, as well as refining and filtration to
assure the standard quality of the alloy, are increasingly
necessary.
The development of new alloys with the required
properties (e.g., tensile properties, workability, heat
treatment, deformation, etc.) could be achievable with
more flexible compositional limits. It would be
necessary to develop such recycling-friendly wrought
compositions and demonstrate to customers the ability to
tailor end properties and the economic benefits created
by high contents of scrap.
The following advancements in technology will be
necessary to achieve the production of any wrought
alloys from scrap without ecological problems:
• Develop and design melting furnaces that minimize
the melt loss (oxidation and dross formation during
remelting) and consumption of melting additives,
improve cost effectiveness and productivity, increase
safety and reduce emissions.
• Develop a low-cost process for metal purification to
enable the production of primary alloys from
recycled scrap, including the methods to remove
specific impurities such as Mg, Fe, Pb, Li, Si, and Ti.
• Develop new, scrap-tolerant wrought alloys that
better match scrap to the specifications for an
increased utilization.
However, it is important to note that until now, there
have been no effective methods for fulfilling the above
requirements technically and economically. Most of the
investigations (e.g., metal purification) are still at the
stage of fundamental or early applied research, with their
progress being uncertain and not foreseeable. Hence, the
earliest eventual implementation at the industrial level
might be expected in the coming decades.
9 CONCLUSION
Because of high costs and shortages of raw materials
(primary aluminium and clean scrap), the main challenge
for the producers of wrought aluminium alloys and semis
are: (i) running the production with alternative, cost-
effective sources of aluminium; and (ii) with the sources
of metal that are more easily available.
According to the general estimation that between
3–10 % of LME is the average amount of the new added
value achieved by remelting contaminated scrap, the
consumption of low-grade scrap, which is already
frequently practised by the producers of cast alloys, is
also being increasingly introduced by remelters.
However, in contrast to mixed scrap for refiners,
scrap batches for remelting should be compositionally
well correlated with the chemical composition of the
wrought alloy to be produced (preferably consisting of
one alloy) and clean enough (not oxidized or conta-
minated with non-metallic impurities). Traditionally,
remelters were defined as the producers of wrought
alloys, mainly from clean and sorted wrought-alloy scrap
and also distinguished from refiners by a lack of refining
capability.
Recent developments in the remelting technology and
inside the global recycling industry – together with the
actual global economic crisis – started to change this
traditional framework toward a new mentality of
remelters. Following the opportunities for creating a new
added value in their niche business, remelters reconciled
the production of wrought alloys from less clean, the
so-called metallurgically clean scrap, which can be
contaminated even with high amounts of various
non-metallic (e.g., organic) impurities. In addition, they
became familiar with achieving the proper composition
of scrap batches before loading the scrap into the furnace
(through the refining of scrap), avoiding the more expen-
sive dilution of impurities by primary aluminium during
the melting. To this end, several pre-melting operations
(scrap sorting and separation, as well as the in-house-
scrap batch compositional blending) were integrated into
the production chain, together with some post-melting
operations, such as the traditional molten-metal refining.
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22 Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23
With all these changes contributing essentially to the
creation of a new added value, a new mentality of
remelters closer to the refining-production practice was
established inside the EU, increasing the importance of
contaminated scrap as a long-term source of aluminium
for wrought alloys. Actually, remelters well understood
that the most significant part of the new added value is
created through proper scrap buying and sorting, while
only the remainder is gained by advanced remelting.
Thus, a kind of "scrap refining" practice must be intro-
duced to keep different aluminium alloys separated at
some appropriate level, considering both the metallur-
gical and economic point of view. The key issue is to
achieve the right alloy composition of a scrap mixture
before melting and not at the end of melting by diluting
the impurity content to the required level. The only way
to achieve this is by being fully acquainted with the scrap
quality through an excellent knowledge of the scrap
market, the individual scrap suppliers and an internal
knowledge of scrap sampling.
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Parkway, NW 2007, 101
14 U. M. J. Boin, M. Bertram, JOM, 57 (2005) 8, 26
15 A&L, The aluminium figures, Edimet (www.edimet.com)
16 Ch. Schmitz, Handbook of Aluminium Recycling, Vulkan-Verlag,
Essen, Germany 2006, 74
17 P. Li, S. Guldberg, H. O. Riddervold, R. Kirchain, In EPD Congress
2005, Ed. M. E. Schlesinger, TMS, Warrendale 2005, 1159
18 M. E. Schlesinger, Aluminum Recycling, CRC Press, Broken Sound
Parkway, NW 2007, 171
19 M. E. Schlesinger, Aluminum Recycling, CRC Press, Broken Sound
Parkway, NW 2007, 172
V. KEVORKIJAN: CHALLENGES AND ADVANTAGES OF RECYCLING WROUGHT ALUMINIUM ALLOYS ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 13–23 23

M. EROL, E. ÇELIK: GRAPHITE-FLAKE CARBON-BLACK-REINFORCED POLYSTYRENE-MATRIX COMPOSITE ...
GRAPHITE-FLAKE CARBON-BLACK-REINFORCED
POLYSTYRENE-MATRIX COMPOSITE FILMS
DEPOSITED ON GLASS-FIBER WOVEN FABRICS AS
PLANE HEATERS
KOMPOZIT POLISTIRENA, OJA^AN Z GRAFITNIMI LUSKAMI IN
SAJAMI, NANESEN NA TKANINO IZ STEKLENIH VLAKEN ZA
PLO[^ATE GRELNIKE
Mustafa Erol1,2,3, Erdal Çelik1,2
1Dokuz Eylul University, Department of Metallurgical and Materials Engineering, Buca, 35160 Izmir, Turkey
2Dokuz Eylul University, Center for Production and Applications of Electronic Materials (EMUM), Buca, 35160 Izmir, Turkey
3Dokuz Eylul University, Graduate School of Natural and Applied Sciences, Buca, 35160 Izmir, Turkey
m.erol@deu.edu.tr
Prejem rokopisa – received: 2012-05-07; sprejem za objavo – accepted for publication: 2012-07-04
Graphite-flake carbon-black/polystyrene composite films as plane heaters are promising materials since they are smarter than
the traditional heating elements. In the present study, we are concerned mainly with the production, characterization and
industrial application of graphite-flake carbon-black-reinforced polystyrene-matrix composite films deposited on glass-fiber
woven fabrics as plane heaters. Within this scope, graphite flakes and carbon-black powders were dispersed in polystyrene gel
and deposited on glass-fiber woven fabrics with different weight ratios. Subsequently, the films were dried at 60 °C for 30 min
in the air. Structural and surface properties of the produced films were characterized with XRD and SEM, respectively.
Electrical and heating properties were determined with a hand-made experimental setup containing a multimeter and a
thermocouple. It was found that uniform and partially ordered films depending on the weight ratio and percolation threshold
were obtained as plane heaters. Planar heating of up to 60 °C was observed with a 24-V DC power supply.
Keywords: plane heater, composite, graphite flake, carbon black, conductive polymer
Kompozitne plasti polistirena z grafitnimi luskami – sajami so obetajo~i materiali za plo{~ate grelnike, ker so „pametnej{i” od
tradicionalnih grelnih elementov. V tej {tudiji obravnavamo proizvodnjo, karakterizacijo in industrijsko aplikacijo kompozitne
tanke plasti polistirena, oja~anega z grafitnimi luskami – sajami, nanesenega na tkanino iz steklenih vlaken, za plo{~ate grelnike.
S tem namenom so bile v tej {tudiji grafitne luske in saje razpr{ene v polistirenskem gelu, nanesene v razli~nem masnem
razmerju na tkanino iz steklenih vlaken. Nato so bile tanke plasti su{ene pri 60 °C 30 min na zraku. Strukturne in povr{inske
lastnosti nastalih tankih plasti so bile ocenjene z XRD in SEM. Elektri~ne in grelne lastnosti so bile dolo~ene z ro~no izdelano
sestavo, ki vsebuje multimeter in termoelement. Ugotovljeno je bilo, da je pri plo{~atih grelnikih enakomerna in delno urejena
plast odvisna od razmerja mas in dele`a pronicanja. Pri uporabi enosmerne napetosti 24 V je bilo opa`eno segrevanje plo{~atih
grelcev do 60 °C.
Klju~ne besede: plo{~ati grelnik, kompozit, grafitne luske, saje, prevoden polimer
1 INTRODUCTION
Polymers with their specific nature are known as
good insulators for electronic applications. Develop-
ments of the polymers, together with the new researches,
have focused on desired conductivities and a wide range
of application for decades. The conductivity of polymers
can be obtained in two ways: (a) by producing a polymer
that is intrinsically conductive or can be made so by
doping and (b) by loading an electrically insulating
matrix with conductive fillers1. The fillers, which involve
conductivity, are generally based on metallic materials
such as Ni, Cu, Ag, Al and Fe2–4 as well as carbon deri-
vates such as carbon black, carbon fiber, graphite and
carbon nanotube5–8.
Natural graphite is mostly used in refractories, steel-
making (as a heating element/electrode), expanded
graphite, brake linings, foundry facings and lubricants.
The heating property of graphite is a significant issue for
electric-arc furnaces in the steel industry. As heating has
been an important process for humankind for ages, the
above issue has also been present since the discovery of
fire up until today’s modern heating technologies. With
respect to heating, fires, stoves, heat exchangers, air
conditioners, furnaces, irons, floor heaters, etc. are the
products that fulfill this requirement. These heating
systems generally use fossil fuels (wood, petroleum,
natural gas, etc), as well as solar energy and electricity.
Due to the physical nature of heated air, it flows from
bottom to top in a heated place. Taking this into account,
efficient systems for heating, like floor heaters, can be
selected. There are many detailed researches on floor
heating in several reports9–12. In addition, due to their
conductive properties, graphite and carbon-black pow-
ders were incorporated into its structure to provide for
the resistivity of the final composite structure. Never-
theless, to the best of our knowledge, no experimental
work has been reported in the literature on the self-
heating properties of polymer-graphite and carbon-black
Materiali in tehnologije / Materials and technology 47 (2013) 1, 25–28 25
UDK 678.7:621.793 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)25(2013)
composites and their application as plane heaters using
composite films.
The fabrication of uniform graphite-flake carbon-
black/polystyrene composite films is impossible by using
conventional methods such as melting and heating
because of several difficulties. To circumvent this prob-
lem, as explained in ref.1, a similar gelation technology
was employed to dissolve polystyrene pellets in a
quickly evaporating solvent. Graphite-flake and carbon-
black powders were dispersed in a polystyrene-chloro-
form gel matrix with different concentrations. The
obtained gel containing these powders was deposited on
glass-fiber woven fabrics to obtain the intended plane
heaters. Then the structure, the surface morphology, as
well as electrical and heating properties of the deposited
composite films were investigated.
2 EXPERIMENTAL DETAILS
Graphite flakes (Tupras), carbon black (Selen Chemi-
stry), polystyrene pellets (Tupras Petkim) and chloro-
form (Aldrich) were commercially provided to produce a
new-generation heating element. Polystyrene pellets
were dissolved in chloroform using an ultrasonic bath for
60 min and subsequently a viscous gel was obtained
prior to depositing the composite films. Graphite-flake
and carbon-black powders were basically added to obtain
the gel as schematically illustrated in Figure 1. After
these processes, graphite-flake carbon-black-reinforced
polystyrene-matrix mixtures were prepared for the
plane-heating composite element. The strategy for
determining the composite samples was based on a
composite including the total powder weight against the
polystyrene weight with a ratio of 0.6 that is referred to
as X. Here the coefficient of X as 1, 2 and 4 means that
the ratios are 0.6, 1.2 and 2.4, respectively. Beside this,
the weight ratio of carbon black/graphite flake was
determined as C/G = 0.62 for the composites including
1X, 2X and 4X. In order to determine individual effects
of the powders on the properties, the composites
containing just carbon black and just graphite were
prepared as 1XC and 1XG, respectively. The details and
definitions of the above mixture specifications were
noted in Table 1.
Table 1: Mixture specifications and content data
Tabela 1: Pregled me{anic in podatki o vsebnosti
Precursor 4X* 2X* 1X* 1XG* 1XC*
Polystyrene (g) 1 1 1 1 1
Chloroform (ml) 8 8 8 8 8
Carbon black (g) 0.92 0.46 0.23 0.6 0
Graphite (g) 1.48 0.74 0.37 0 0.6
*X indicates a composite having the mass ratio C/G = 0.62
Coefficients of X indicate (C + G)/P
As the substrate choice is an important parameter for
any film-deposition technique, glass-fiber, woven fabric
substrates with a planar density of 100 g/m2 were chosen
because of their flexibility and high-temperature service
abilities. The produced mixtures were spray painted on
these substrates using an air-compressor-based mecha-
nism. After that deposited films were dried to remove the
solvent at 60 °C for 30 min on a hot plate for three times.
X-ray diffraction (XRD, Rigaku D/MAX-2200/PC)
patterns of the coatings were determined to identify the
phase structure using a diffractometer with CuK
irradiation. The surface properties of the films were
examined with the help of scanning electron microscopy
(SEM, JEOL JSM 6060).
M. EROL, E. ÇELIK: GRAPHITE-FLAKE CARBON-BLACK-REINFORCED POLYSTYRENE-MATRIX COMPOSITE ...
26 Materiali in tehnologije / Materials and technology 47 (2013) 1, 25–28
Figure 2: Experimental setup for determining electrical and heating
properties of composite films
Slika 2: Eksperimentalni sestav za dolo~anje elektri~nih in grelnih
lastnosti kompozitnih tankih plasti
Figure 1: Schematic illustration of the process
Slika 1: Shematski prikaz procesa
Electrical and heating properties of the composite
coatings were determined using the basic experimental
setup as illustrated in Figure 2. This basic setup con-
sisted of a digital thermometer, a 24-V direct-current
(DC) power supply and a multimeter (used as an amper-
meter in this circuit). Square specimens with the dimen-
sions of 30 mm × 30 mm were cut and electrodes were
painted on two edges with a silver paste as depicted with
the orange color in Figure 2. The specimens were fixed
to the circuit using copper jaws. The circuit current and
the time versus the temperature data were noted from
this set up. The resistances of the film were basically
calculated using Ohm’s law.
3 RESULTS AND DISCUSSION
The XRD patterns of the produced composite films
on the glass-fiber woven fabrics are demonstrated in
Figure 3 in order to prove that their structure consists of
both graphite flakes and carbon-black powder. As seen
from the XRD patterns, the main peaks of graphite and
carbon were determined at the diffraction angle of 26.6°.
With respect to increasing the total powder amount,
higher proportional-intensity data was recorded retaining
the mixture contents listed in Table 1. In addition,
crystalline-like polystyrene structure was clearly found
in this data. This pattern proves that polystyrene, gra-
phite flakes and carbon-black powder are physically
mixed in the composite structure as no structural or
chemical change was observed. As expected, another
observation of the XRD results showed a highly amor-
phous band of specimen 1XG and a highly crystalline
structure of 1XG.
Microstructural properties of the composite films
observed via SEM are shown in Figure 4. As can be
seen from the micrographs, an efficient composite
coating was prepared for the heating elements. It should
be noted that the round areas were considered to be the
bubbles left after the evaporation of the solvent due to an
increase in the pore number and size against the relative
solvent amount. It is also important to note here that the
other areas on the micrographs were found to be poly-
styrene matrices encapsulating powder agglomerations.
The electrical behavior of the films was determined
with the experimental setup (Figure 2 for details) and
the measured current data was used to calculate the
resistance values according to the basic Ohm’s law. The
measured and calculated electrical properties are listed in
Table 2. In the light of electrical measurements, the
samples 1X, 1XG and 1XC were found to be highly
insulating according to the resistance values. This is
because the above particle/composite weight ratio of 0.6
is found to be lower than the intended structure. Perco-
lation threshold is an important phenomenon, as reported
in the research by S. Isaji et al.1; however, there is no
study of a polystyrene matrix filled with both graphite
M. EROL, E. ÇELIK: GRAPHITE-FLAKE CARBON-BLACK-REINFORCED POLYSTYRENE-MATRIX COMPOSITE ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 25–28 27
Figure 4: SEM micrographs of the composite films on glass-fiber
woven fabrics including: a) 4X, b) 2X, c) X, d) 1XC, e) 1XG
Slika 4: SEM-posnetki kompozitne tanke plasti na tkanini iz steklenih
vlaken: a) 4X, b) 2X, c) X, d) 1XC, e) 1XG
Figure 3: XRD patterns of the composite films on glass-fiber woven
fabrics
Slika 3: Rentgenski posnetki kompozitnih tankih plasti na tkanini iz
steklenih vlaken
and carbon. In this study, the ratio of the powders loaded
by us can be regarded as lower than the percolation
threshold for the above samples. In this way, heating was
obtained from these samples, as noted in Figure 5,
which clarifies the variation of temperature with time
and sample type. According to Figure 5, the samples
such as 4X and 2X were determined to be good candi-
dates for a plane heating that can reach the temperatures
of 65 °C and 35 °C in 90 min, respectively.
Table 2: Electrical properties of the samples
Tabela 2: Elektri~ne lastnosti vzorcev
Sample code DC voltage(V) Current (mA)
Resistance
(kW)
4X 24 9.611 2.49
2X 24 19.04 1.26
1X 24 0.25 96
1XC 24 0.02 1200
1XG 24 0.01 2400
It is interesting that the samples evaluated in this
study were employed with a power supply unit of 24-V
DC. As a result, the plane-heater coatings with large
surfaces are applicable with 110-AC and 220-AC city
voltages, as listed in Table 2. It also needs to be pointed
out that this heating system does not emit any harmful
gaseous or liquid impurities unlike the heating systems
consuming fossil fuels. In addition, an easy layout and
application efficiency make the system a good candidate
heating device.
4 CONCLUSIONS
Graphite-flake carbon-black-reinforced polystyrene-
matrix composite films were successfully deposited on
glass-fiber woven substrates as plane heaters. Structural
and microstructural results indicate a good compatibility
between the powders and the matrix without any
chemical interaction among them and a good surface
quality with a partial orientation of the reinforcement
powders. Samples 4X and 2X were found to be good
candidate materials for plane-heating systems or devices.
The composite films are believed to be the new-gene-
ration heating elements due to being cost-effective,
environmentally friendly and energy friendly. For these
reasons, the composite films will start a new era in the
heating sectors.
Acknowledgement
This study was financially supported by Ayçe
Mühendislik Co. and the Turkish Ministry of Science,
Industry and Technology with the project code
00360-STZ-2009-1.
5 REFERENCES
1 S. Isaji, Y. Bin, M. Matsuo, Electrical conductivity and self-tem-
perature-control heating properties of carbon nanotubes filled poly-
ethylene films, Polymer, 50 (2009), 1046–1053
2 D. Bloor, K. Donnelly, P. J. Hands, P. Laughlin, D. Lussey, A metal-
polymer composite with unusual properties, J. Phys. D: Appl. Phys.,
38 (2005), 2851–2860
3 P. V. Notingher, D. Panaitescu, H. Paven, M. Chipara, Some cha-
racteristics of conductive polymer composites containing stainless
steel fibers, Journal of Optoelectronics and Advanced Materials, 6
(2004), 1081–1084
4 T. Yamamoto, E. Kubota, A. Taniguchi, S. Dev, K. Tanaka, K. Osa-
kada, Electrically conductive metal sulfide polymer composites pre-
pared by using organosols of metal sulfides, Chem. Mater., 4 (1992),
570–576
5 S. Geethaa, K. Kannan, S. Kumarb, S. Meenakshia, M. T. Vijayana,
D. C. Trivedia, Synergetic effect of conducting polymer composites
reinforced E-glass fabric for the control of electromagnetic radia-
tions, Composites Science and Technology, 70 (2010), 1017–1022
6 A. Russameeden, J. Pumchusak, Preparation of fiber-reinforced
electrically conducting polypropylene composites by wet-lay process
for use as bipolar plate in a proton exchange membrane fuel cell,
Journal of Metals, Materials and Minerals, 18 (2008), 121–124
7 A. B. Kaiser, Y. W. Park, Current-voltage characteristics of conduct-
ing polymers and carbon nanotubes, Synthetic Metals, 152 (2005),
181–184
8 K. R. Reddy, B. C. Sin, K. S. Ryu, J. C. Kim, H. Chung, Y. Lee,
Conducting polymer functionalized multi-walled carbon nanotubes
with noble metal nanoparticles: Synthesis, morphological characte-
ristics and electrical properties, Synthetic Metals, 159 (2009),
595–603
9 B. Chen, J. J. Liu, Y. Y. Sun, Study on thermal performance analysis
method of direct-gain solar dwellings with floor heating, Proceedings
of ISES World Congress, 3 (2009), 804–808
10 R. Zeng, X. Wang, W. Xiao, Y. Zhang, Q. Zhang, H. Di, Thermal
performance of phase change material energy storage floor for active
solar water-heating system, Front. Energy Power Eng. China, 4
(2010) 2, 185–191
11 Z. P. Song, R. Z. Wang, X. Q. Zhai, An experimental and simulation
study on performance of solar-powered floor heating system, Pro-
ceedings of ISES World Congress 2007, 5 (2009), 2224–2228
12 Z. G. Lin, S. Y. Zhang, X. Z. Fu, Y. Wang, Investigation of floor
heating with thermal storage, Journal of Central South University of
Technology, 13 (2006), 399–403
M. EROL, E. ÇELIK: GRAPHITE-FLAKE CARBON-BLACK-REINFORCED POLYSTYRENE-MATRIX COMPOSITE ...
28 Materiali in tehnologije / Materials and technology 47 (2013) 1, 25–28
Figure 5: Time versus the obtained temperature results for the
heating-element samples
Slika 5: Odvisnost ~asa in dose`ene temperature pri ogrevanju
vzorcev grelnih elementov
J. EKART et al.: POSTOPKI PRIPRAVE ZA PROIZVODNJO TRDNIH GORIV IZ NENEVARNIH ODPADKOV
POSTOPKI PRIPRAVE ZA PROIZVODNJO TRDNIH
GORIV IZ NENEVARNIH ODPADKOV
THE PROCESSES OF PREPARATION FOR THE PRODUCTION OF
SOLID FUELS FROM NON-HAZARDOUS WASTE
Janez Ekart1, Niko Samec1, Filip Kokalj1, Brigita Polanec2
1Univerza v Mariboru, Fakulteta za strojni{tvo, Smetanova 17, 2000 Maribor, Slovenija
2MD In`eniring, d. o. o., Lo{ka ulica 8, 2000 Maribor, Slovenija
janez.ekart@gmail.com
Prejem rokopisa – received: 2012-05-07; sprejem za objavo – accepted for publication: 2012-09-18
Namen raziskovalnega dela v predstavljenem ~lanku je bil z matemati~nim modelom poiskati re{itev za pripravo trdnih goriv iz
nenevarnih odpadkov, ki se lahko predelajo v trdno gorivo v skladu z veljavno nacionalno zakonodajo. Nenevarni odpadki so
definirani z DIREKTIVO 2008/98/ES Evropskega parlamenta in Evropskega sveta (19. 11. 2008) – Priloga III1. Z uporabo
matemati~nega modela in njegovim orodjem ob~utljivosti, vezanim na posamezne lastnosti vhodnih odpadnih materialov, smo iz
razpolo`ljivih odpadnih materialov z znanimi parametri njihovega organskega in anorganskega dela poiskali re{itev za njihovo
najbolj optimalno izrabo. Iz razpolo`ljivih odpadnih materialov smo v maksimalnem masnem dele`u pripravili zahtevano trdno
gorivo z vsemi potrebnimi kakovostnimi lastnostmi z najvi{jim mogo~im kakovostnim razredom in iz preostanka razpolo`ljivih
masnih tokov ostanek trdnega goriva slab{e kakovosti, ki ga pa {e vedno lahko plasiramo na trgu z mo`nostjo sose`iga v
cementarnah ali se`igalnicah odpadkov. Pri ugotavljanju kakovosti nenevarnih odpadkov in trdnega goriva je bilo pomembno
njihovo pravilno vzor~enje, ki se je izra`alo v stopnji homogenizacije vzorcev, ki imajo vsak zase svoje kemijske, fizikalne in
energijske lastnosti. Analiza kovinskega dela nenevarnih odpadkov je bila pomembna zaradi podatkov o pri~akovanih emisijskih
vrednostih v dimnih plinih, lete~em pepelu, pepelu in `lindri, ki nastane pri sose`igu trdnega goriva v kurilni napravi. Analiza
nekovinskega dela nenevarnih odpadkov je bila pomembna z vidika energijske vrednosti trdnega goriva in tehnolo{kih ter
okoljskih posledic. Rezultati raziskovalnega dela so pokazali, da je mogo~e iz razpolo`ljivih nenevarnih odpadkov dose~i
sorazmerno majhen masni dele` visokokakovostnega trdnega goriva iz odpadkov glede na celotno razpolo`ljivo maso odpadkov,
ki je sedaj v Sloveniji.
Klju~ne besede: nenevarni odpadki, trdno gorivo, kurilna vrednost, matemati~ni model, razred trdnega goriva
The purpose of the research work in the presented article was to find a solution, by means of a mathematic model, to the
problem of how to prepare solid fuels from non-hazardous waste that can be processed into solid fuels according to the national
law in force. Non-hazardous waste is defined by the DIRECTIVE 2008/98/ES of the European Parliament and the Council
(19.11.2008) – Annex III.1 With the application of a mathematical model and its sensitivity tool, depending on the individual
characteristics of the input waste material, the optimal utilization of different waste streams at disposal based on organic and
inorganic parameters for the most optimal solution for exploitation was calculated. From the waste material at disposal the
desired solid fuel was made in the maximum mass amount with all the necessary qualitative characteristics with the highest
quality class and from the rest of the available mass stream solid fuel of lower quality that can still be placed on the market for
incineration at cement kilns or incinerators was made. When defining the quality of non-hazardous waste and solid fuel one of
the most important tasks of the research was to sample the waste and the fuel correctly, which was reflected in the degree of
sample homogenization to find out the chemical, physical and energy characteristics. The analysis of the metal part of the
non-hazardous waste was important because of the data about the expected air-emission values of the flue gas, the fly ash, the
ash and the slag that occurs when incinerating solid fuel in the heating system. The analysis of the non-metal part of the
non-hazardous waste was important from the view of the calorific value of the solid fuel and of the technological and
environmental impact. The results of the research showed that it is possible to achieve from the non-hazardous waste at disposal
a relatively small mass fraction of high-quality solid fuel from waste according to the whole amount of mass available in
Slovenia at present.
Keywords: non-hazardous waste, solid fuels, calorific value, mathematical model, classification of solid fuels
1 UVOD
Za zmanj{anje obremenitev odlagali{~ in pove~evanje
izkori{~anja potenciala odpadkov so na podro~ju rav-
nanja z odpadnimi materiali nujne druge re{itve v
primerjavi z uveljavljenimi v na{i dr`avi. Ker imajo vsi
organski odpadni materiali neko kurilno vrednost, je
smiselna tudi re{itev z energijsko izrabo tistih odpadnih
materialov, ki niso primerni za snovno izrabo. Trdna
goriva iz odpadnih materialov imajo vi{jo stopnjo ener-
gijske izrabe odpadkov v primerjavi s klasi~nimi se`igal-
nicami odpadkov, saj temeljijo na izrabi alternativnih
energijskih virov in ne na odstranjevanju odpadkov, kot
to velja za klasi~ne se`igalnice.
Te`i{~e raziskovalnega dela sta uporabljena metoda
laboratorijskih analiz razpolo`ljivih odpadnih materialov
in matemati~ni model me{anja teh odpadnih materialov s
ciljem proizvodnje trdnih goriv z `elenimi lastnostmi.
Dodatna laboratorijska analiza trdnega goriva in analiza
emisijskih vrednosti pri se`igu trdnega goriva je
zadostna osnova, da v nadaljnjem raziskovalnem delu
poi{~emo odmike od pri~akovanih rezultatov in temu
primerno z analizo ob~utljivosti matemati~nega modela
izvr{imo korekcijo dele`ev odpadnih materialov v
trdnem gorivu.
Materiali in tehnologije / Materials and technology 47 (2013) 1, 29–35 29
UDK 519.61/.64:662.62 ISSN 1580-2949
Izvirni znanstveni ~lanek/Original scientific article MTAEC9, 47(1)29(2013)
2 EKSPERIMENTALNI DEL
Za dolo~itev fizikalnih in kemijskih lastnosti trdnega
goriva imamo na razpolago laboratorijske analize razpo-
lo`ljivih frakcij nenevarnih odpadkov, in sicer:
– vsebnost vlage,
– vsebnost pepela,
– GCV – zgorevalna toplota (MJ/kg),
– masni dele` Cl (%),
– vsebnost Hg (mg/MJ),
– vsebnost Cd (mg/kg),
– masni dele` S (%).
Pogoj za pridobivanje laboratorijskih rezultatov za
frakcije razpolo`ljivih odpadkov in trdnega goriva sta
bila njihovo predhodno vzor~enje in homogenizacija.
Vzor~enje in homogenizacija vzorcev sta potekala v
skladu s tehni~nimi specifikacijami in standardi: EN
14899:2005,2 SIST EN 15442:2011,3 SIST EN 15443:
20114 in SIST-TS CEN/TS 15413:2007.5 Za homoge-
nizacijo vzorcev frakcij odpadkov in trdnega goriva sta
bila uporabljena:
– rezalni mlin Retsch SM 2000 za ve~stopenjsko
mletje 0,8 cm, 0,4 cm ter 0,2 cm in
– centrifugalni mlin Retsch ZM 200 za homogeniza-
cijo vzorcev.
Dolo~evanje vlage je potekalo v skladu s standardom
SIST EN 15414-3:2011,6 uporabljena je bila `arilna
komora Nabertherm GmbH.
Dolo~evanje zgorevalne toplote (zgornje kurilne
vrednosti) je potekalo s kalorimetrom IKA Werke C5000
in z uporabo adiabatnega delovnega procesa z validirano
metodo po DIN 51900, kar je v skladu s standardom
SIST EN 15400:2011.7
Preostale uporabljene metode:
– Hg (mg/kg) v suhi snovi – ISO 5666:1999, pogl. 5,
modif.8
– Cd (mg/kg) v suhi snovi – SIST EN ISO 17294-2:
20059
– S (%) – izlu`ek – ASTM D 4239 (metoda C):199710
– Cl (%) – metoda titracije s titratorjem Metrohm
TITRANDO 809 po SIST EN 15408:201111
2.1 Uporaba matemati~nega modela
Pri izdelavi in uporabi matemati~nega modela smo
izhajali iz variante, s katero bi iz razpolo`ljivih frakcij
odpadkov pridobili trdno gorivo z maksimalno mogo~o
kakovostjo in ostanek trdnega goriva s slab{o kakovostjo.
Na osnovi predpisanih meril za dolo~itev kakovost-
nega razreda goriva so bili upo{tevani naslednji osnovni
parametri: GCV, Cl, Hg, Cd in S.
Celotna masa odpadkov za predelavo v trdno gorivo
je izra`ena z ena~bo:
m ml i
i
n
=
=
∑
1
n  N (1)
mI – skupna masa vseh frakcij odpadkov, ki so na razpo-
lago (kg)
mi – masa i-te frakcije odpadkov (kg)
Podatki za maso posamezne frakcije so bili pridob-
ljeni z eviden~nimi listi, ki so pri prevzemu odpadkov
obvezni.
Ker je bilo na razpolago ve~ frakcij odpadkov, je bilo
pri sestavi trdnega goriva treba upo{tevati njihove last-
nosti in dele`e. Dele`i frakcij odpadkov so bili izra`eni z
ena~bo:
di = mi/mI n  N (2)
di – masni dele` posamezne frakcije odpadkov
mi – masa i-te frakcije odpadkov
mI – skupna masa odpadkov
V imenovalcu je bila celotna masa vseh razpolo`lji-
vih frakcij odpadkov.
Posamezen parameter v trdnem gorivu je bil izra~u-
nan po ena~bi:
G
G m
mI j
I j
i
n
I
l
( )
( )
=
⋅
=
∑
1
n ∈ N (3)
j – parameter (GCV, Cl, Cd, Hg, S)
GI(j) – vrednost j-tega parametra skupnega goriva iz od-
padkov
Gi(j) – vrednost ra~unanega parametra i-te frakcije odpad-
kov po j-tem parametru
mi – masa i-te frakcije odpadkov
mI – skupna masa odpadkov
V Sloveniji imamo kakovost trdnega goriva oprede-
ljeno z Uredbo,12 ki definira kakovostne razrede trdnega
goriva v Prilogi 3 (tabela 1).
Z matemati~nim modelom so bili izra~unani relativni
dele`i posameznih frakcij odpadkov, ki so nam bili na
razpolago, z upo{tevanjem njihovih posameznih para-
metrov v odvisnosti od referen~nega parametra, katerega
razred za trdno gorivo je dolo~il prevzemnik.
Relativne vrednosti parametrov posamezne frakcije
odpadkov so bile izra~unane po ena~bi:
J. EKART et al.: POSTOPKI PRIPRAVE ZA PROIZVODNJO TRDNIH GORIV IZ NENEVARNIH ODPADKOV
30 Materiali in tehnologije / Materials and technology 47 (2013) 1, 29–35
Tabela 1: Podatki osnovnih frakcij odpadkov
Table 1: Data of the basic waste fractions
Frakcije odpadkov Masa, m/t GCV/(MJ/kg) w(Cl)/% (Cd)/(mg/kg) (Hg)/(mg/MJ) w(S)/%
Lahka frakcija MKO 120 15,9 0,48 1,8 0,22 0,02
Odpadki iz industrije (Gorenje) 50 21,7 0,55 2 0,32 0,12
Ostanek po razvr{~anju embala`e 30 14,2 0,84 1,6 0,10 0,02
p
G
Gi j
i j
j
( )
( )
( ) .
=
ref
(4)
pi(j) – relativna vrednost i-te frakcije po j-tem parametru
Gi(j) – vrednost ra~unanega parametra j i-te frakcije
G(j)ref. – `elena vrednost parametra j
Tako so bile izra~unane relativne vrednosti posamez-
nih parametrov v odnosu na `elene (referen~ne) para-
metre za vse frakcije odpadkov.
Da bi lahko dolo~ili vrednosti optimalnih lastnosti
trdnega goriva, smo uporabili izlo~itvena merila za vsak
parameter in za vsako frakcijo odpadkov posebej (if
relativna vrednost j-tega parametra ≥1, zapi{i 1, druga~e
pa 0). Z upo{tevanjem izlo~itvenih meril po posameznih
parametrih j je bila za optimalne lastnosti trdnega goriva
dolo~ena nova skupna masa osnovnih frakcij odpadkov,
ki je bila izra~unana po ena~bi:
m m d pv j I i j
i
n
( ) ( )( )= ⋅
=
∑
1
n  N (5)
mV(j) – nova masa po izlo~itvi
Za prime{anje posameznih frakcij odpadkov smo
dolo~ili variacijski razmik (VR) med lastnostmi optimal-
nega trdnega goriva in trdnega goriva po zahtevah pre-
vzemnika:
VR G G j= −opt. ref( ) . (6)
Gopt. – vrednost j-tega parametra optimalnega trdnega go-
riva
G(j)ref. – vrednost j-tega parametra trdnega goriva, ki
ustreza zahtevam prevzemnika
Parameter prime{anja Gj dolo~imo z variacijskim
razmikom in maso po izlo~itvi. ^e je parameter Gj enak
0, je prime{anje dodatne frakcije dovoljeno, sicer ne. Za
izra~un teoreti~no mogo~e mase frakcije odpadkov za
prime{anje smo uporabili ena~bo:
m
VR m
G Gj
V j
j i j
dom
ref
( )
( )
( ) . ( )
=
⋅
−
(7)
mdom(j) – teoreti~na mogo~a masa prime{anja za
parameter j
Za prime{anje smo pri izra~unu po matemati~nem
modelu upo{tevali, da lastnosti enega izmed osnovnih
frakcij odpadkov ustrezajo lastnostim referen~nega trd-
nega goriva, kot ga zahteva prevzemnik. Zato smo dolo-
~ili, da tej frakciji odpadkov prime{amo razpolo`ljivi
preostali osnovni frakciji odpadkov s ciljem, da iz
obstoje~e osnovne mase odpadkov zagotovimo maksi-
malno mogo~o maso trdnega goriva, ki bo ustrezalo
referen~nim vrednostim. Glede na lastnosti frakcij
odpadkov smo dolo~ili prioritetni red prime{anja. Kot
prvi prioritetni red prime{anja smo uporabili frakcijo
odpadkov, ki je v najve~jem dele`u posameznih para-
metrov j1,5 bila najbli`je parametrom referen~nega
trdnega goriva (kot ga zahteva prevzemnik). Ker smo
imeli tri dodatne frakcije odpadkov in pet parametrov, ki
smo jih analizirali v posamezni frakciji odpadkov za
prime{anje, smo uvedli matriko za prime{anje s tremi
stolpci in petimi vrsticami. Pri tem je prvi stolpec
pomenil frakcijo odpadkov V1, drugi stolpec frakcijo
odpadkov V2 in tretji stolpec frakcijo odpadkov V3.
Vrstice so pomenile posamezne parametre j1,5. Merilo za
formiranje matrike:
– prva vrstica, prvi stolpec matrike (1,1) : if V1 =
razredref. – vpi{i 1, druga~e vpi{i 0
Za prvi prioritetni red prime{anja smo dobili matriko:
0 1 0
0 1 0
0 1 0
0 1 0
0 1 0
Rezultat matrike pove, da smo v prvem prioritetnem
redu prime{anja uporabili frakcijo odpadkov V2. Po
enakem postopku smo dobili matriko drugega in tretjega
prioritetnega reda prime{anja. Po dolo~itvi prioritetnih
redov prime{anja frakcij odpadkov smo dolo~ili koli~ino
(maso) prime{anja frakcije odpadkov na podlagi meril
posameznih parametrov j1,5 in ena~be:
m m Pj P n jxi( ) . ( )1 = ⋅dej (8)
m(j)P1 – masa dome{avanja frakcije odpadkov prve priori-
tete po j-tem parametru
mdej. – dejanska masa prime{anja vseh frakcij odpadkov
Pn(jxi) – pripadajo~i prioritetni red me{anja
Za izra~un vrednosti skupne mase smo uporabili
ena~bo (primer parameter GCV):
m m mjS j P V jn = +( ) ( )1 (9)
mjSn – skupna masa n-tega materiala po j-tem parametru
m(j)P1 – masa prime{anja prvega prioritetnega reda po
j-tem parametru
Po enakem sistemu smo izra~unali koli~ine prime-
{anja za druge parametre frakcije odpadkov V1 in nato
koli~ine prime{anja v drugem in tretjem prioritetnem
razredu.
Kon~no maso prime{anja za optimalno trdno gorivo
smo izra~unali po ena~bi:
m m m p
i
n
iTAG ref.
= +
=
∑
1
(10)
mTAG – masa kon~nega goriva `elenega razreda
mref. – masa vzorca trdnega goriva, ki izpolnjuje pogoje
`elenega razreda
mpi – vsota mas prime{anih frakcij odpadkov k vzorcu
trdnega goriva, ki izpolnjuje pogoje `elenega razreda
Po prime{anju frakcij odpadkov v prvem, drugem in
tretjem prioritetnem redu prime{anja smo dobili trdno
gorivo, ki je izpolnjevalo pogoje `elenega razreda v
celotni masi TAG in ostanek trdnega goriva slab{ega
kakovostnega razreda, ki bi ga lahko hkrati se`igali v
cementarnah ali se`igalnicah odpadkov.
Slika 1 prikazuje diagram poteka priprave refe-
ren~nega trdnega goriva in ostanka trdnega goriva. Na
J. EKART et al.: POSTOPKI PRIPRAVE ZA PROIZVODNJO TRDNIH GORIV IZ NENEVARNIH ODPADKOV
Materiali in tehnologije / Materials and technology 47 (2013) 1, 29–35 31
sliki 2 je prikazan model osnovnega masnega toka
frakcij odpadkov in prime{anja lo~enih frakcij z uporabo
matemati~nega modela, s ~imer dose`emo produkt
trdnega goriva po zahtevah prevzemnika in ostanek
trdnega goriva slab{e kakovosti.
Ker smo imeli na razpolago tri dodatne frakcije
odpadkov, smo k ostanku trdnega goriva prime{ali lo~ene
frakcije odpadkov po naslednji ena~bi:
m m m p
i
n
iost. 0
= +
=
∑
1
n  N (11)
most. – masa ostanka goriva po prime{anju lo~enih frakcij
odpadkov
m0 – masa ostanka goriva pred prime{anjem
mpi – vsota mas prime{anih frakcij odpadkov k ostanku
trdnega goriva
3 REZULTATI
Na razpolago smo imeli 200 t osnovnih frakcij od-
padkov in 50 t lo~enih frakcij za prime{anje. V koli~ini
200 t odpadkov so bili taki s parametri, dolo~enimi na
podlagi njihovega vzor~enja in laboratorijskih analiz
(tabela 1). Podatke o lo~enih frakcijah za prime{anje
prikazuje tabela 2.
Z uporabo celotne osnovne mase frakcij odpadkov in
njihovih lastnosti, ki jih prikazuje tabela 1, smo z upo-
rabo ena~be (3) izra~unali parametre trdnega goriva
(tabela 3).
V skladu s Prilogo 3 Uredbe12 smo pridobljeno
gorivo zaradi kurilne vrednosti uvrstili v razred 3. Drugi
parametri so ustrezali trdnemu gorivu razreda 2, nekateri
celo razredu 1.
J. EKART et al.: POSTOPKI PRIPRAVE ZA PROIZVODNJO TRDNIH GORIV IZ NENEVARNIH ODPADKOV
32 Materiali in tehnologije / Materials and technology 47 (2013) 1, 29–35
Slika 1: Procesni diagram priprave referen~nega trdnega goriva in
ostanka trdnega goriva
Figure 1: Flow chart of the preparation of the reference solid fuel and
its residues
Slika 2: Model osnovnega masnega toka in dome{avanja lo~enih frakcij odpadkov z uporabo matemati~nega modela
Figure 2: Model of the basic mass stream and co-mixing of the separated waste fractions with the use of the mathematical model
Tabela 2: Lo~ene frakcije odpadkov za dome{avanje
Table 2: Separated waste fractions for co-mixing
Frakcije odpadkov Masa, m/t GCV/(MJ/kg) w(Cl)/% (Cd)/(mg/kg) (Hg)/(mg/MJ) w(S)/% 4
Tekstil 20 30,5 0,23 2 0,10 0,01
Odpadni les 20 7,4 0,19 2 0,32 0,12
Rumene vre~e 10 20,9 0,7 2 0,10 0,01
Potencialni prevzemnik je zahteval trdno gorivo
razreda 2, kar je za doseganje vi{je kurilne vrednosti
goriva pomenilo potrebo po prime{anju k osnovnemu
trdnemu gorivu razreda 3. Za pripravo trdnega goriva
razreda 2 smo z izlo~itvenimi merili z uporabo ena~be
(5) izra~unali maso trdnih goriv iz osnovnih frakcij
odpadkov, pri ~emer smo dobili trdno gorivo kakovosti,
ki jo prikazuje tabela 4.
Sicer smo `e z uporabo izlo~itvenih meril pri osnovni
masi frakcij odpadkov teoreti~no dosegli dve kakovosti
trdnega goriva, od katerih je eno ustrezalo razredu 2 kot
na{emu referen~nemu gorivu. S prime{anjem frakcije
odpadkov iz osnovnega masnega toka k referen~nemu
gorivu smo `eleli pove~ati maso referen~nega goriva. Z
uporabo ena~b od (6) do (10) smo s prime{anjem z upo-
rabo osnovnih frakcij odpadkov dosegli pove~anje mase
referen~nega goriva razreda 2 in zmanj{ali maso ostanka
trdnega goriva, pri kateri se je zmanj{ala vrednost GCV
in je zato bil uvr{~en v razred 4.
Rezultati prime{anja v obsegu osnovnega masnega
toka frakcij odpadkov so prikazani v tabeli 5.
S prime{anjem dodatne lo~ene frakcije odpadkov V4
– tekstil k ostanku trdnega goriva iz osnovne mase
frakcij odpadkov po ena~bi (11) smo v prvem koraku
izbolj{ali kakovost ostanka trdnega goriva iz osnovne
mase frakcij odpadkov, tako da je ostanek trdnega goriva
iz razreda 4 pre{el v razred 3. S prime{anjem vseh
dodatnih lo~enih frakcij odpadkov k ostanku trdnega
goriva iz osnovnega masnega toka frakcij odpadkov smo
dobili trdno gorivo razreda 4, vendar z bistveno vi{jim
GCV, kot ga je imel ostanek iz osnovnega masnega toka
frakcij odpadkov. Vsi drugi parametri so ustrezali raz-
redu 1 oziroma razredu 2.
Rezultate prime{anja lo~enih frakcij odpadkov k
ostanku trdnega goriva iz osnovnega masnega toka
prikazuje tabela 6.
4 DISKUSIJA
Fizikalno-matemati~na izhodi{~a niso upo{tevala
vseh procesov, ki potekajo pri predelavi med proizvod-
njo, kot so na primer mletje in izlo~evanje anorganskih
materialov, temve~ smo obravnavali samo vhod in izhod
procesa. Matemati~ni model omogo~a izra~un parame-
trov trdnega goriva s poljubnim {tevilom frakcij odpad-
kov, vendar je z vidika tehnolo{ke re{itve doziranja
frakcij odpadkov primerno ra~unati z isto~asnim dozira-
njem do maksimalno treh frakcij odpadkov. Za dolo~e-
vanje lastnosti goriva je bil upo{tevan zakon o ohranitvi
mase, zato so bili za vsak masni tok frakcije odpadkov
pomembni vsi parametri.
Ker je proizvodnja goriva zasnovana na mehanski
obdelavi, so se lastnosti trdnega goriva izra`ale samo v
velikosti delcev in homogenizaciji. Izlo~ene kovine niso
pomenile pomembnega dele`a in v matemati~nem mo-
delu v masnem toku niso bile upo{tevane. Na podlagi
analiti~nih podatkov laboratorijskih analiz se je njihovo
izlo~anje v procesu kazalo v pove~anju vrednosti GCV
trdnega goriva.
Strokovna literatura za posamezne parametre trdnih
goriv obravnava problematiko trdnih goriv parcialno, ne
obravnava pa meril, kako dose~i konfekcionirano trdno
gorivo na podlagi izbire odpadnih materialov, ki imajo
vsak zase dolo~ene kemijske, fizikalne in energijske
lastnosti. V strokovnih ~lankih, ki obravnavajo podro~je
trdnih goriv, podro~ja uporabe matemati~nega modela za
izra~un kemijskih in fizikalnih lastnosti trdnih goriv ni
bilo mogo~e najti. Sicer so v strokovni literaturi13 nave-
dene ugotovitve o tem, da je osnova postopka ocenje-
vanja lastnosti celotno stanje materialov. Uporaba ~iste
mehanske separacijske tehnike zagotavlja, da se zmes z
vidika materialov vhodnih odpadkov ne preoblikuje in da
razpolo`ljive separacijske tehnike ne pu{~ajo materialov
v sistemu. Z uporabo zakona o ohranitvi mas in energije
je stanje na vhodu enako stanju vsote izhodnih masnih
tokov. Velja ena~ba:
Xinput (S) = Minput × C input (S) = MP CP si i
i
k
⋅
=
∑ ( )
1
(12)
pri ~emer je:
X – obremenitev
k – {tevilo izhodov
C – koncentracija
M – referen~na masa funkcionalne enote (kilogram
ostanka na tono vstopnega materiala)
P1 – indikator izhod 1
Pi – indikator izhod i
S – substanca elementa
S preizkusi je bilo ugotovljeno, da so kemijski ele-
menti, kot so klor, kadmij in svinec, v odpadkih pogosto
koncentrirani. Tudi z optimirano tehniko lo~evanja je
mo`nost za njihovo zmanj{anje omejena zaradi razdelit-
ve koncentracije teh elementov v razli~nih komponentah
masnega toka odpadkov. V nasprotju s Cd in Pb je pri-
spevek cinka v odpadkih bolj enakomerno porazdeljen.
Podatki o koncentracijah nevarnih snovi v odpadkih in
njihova porazdelitev v gospodinjskih odpadkih (me{ani
komunalni odpadki) ter rezultati lo~evanja z uporabo
razli~nih tehnik lo~evanja ka`ejo, da so mehanske
operacije lo~evanja zadostne za izlo~anje odpadkov, ki
vsebujejo nevarne snovi.
Ugotovljeno je, da dele` plastike iz me{anih komu-
nalnih odpadkov nara{~a z lo~evanjem le-teh v bobna-
stem situ od 26,3 % na 57,8 %. Prav tako vsebnost vlage
v me{anih komunalnih odpadkih pade s tem na~inom
lo~evanja od 50 % na 40 %, kar je z vidika priprave
trdnega goriva pomemben podatek. Vendar pa je vpliv
predhodno sortiranih trdnih odpadkov na objekte za
pridobivanje energije v razli~nih dr`avah {e vedno nego-
tov zaradi inherentne kompleksnosti sestave trdnih
odpadkov.14
[tudija uporabe RDF je ugotovila, da njegova upo-
raba v cementnih pe~eh namesto premoga ali koksa
ponuja okoljske prednosti pri zmanj{anju toplogrednih
J. EKART et al.: POSTOPKI PRIPRAVE ZA PROIZVODNJO TRDNIH GORIV IZ NENEVARNIH ODPADKOV
Materiali in tehnologije / Materials and technology 47 (2013) 1, 29–35 33
plinov, medtem ko nastanek polutantov v plinastem
agregatnem stanju ni pomemben dejavnik.15 Zanimiva je
tudi primerjava med me{animi komunalnimi odpadki in
RDF za primer energijske izrabe na obmo~ju Castilla y
Leona ([panija). Tako je bila ugotovljena vsebnost vlage
v organskem delu me{anih komunalnih odpadkov
58,33 % in v RDF 42,26 %, pri analizi vsebnosti vlage v
papirju in kartonu me{anih komunalnih odpadkov pa je
ugotovljeno 55,59 % vlage, v RDF pa samo 16,72 %.
Posledi~no temu je bila tudi GCV v dostavljenem stanju
me{anih komunalnih odpadkov 10,16 kJ/kg, v RDF pa
18,28 kJ/kg.16
Proizvajalci trdnih goriv lahko matemati~ni model
uporabljajo samo ob izpolnjevanju pogojev za zago-
tavljanje kakovosti trdnega goriva iz odpadkov v skladu s
standardom SIST EN15358:2011 Trdna alternativna
goriva – Sistemi vodenja kakovosti,17 ki vklju~ujejo
stalno kontrolo odpadkov na vhodu in produkta trdnega
goriva na izhodu z njihovim vzor~enjem ter izvedbo
laboratorijskih analiz.
Raziskovalno delo, ki je predstavljeno v tem ~lanku,
daje mo`nost nadaljnje raziskave trdnih goriv predvsem
na osnovi ugotovljenih emisij snovi v zrak pri zgorevanju
trdnega goriva. Z uporabo teoreti~nih prera~unov emisij
je mogo~e izvesti primerjave z eksperimentalno dolo~e-
nimi vrednostmi emisij snovi v dimnih plinih pri zgore-
vanju trdnega goriva iz odpadkov. Za celovito bilanco
transformacije {kodljivih snovi pa so bile opravljene tudi
analize pepela in `lindre, ki sta ostala po zgorevanju. To
omogo~a postavljanje matemati~nega modela, ki bo
napovedoval vpliv zgorevanja posameznega goriva na
okolje `e na osnovi temeljite analize vzorca trdnega
goriva.
Zahvala
Avtorji smo hvale`ni Javni agenciji za tehnolo{ki
razvoj Republike Slovenije in Evropski uniji oziroma
Evropskemu skladu za regionalni razvoj za sofinan-
ciranje tega razvojno investicijskega projekta (RIP 09) z
naslovom "Obnovljivi viri energije iz odpadkov –
postopki predelave odpadkov v trdno gorivo in njegova
energijska izraba s se`igom in uplinjanjem".
Acknowledgement
The authors gratefully acknowledge the Public
Agency for Technology of the Republic of Slovenia and
the European Union; the European Fund for Regional
Development for co-funding this research investment
project in the framework RIP 09 with the title "Renew-
able energy from waste – the processes of waste
processing into solid fuel and its energy recovery with
combustion and gasification".
5 LITERATURA
1 DIREKTIVA 2008/98/ES Evropskega parlamenta in Evropskega
sveta (19. 11. 2008) – Priloga III
2 EN 14899:2005; Karakterizacija odpadkov-vzor~enje odpadkov –
Okvirno navodilo za pripravo in uporabo na~rta vzor~enja
3 SIST EN 15442:2011; Trdno alternativno gorivo – Metode vzor~enja
4 SIST EN 15443:2011; Trdno alternativno gorivo – Metode za pri-
pravo laboratorijskega vzorca
5 SIST TS CEN/TS 15413:2007; Trdno alternativno gorivo – Metode
za pripravo testnega vzorca iz laboratorijskega vzorca
J. EKART et al.: POSTOPKI PRIPRAVE ZA PROIZVODNJO TRDNIH GORIV IZ NENEVARNIH ODPADKOV
34 Materiali in tehnologije / Materials and technology 47 (2013) 1, 29–35
Tabela 3: Parametri trdnega goriva z uporabo celotne mase osnovnih
frakcij odpadkov
Table 3: Parameters for the solid fuel produced with the total mass of
the basic waste fractions
Parameter Vrednost Razred
Masa, m/kg 200000
GVC/(MJ/kg) 17,095 3
w(Cl)/% 0,5515 2
(Hg)/(mg/MJ) 0,013044861 1
(Cd)/(mg/kg) 1,82 2
w(S)/% 0,045 1
Tabela 4: Trdno gorivo iz osnovnih frakcij odpadkov z uporabo
izlo~itvenih meril
Table 4: Solid fuel produced from basic fraction using the eliminating
criteria
Trdno gorivo Razred 2 Razred 3
Parameter Enota Vrednost Razred Vrednost Razred
masa, m kg 50000 150000
GVC MJ/kg 21,7 2 15,56 3
w(Cl) % 0,550 2 0,552 2
(Hg) mg/MJ 0,0147 1 0,01248 1
(Cd) mg/kg 2,0 2 1,76 2
w(S) % 0,12 1 0,02 1
Tabela 5: Trdno gorivo z dome{avanjem v obsegu osnovnega masnega
toka
Table 5: Solid fuel produced with co-mixing from the basic mass
stream
Trdno gorivo Razred 2 Razred 4
Parameter Enota Vrednost Razred Vrednost Razred
masa, m kg 70731,7 129268,3
GVC MJ/kg 24,279 2 13,164 4
w(Cl) % 0,456 2 0,6036 3
(Hg) mg/MJ 0,0114 1 0,0139 1
(Cd) mg/kg 2,0 2 1,7215 2
w(S) % 0,0878 1 0,0216 1
Tabela 6: Trdno gorivo z dome{avanjem dodatnih lo~enih frakcij
odpadkov
Table 6: Solid fuel produced with co-mixing from the additional
separated waste fractions
Trdno gorivo Dome{avanje V4 Dome{avanje V4,V5 in V6
Parameter Enota Vrednost Razred Vrednost Razred
masa, m kg 149268 179268
GVC MJ/kg 15,487 3 14,887 4
w(Cl) % 0,5535 2 0,521 2
(Hg) mg/MJ 0,0126 1 0,016 1
(Cd) mg/kg 1,7588 2 1,799 2
w(S) % 0,02 1 0,049 1
6 SIST EN 15414-3:2011; Trdno alternativno gorivo – Dolo~evanje
vsebnosti vlage z metodo su{enja v su{ilni komori – 3. del: Vlaga v
preizkusnem vzorcu
7 SIST EN 15400:2011; Trdno alternativno gorivo – Dolo~evanje
kurilne vrednosti
8 ISO 5666:1999, pogl. 5, modif.; metoda za dolo~evanje `ivega srebra
(standard, korigiran z ISO 12846:2012)
9 SIST EN ISO 17294-2:2005; Dolo~evanje te`kih kovin z uporabo
induktivno sklopljene plazme z masnim selektivnim detektorjem
(ICP-MS)
10 ASTM D 4239 (metoda C):1997; standardna testna metoda za dolo-
~evanje vsebnosti `vepla v zgorevalni komori z visokotemperaturnim
re`imom
11 SIST EN 15408:2011; Trdna alternativna goriva – Metode za dolo-
~evanje `vepla (S), klora (Cl), fluora (F) in broma (Br)
12 Uredba o predelavi nenevarnih odpadkov v trdno gorivo, Ur. l. RS {t.
57/2008
13 V. S. Rotter, T. Kost, J. Winkler, B. Bilitewski, Material flow analysis
of RDF Production processes, Waste management, (2004),
1005–1021
14 N. B. Chang, Y. H. Chang, W. C. Chen, Evaluation of heat value and
its prediction for refuse-derived fuel, The Science of the Total
Environment, (1997), 139–148
15 G. Genon, E. Brizio, Perspectives and limits for cement kilns as a
destination for RDF, Waste management, (2008), 2375–2385
16 C. Montejo, C. Costa, P. Ramos, M. C. Marquez, Analysis and
comparison of municipal solid waste and reject fraction as fuels for
inceneration plants, Applied Thermal Engineering, (2011),
2135–2140
17 SIST EN15358:2011 Trdna alternativna goriva – Sistemi vodenja
kakovosti
J. EKART et al.: POSTOPKI PRIPRAVE ZA PROIZVODNJO TRDNIH GORIV IZ NENEVARNIH ODPADKOV
Materiali in tehnologije / Materials and technology 47 (2013) 1, 29–35 35

M. BABI^ et al.: PROBLEMS ASSOCIATED WITH A ROBOT LASER CELL USED FOR HARDENING
PROBLEMS ASSOCIATED WITH A ROBOT LASER CELL
USED FOR HARDENING
PROBLEMATIKA ROBOTSKEGA LASERSKEGA KALJENJA
Matej Babi~1, Matja` Milfelner2, Igor Beli~3, Peter Kokol4
1Emo-Orodjarna, d. o. o., Be`igrajska cesta 10, 3000 Celje, Slovenia
2Tic-Lens, d. o. o., Be`igrajska cesta 10, 3000 Celje, Slovenia
3Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
4University of Maribor, Faculty of Health Sciences, @itna ulica 15, 2000 Maribor, Slovenia
babicster@gmail.com
Prejem rokopisa – received: 2012-05-29; sprejem za objavo – accepted for publication: 2012-07-30
Laser hardening is a surface-hardening process. It is used exclusively on ferrous materials suitable for hardening, including steel
and cast iron with a carbon content of more than 0.2 %. This article describes robot laser hardening, the results of previous
work, research and experience with robot laser hardening. The second part of the paper describes the problems associated with
robot laser hardening at different angles. We wanted to find the impact of the angles on the hardness of the material. Therefore,
we directed the laser beam at different angles, including perpendicular, in the process of hardening. We made test patterns of a
standard label on the materials of DIN standard 1.2379.
Keywords: hardening, robot, laser, parameters
Lasersko kaljenje je proces povr{inskega utrjevanja. Uporablja se izklju~no za `elezne materiale, ki so primerni za kaljenje in
vsebujejo ve~ kot 0,2 % ogljika. V ~lanku opisujemo robotsko lasersko kaljenje, navajamo rezultate dosedanjega dela in
raziskav ter izku{nje z laserskim kaljenjem. Drugi del opisuje problematiko robotskega laserskega kaljenja pri razli~nih kotih.
@eleli smo ugotoviti, kako kot laserskega `arka vpliva na trdoto materiala. Kot laserskega `arka smo spreminjali glede na smer
potovanja, kakor tudi pravokotno na smer potovanja laserskega `arka. Naredili smo vzorce standardne oznake po DIN standardu
1.2379.
Klju~ne besede: kaljenje, robot, laser, parametri
1 INTRODUCTION
Laser hardening1–5 is a metal-surface treatment pro-
cess that complements the conventional flame- and
induction-hardening processes. A high-power laser6–10
beam is used to heat the metal surface rapidly and
selectively to produce hardened case depths of up to 1.5
mm with hardness values of up to 65 HRc. This has a
hard martensitic microstructure providing improved
properties, such as wear resistance and increased
strength. To harden the workpiece, the laser beam
usually warms the outer layer to just under the melting
temperature (about 900 °C to 1400 °C). Once the desired
temperature is reached, the laser beam starts moving. As
the laser beam moves, it continuously warms the surface
in the processing direction.11,12 The high temperature
causes the iron atoms to change their position within the
metal lattice (austenization). As soon as the laser beam
moves away, the hot layer is very rapidly cooled by the
surrounding material in a process known as self-hard-
ening. This rapid cooling prevents the metal lattice from
returning to its original structure and producing marten-
site. The laser beam hardens the outer layer or case of
the workpiece. The hardening depth of the outer layer is
typically from 0.1 mm to 1.5 mm. However, on some
materials, it may be 2.5 mm or more. A greater harden-
ing depth requires a larger volume of the surrounding
material to ensure that the heat dissipates quickly and the
hardening zone cools fast enough. Relatively low power
densities are needed for hardening. At the same time, the
hardening process involves the treatment of extensive
areas of the surface. That is why the laser beam is shaped
so that it irradiates an area that is as large as possible.
This irradiated area is usually rectangular. Scanning
optics are also used in hardening. They are used to move
a laser beam with a round focus back and forth very
rapidly, creating a line on the work piece with a power
density that is virtually uniform. This method makes it
possible to produce hardened tracks up to 60 mm wide.
2 EXPERIMENTAL METHOD AND MATERIALS
PREPARATION
A robot laser cell can be used to provide the heat
necessary for a treatment process. The absorbed
radiation from the laser of the laser cell heats up the
surface layer to a temperature where austenite can form.
In this work we research how the parameter of angle
impacts on the hardness of the material. We used a
RV60-40 robot laser cell from Reis Robotics, which is a
leading technology company for robotics and system
integration. The articulated-arm robot series is the most
important robot kinematics for industrial use. As 6-axes
universal robots with high path speeds and large work
envelopes the RV-robots are especially suited for the
Materiali in tehnologije / Materials and technology 47 (2013) 1, 37–41 37
UDK 621.785.5:620.178 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)37(2013)
tough demands of path-related tasks. The design based
on FEM and CAD stands out due to its excellent static
and dynamic behaviour. Their robotic automation solu-
tions are used by all major application fields, such as
solar energy, foundry, welding and hardening. The Reis
Robotics group comprises three German subsidiaries and
eight international subsidiaries as well as representative
agencies in many countries. The laser beams have a
rectangular shape. We used 5 mm × 13 mm optics, which
means that with this optic we hardened a width of
approximately 13 mm. Robot lasers work continuously
with wavelength of 700–1000 nm. The maximum power
of a robot laser cell is 3000 W. However, we hardened
specimens with a 2000 W output power. The specimen
was material of DIN standard 1.2379. We hardened the
material at 2 mm/s using 1100 °C. There are different
and interesting problems regarding the robot laser
hardening of metals. The problem can be represented
geometrically, as seen in Figures 1 to 3.
Similar problems arise in the following situation. We
harden materials at an incidence angle of  	 90°. Fig-
ure 3 shows the situation where we changed the angles
in different directions. We see that the upper part of the
beam has a longer travel time than the lower part of the
beam. This means that the lower part of the hardened
piece is better than the upper. The workpiece will not be
evenly hardened and the final result of the laser harden-
ing will not be good.
To analyse the results we used the method of intelli-
gent system, i.e., a neural network. Neural networks are
model-less approximators; they are capable of perform-
ing an approximation – modelling operations regardless
of any relational knowledge of the nature of the
modelled problem. This relational knowledge is typically
represented by a set of equations describing the observed
variables and constants used to describe the system’s
dependencies. A common use of neural networks is
multi-dimensional function modelling13,14, i.e., the
re-creation of the system’s behaviour on the basis of a set
of known discrete points representing the various states
of the system. We used feedforward neural networks
with supervised training algorithms.
3 RESULTS
We are interested in the hardness of the robot
laser-hardened material as we change the incidence angle
of the laser beam on the substrate material. We have two
options. Firstly, we can change the angle with regard to
the direction of the laser beam. Here, we also have we
two options. In this situation we have conducted tests for
angles of   {15°, 30°, 45°, 60°, 75°, 90°} between the
right-hand side of the laser beam and the material surface
(Figure 1). However, we have conducted tests for angles
of   {15°, 30°, 45°, 60°, 75°, 90°} between the left-
hand side of laser beam and the material surface (Figure
2). This means that we made six tests for each option. In
these two options the width of the hardening is un-
changed. Second, we can change the angle with regard to
the perpendicular direction of the laser beam. We have
conducted tests for angles of   {15°, 30°, 45°, 60°,
75°, 90°}. In these options we have different widths of
hardening, because we change the angle with regard to
the perpendicular direction of the laser beam. The results
are presented in Figure 9. We varied the amounts of
power supplied to the laser beam when we made tests on
the tool steel 1.2379. In all the tables we present the
hardness before hardening, after hardening and the
average hardness after hardening.
M. BABI^ et al.: PROBLEMS ASSOCIATED WITH A ROBOT LASER CELL USED FOR HARDENING
38 Materiali in tehnologije / Materials and technology 47 (2013) 1, 37–41
Figure 3: Problem 3 of robot laser hardening: the lateral incidence
angle of the laser beam on the material surface and the beam move-
ment direction
Slika 3: Tretji primer laserskega kaljenja: spreminjanje lateralnega
vpadnega kota laserskega `arka glede na povr{ino materiala in smer
gibanja `arka
Figure 2: Problem 2 of the robot laser hardening: the variation of the
incidence angle   (0°, 90°) between the left-hand side of laser beam
and the material surface
Slika 2: Drugi primer laserskega kaljenja: spreminjanje vpadnega kota
  (0°, 90°) med levo stranjo laserskega `arka in povr{ino materiala
Figure 1: Problem 1 of the robot laser hardening: the variation of the
incidence angle   (0°, 90°) between the right-hand side of laser
beam and the material surface
Slika 1: 1. Primer laserskega kaljenja: spreminjanje vpadnega kota 
 (0°, 90°) med desno stranjo laserskega `arka in povr{ino materiala
3.1 Variation of the incidence angle with regard to the
direction of the laser beam
Again, we have two options (Figures 1 and 2). First,
we change the angle with regard to the direction of the
laser beam (the problem is presented in Figure 1). The
results of the measurements are shown in Table 1.
Table 1: Relationship between the angles and the hardness
Tabela 1: Povezava med kotom  in trdoto
/° Hardness afterhardening (HRc)
Average
hardness after
hardening
(HRc)
Hardness
before
hardening
(HRc)
15 62, 63, 62, 56 61.5 9
30 59, 59, 59, 63, 62 60.4 9
45 61, 61, 61 ,60, 61 60.8 9
60 61, 62, 61, 50, 56 58 9
75 61, 62, 61, 50, 61 59 9
90 61, 61, 61, 48, 62 58.6 9
All the data from Table 1 we analysed with the
neural network. Figure 4 shows the relationship between
the incidence angle and the hardness.
Second, we changed the angle with regard to the
direction of the laser beam (the problem is presented in
Figure 2). The results of the measurements are shown in
Table 2.
Table 2: Relationship between the angles and the hardness
Tabela 2: Povezava med kotom  in trdoto
/° Hardness afterhardening (HRc)
Average
hardness after
hardening
(HRc)
Hardness
before
hardening
(HRc)
15 61, 69, 52, 57, 56 55 9
30 54, 57, 58, 56, 56 56.2 9
45 51, 56, 56, 49, 54 53.2 9
60 52, 55, 54, 54 ,56 54.2 9
75 50, 55, 56, 23, 43 45.4 9
90 58, 59, 57, 60, 59 58.6 9
All the data from Table 2 we analysed with the
neural network. Figure 5 shows the relationship between
the incidence angles and the hardness.
3.2 Variation of the incidence angle with regard to the
perpendicular direction of the laser beam
In this case we changed the angle with regards to the
perpendicular direction of the laser beam (Figure 3). We
chose the same angles of   {15°, 30°, 45°, 60°, 75°,
M. BABI^ et al.: PROBLEMS ASSOCIATED WITH A ROBOT LASER CELL USED FOR HARDENING
Materiali in tehnologije / Materials and technology 47 (2013) 1, 37–41 39
Figure 6: Relationship between the angles and the hardness. The
modelling of the relationship was obtained using the four-layer neural
network (from Table 3).
Slika 6: Razmerje med vpadnimi koti in trdoto. Modeliranje razmerja
je bilo narejeno s {tiri-nivojskim nevronskim sistemom (podatki so v
tabeli 3).
Table 3: Relationship between the angles and the hardness
Tabela 3: Povezava med kotom  in trdoto
/° Hardness afterhardening (HRc)
Average
hardness after
hardening
(HRc)
Hardness
before
hardening
(HRc)
15 49, 48, 49, 61, 60 53.4 9
30 56, 57, 57, 63, 63 59.2 9
45 54, 56, 55, 64, 61 58 9
60 56, 57, 58, 63, 63 59.4 9
75 57, 57, 59, 63, 63 59.8 9
90 57, 60, 58, 58, 60 58.6 9
Figure 4: Relationship between the angles and the hardness. The
modelling of the relationship was obtained using the four-layer neural
network (from Table 1).
Slika 4: Razmerje med vpadnimi koti in trdoto. Modeliranje razmerja
je bilo narejeno s {tiri-nivojskim nevronskim sistemom (podatki so v
tabeli 1).
Figure 5: Relationship between the angles and the hardness. The
modelling of the relationship was obtained using the four-layer neural
network (from Table 2).
Slika 5: Razmerje med vpadnimi koti in trdoto. Modeliranje razmerja
je bilo narejeno s {tiri-nivojskim nevronskim sistemom (podatki so v
tabeli 2).
90°}. There is no need to consider two options since
those options are symmetrical.
All the data from Table 3 we analysed with the
neural network. Figure 6 shows the relationship between
the incidence angles and the hardness.
4 DISCUSSION
By using angular functions we can calculate the
width of the hardening at a certain angle. Here, the
following information is already known: the width of the
optics (d) and the angle () under which the hardening is
conducted. The hardening width is calculated. The width
of the beam optics represents one side of a right-angle
triangle, the angle () of hardening is the right-angle
triangle’s opposite side, which was marked with the
width of the optics (d). The beam hardening of the
workpiece is the hypotenuse of the right-angle triangle,
denoted by x. After delivery the sinus is
sin  = d/x , d  {5, 8, 13, 16, 23, 28} mm,
  (0°, 180°) (1)
By changing the angle  of the longitudinal harden-
ing of the workpiece, we can achieve different degrees of
hardness in the materials (section 3.1). By changing the
angle  with regard to the transverse hardening and the
different sizes of optics, we can achieve a different width
of hardening at a given time. Figure 7 shows that the
maximum power is used when the laser beam falls below
the minimum angle in our study, i.e., 15°. In this position
we achieve the area of hardening, but for that we require
more time and power. An imprecisely measured width
due to the hardening occurs as a small deviation from the
calculated width of the hardening with equation (1).
However, the measurement results are fairly accurate.
We are interested in conditions that return better results.
In our case the most favourable solution to the problem
is the hardness of the material.
Figure 8 shows the relationship between the hardness
and the angle of hardening. In this graph we can see a
comparison of the three different types of robot laser
hardening by changing the angle of the laser beam.
Figure 9 presents the relationship between the width
and the angle of hardening with regards to the perpendi-
cular direction of the laser beam. We can see that by
increasing the angle, the width of the hardening
decreases, which we can prove with equation (1).
We analysed the graph with two different methods.
First, we used linear regression. Second, we used the
modelling of the relationship that was obtained by the
four-layer neural network. For Figures 4, 5 and 6 we
calculated the correlation coefficient, which represents
the size of the linear connection of the hardness and the
fractal dimension. The correlation coefficient R for graph
1 is –0.8324, for graph 3, –0.88263 and for graph 5,
0.65799. We can see that the correlation coefficients are
not similar. Because the correlation coefficients are not
0, the variable hardness and the angles of hardening are
correlated. Smaller values of the angles tend to be linked
to the hardness values, which tell us that there is a
negative correlation coefficient. This is presented in
Figures 4 and 6. But in Figure 5 we have a different
situation – a positive correlation coefficient. The purpose
of this work has been to study how the angles of the
robot laser cell impact on the hardness of the specimens.
The presented problem could be solved in order to
modify the laser beam’s intensity across the width. This
means that the first laser beam is divided into several
parts. Then each part of the laser beam is divided into the
M. BABI^ et al.: PROBLEMS ASSOCIATED WITH A ROBOT LASER CELL USED FOR HARDENING
40 Materiali in tehnologije / Materials and technology 47 (2013) 1, 37–41
Figure 9: Relationship between the width and the angle of hardening
with regards to the perpendicular direction of the laser beam
Slika 9: Razmerje med {irino kaljenja in kotom kaljenja glede na
pravokotno smer kaljenja laserskega `arka
Figure 7: Relationship between the power and the angles of hardening
Slika 7: Razmerje med mo~jo in kotom, pod katerim kalimo
Figure 8: Relationship between the hardness and of the angle of
hardening
Slika 8: Razmerje med trdoto in kotom, pod katerim kalimo
specified strength. The part of the laser beam that made
the longest journey gave most of the power to that part of
the beam, and the part that had the shortest path to the
device, gave the smallest amount of laser beam intensity.
5 CONCLUSION AND FUTURE WORK
Robot laser hardening is very useful in the auto-
motive (e.g., machine parts for transmission shafts,
axles, running surface, torsion springs, gears), military
and aerospace industries. The process has several advan-
tages over conventional induction hardening. However,
even in the robot laser-hardening process there are
problems, as described in this paper. Thus, we still have
enough unsolved problems in robot laser hardening.
Robot laser cells have several parameters that affect the
final result of the hardening. These laser parameters are
the power, the energy density, the focal distance, the
energy density at the focus, the focal position, the
temperature and the speed of germination. In the future
we want to explore how different angles of the laser
beam in the hardening process affect the hardened
patterns in:
• dual robot laser-beam hardening (laser beam is
divided into two parts),
• pixel robot laser hardening,
• robot laser hardening with changes to the velocity
and the temperature of the laser beam.
6 REFERENCES
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M. BABI^ et al.: PROBLEMS ASSOCIATED WITH A ROBOT LASER CELL USED FOR HARDENING
Materiali in tehnologije / Materials and technology 47 (2013) 1, 37–41 41

DJ. VUKELIC et al.: BURNISHING PROCESS BASED ON THE OPTIMAL DEPTH OF A WORKPIECE PENETRATION
A BURNISHING PROCESS BASED ON THE OPTIMAL
DEPTH OF WORKPIECE PENETRATION
POSTOPEK GLADILNEGA VALJANJA NA OSNOVI OPTIMALNE
GLOBINE PRODIRANJA V OBDELOVANEC
Djordje Vukelic1, Dragomir Miljanic2, Sasa Randjelovic3, Igor Budak1, Dragan
Dzunic3, Milan Eric3, Marko Pantic3
1University of Novi Sad, Faculty of Technical Sciences, Department for Production Engineering, Trg Dositeja Obradovica 6, 21000 Novi Sad,
Serbia
2Metalik DOO, Trebje{ka 6/26, 81400 Niksic, Montenegro
3University of Kragujevac, Faculty of Engineering, Department for Production Engineering, Sestre Janjic 6, 34000 Kragujevac, Serbia
vukelic@uns.ac.rs
Prejem rokopisa – received: 2012-06-26; sprejem za objavo – accepted for publication: 2012-08-27
This paper deals with the problem of defining the trajectory of a stiff burnishing tool that would be optimal from the point of
view of surface quality. The basic goal of this work is to gain an insight into the very process from the microscopic aspect, with
the primary focus on material flow and roughness variations. Based on theoretical considerations, we planned an experiment
with the aim to verify the initial hypotheses about the analysis of roughness change and a determination of the optimal depth of
the workpiece penetration. Through matching and the superposition of surface profiles formed at various contact pressures, i.e.,
various burnishing forces and various penetration depths of the burnishing ball into the profile roughness, the phenomenon of
roughness change was explained. Theoretical assumptions related to a determination of the optimum tool trajectory have largely
been confirmed from the point of view of surface quality. The balls within the stiff tool system, which follow a predetermined
depth of penetration into the roughness profile, very likely provide optimum surface quality, regardless of the initial machining.
Based on experimental results, it is highly possible that the depth of the penetration of tool (burnishing ball) should equal the
roughness profile height of the previously machined surface. The analysis of results obtained by the measurement of the surface
roughness and the super-positioning of the profiles obtained by burnishing with various burnishing forces, significantly
contributed to the explanation of the roughness peaks’ deformation phenomenon. The proposed burnishing method could be of
special importance in the burnishing of roughly machined surfaces, where Rp reaches high values. Investigations presented in
this paper open a number of new directions, such as the testing of a stiff tool system with various workpiece materials and
burnishing regimes, with different surface roughnesses as the result of the initial machining. We believe that the proposed model
can significantly improve our present knowledge of the burnishing process.
Keywords: burnishing, surface quality, stiff tool system, profile peaks, profile valleys
^lanek obravnava problem dolo~anja optimalne trajektorije togega gladilnega orodja glede na kvaliteto povr{ine. Osnovni cilj
tega dela je vpogled v postopek obdelave na mikroskopskem nivoju, osredinjen na te~enje materiala in spreminjanje hrapavosti
povr{ine. Na osnovi teoreti~nega raziskovanja so avtorji pripravili preizkus za potrditev za~etne hipoteze o spremembi
hrapavosti in dolo~anju optimalne globine prodiranja v obdelovanec. Z usklajevanjem in predpostavljanjem profila povr{ine, ki
nastane pri razli~nih pritiskih, to je z razli~nimi silami pri gladilnem valjanju in za razli~ne globine prodiranja gladilnih kroglic
na profil hrapavosti, je bil pojasnjen pojav spreminjanja hrapavosti. Teoreti~ne predpostavke za dolo~anje optimalne poti orodja
so bile potrjene s stali{~a kvalitete povr{ine. Kroglice v togem orodju, ki sledijo vnaprej dolo~enemu prodiranju v profil
hrapavosti, zelo verjetno zagotavljajo optimalno kvaliteto povr{ine, ne glede na predhodno obdelavo povr{ine. Na osnovi
eksperimentalnih rezultatov je najbolj verjetno, da naj bi bila globina prodiranja orodja (gladilne kroglice) enaka vi{ini profila
hrapavosti predhodno obdelane povr{ine. Analiza rezultatov, dobljenih z meritvami hrapavosti povr{ine, dobljene pri gladilnem
valjanju z razli~nimi silami gladilnega valjanja, je pomembno prispevala k razjasnitvi pojava hrapavosti. Predlagana metoda
glajenja je lahko pomembna pri gladilnem valjanju grobo obdelanih povr{in, kjer Rp dosega velike vrednosti. Raziskave,
predstavljene v tem prispevku, odpirajo {tevilne nove smeri, kot so preizku{anje sistemov togih orodij z razli~nimi materiali
obdelovancev in re`imi gladilnega valjanja pri razli~nih hrapavostih iz predhodne obdelave. Avtorji menijo, da predlagani model
lahko pomembno prispeva k sedanjemu poznanju postopkov glajenja povr{in.
Klju~ne besede: gladilno valjanje, kvaliteta povr{ine, sistem togega orodja, profil vrhov, profil dolin
1 INTRODUCTION
In manufacturing industry, surface roughness plays a
vital role in product performance. Regardless of whether
a product is shaped with or without chip removal, there
are a number of factors that influence workpiece surface
roughness, such as workpiece characteristics (physical
and mechanical properties, chemical composition,
micro-geometry, macro-geometry, etc.), machining
equipment (stiffness, kinematics, sensitivity to heat
transfers, etc.), tool (material, shape, surface quality,
rigidity, wear, etc.), cooling and lubricating agent (che-
mical composition, viscosity, method of application,
etc.), as well as the characteristics of the machining or
forming process (strain, strain rates, stress distribution
inside a workpiece, heat generation, etc.).1–3 In addition
to numerous machining processes (milling, grinding,
polishing, honing, lapping) which contribute to a lower
surface roughness,4 there is also the burnishing process.
Ball/roller burnishing is a cold-finishing process
without chip removal that plastically forms the surface
layer of a workpiece. The purpose of this finishing pro-
cess is not to achieve a dimensional accuracy but a
surface quality with appropriate roughness,5–10 micro-
Materiali in tehnologije / Materials and technology 47 (2013) 1, 43–51 43
UDK 621.7:621.7.019 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)43(2013)
hardness,11–13 wear and corrosion resistance,14–16 fatigue-
and tensile strength.14,15,17 Furthermore, residual tensile
surface stresses, which are the result of previous
machining (turning, milling, etc.), are transformed by
burnishing into compressive stress, thus improving
several mechanical properties.17–20 The penetration depth
of compressive stresses as well as the thickness of
hardened surface layer depend on the workpiece material
and the applied loads. Compressive stresses decrease
from the workpiece surface to the interior, while the
penetration depth can reach up to 1 mm, depending on
the workpiece material and the loads. Such a treatment
achieves a smooth surface with a hard layer on the
workpiece surface, which is the result of the deformation
strengthening caused by intensive plastic deformation of
the surface layer. It also diminishes the surface defects
(as the result of surface plastic deformation) and modi-
fies the surface microstructure.21–23 Among the advanta-
ges of ball/roller burnishing are flexibility, cost-effective-
ness and the possibility to use simple machining
equipment.
In burnishing, the working part of the tool (hard ball
or roller) is rolled over workpiece surface. As the result
of rolling over the surface, high contact (Hertz) pressures
occur, which overstep the yield stress, leading to plastic
forming of the surface layer. Roughness peaks are
deformed, and through plastic flow we begin to fill the
roughness valleys between them, evening out the texture
of the surface roughness.
Burnishing can be used on workpieces of various
materials, such as the bronze,24 aluminum5,17,18,25–27 and
various steels.7,9,10,16,20–23 Brinksmeier et al.21 showed that
the burnishing of workpieces with a large content of
unstable austenite can lead to a deformation-induced
martensitic transformation. This proved that it is possible
to include martensitic transformation in a mechanical
surface treatment without introducing additional thermal
processes. Besides using burnishing on various material
workpieces, this process can also be applied on work-
pieces of various geometries, which makes it practical
for treating outer and inner cylindrical surfaces as well as
small- and large-area flat surfaces.
The tools that are used in this process feature a ball
or a roller as the working element whose design should
provide smooth rolling over the workpiece surface,
without sliding and the occurrence of adhesive bonding
during work. There are various design solutions that
provide the free rolling of the working element, using
backing-up balls17,24 or roller bearings, as is the case in
this work. The force with which the ball presses against
the workpiece surface is most often regulated by cali-
brated springs,24–27 although different solutions also exist,
featuring pressurized fluid,21–23,28 flexible tool holders,29
etc. In addition, there are tool carriers that are specially
designed for application on large-area flat surfaces,
which reduce the processing time while being able to
accommodate several simple burnishing tools.17,30
Numerous researchers have studied this process,
investigating the influence of ball/roller material and
dimensions, workpiece material, tool geometry, and
process parameters, i.e., the burnishing speed, feed,
pressure force, and number of passes. There are a num-
ber of papers that investigate the optimization of process
parameters.10,12,18,19,24–26 However, achieving an optimum
set of burnishing parameters for a specific workpiece
material requires a large number of experiments. In order
to limit that number, Response Surface Methodology
(RSM)17 and the Taguchi method24 are employed. For
this reason it is very important to develop an appropriate
mathematical model that predicts the surface-quality
parameters with the required accuracy.17,30,31
Furthermore, there are a number of papers that
compare burnishing to alternative processes that provide
similar output results.9,14,32–34 Also, some authors have
investigated the surface quality achieved by burnishing,
considering previous machining. For example, Bouzid et
al.35 have established that the best burnishing results are
achieved with grinding as the initial machining. There
are some authors who used finite-element analyses
(FEA) to model the burnishing process, achieving satis-
factory results.34,36
The papers analyzed in this section allow us to
conclude that the burnishing force is the most influential
parameter regarding surface roughness, hardness, thick-
ness of the hardened layer, as well as the likelihood of
surface damage (layer peeling) during processing.
Gharby et al.30 have investigated the burnishing of AISI
1010 steel plates and established a limit of 400 N burn-
ishing force, stating that above this limit the workpiece
surface layers begin to peel off. The same author17 con-
ducted experiments on 1050A aluminium and establi-
shed an optimal burnishing force of 115 N, pointing out
that a combination of large burnishing forces and large
feeds deteriorates the surface roughness. Most of the
investigations dealing with this subject varied the
burnishing force within the range 0–400 N. In this paper
we focused on the influence of the burnishing force on
the surface roughness, while the force was varied within
the range 0–450 N.
The optimal forces are determined for specific
workpiece material types and their characteristics. From
the reviewed literature it can be concluded that the
burnishing force is used as an optimization parameter.
However, considering the very burnishing process (sur-
face strengthening and surface quality), it is better to
consider the magnitude of the contact pressure between
the burnishing ball and the workpiece surface as an
optimization parameter. A numerical determination of
the magnitude and the distribution of contact pressures
requires data on the ball/roller radius, the radius of work-
piece surface curvature (in the case that the burnishing is
performed on a lathe, it is the billet radius, otherwise, for
flat surfaces, an infinite radius of curvature is assumed),
and the data on the module of elasticity and the Poisson
DJ. VUKELIC et al.: BURNISHING PROCESS BASED ON THE OPTIMAL DEPTH OF A WORKPIECE PENETRATION
44 Materiali in tehnologije / Materials and technology 47 (2013) 1, 43–51
coefficient for the workpiece and the ball/roller material.
However, most often the available literature on the
subject of burnishing does not provide these data.
Bearing in mind that the profiles of the machined
surfaces are inherently stochastic, it is extremely difficult
to determine the values of the contact pressures with the
required certainty. All this considered, most investigators
of burnishing process choose to consider the burnishing
force as the parameter of influence, rather than the more
relevant contact pressure.
Considering the large number of material types in
industrial applications and the wide range of their
properties, the broader application of the burnishing
process would require an extensive database with the
optimal contact pressures for all types of materials. This
is especially true if one considers that the optimal con-
tact pressure is not only related to the workpiece material
characteristics but also to the process parameters (feed,
number of passes, initial surface quality, etc.).
For that reason, the focus of this investigation is
placed on establishing the depth to which the stiff
burnishing tool penetrates the roughness valleys, i.e., the
optimal value for this depth which, considering the force
magnitude and other process parameters, provides near-
optimal surface quality.
The available references mostly agree on ranking the
influential parameters (burnishing force, feed, number of
passes and other factors) based on their relative impact
on the surface quality. However, an optimal surface
roughness can be achieved at various levels of burnishing
forces. Therefore, optimal burnishing forces depend on
the workpiece material properties and the ball/roller dia-
meter, the burnishing parameters and the initial surface
quality. It is worth noting that the investigations pub-
lished so far have not looked into the phenomenon of
roughness peaks, which leaves space for an investigation
into the changes that take place in the surface roughness
during the burnishing.
The initial hypothesis in this investigation supposes
that it is possible to define a near-optimal penetration
depth of the burnishing tool into the workpiece rough-
ness profile, which would result in a near-optimal
surface quality, depending on the magnitude of the
burnishing force and other burnishing parameters.
Within this investigation, experiments were con-
ducted with the aim to verify the initial hypotheses
pertaining to an analysis of the surface topography
changes and a determination of the near-optimal tool
trajectory. By monitoring the changes that take place in
the roughness profiles at various depths of tool pene-
tration into the roughness valleys, an optimal penetration
depth was determined that corresponds to the minimal
surface roughness (Ra). In these experiments a specially
designed stiff tool system was employed. It should be
noted that the phenomenon of workpiece surface
strengthening was not within the scope of this investi-
gation. As previously discussed, this study is dealing
primarily with the problem of workpiece surface quality
in burnishing, and also looks into the problem of
optimizing the surface quality as a function of tool (ball)
penetration depth into the workpiece surface roughness
profile. Finally, experimental investigations were
performed on a tempered steel.
2 THEORETICAL CONSIDERATIONS
In this section theoretical considerations of the burn-
ishing process are presented as the rolling of a ball over
the roughness peaks on the previously machined work-
piece (Figure 1). The burnishing is performed with a
tool system of high stiffness, consisting of a ball with a
defined radius R, mounted on a carrier that provides free
rolling.
During the burnishing process the burnishing ball is
also exposed to lateral forces. The theoretical model
proposed in this paper does not consider these forces for
the following reasons:
• The rolling resistance of the ball during relative
motion over the workpiece surface represents the
lateral force. This force can be neglected, considering
the 0.001 order of magnitude of the rolling resistance
coefficient, as well as the relatively high ratio
between the normal (vertical) load and the rolling
resistance.
• The remaining lateral forces can also be neglected if
one considers the exclusively large ratios between the
ball radius and the maximum surface profile peak
height, which is often the case in ball burnishing. In
such conditions the angle between the lines drawn
from the ball centre, connecting the two points of
ball/profile contact, is very small, which means that
the vertical load on the burnishing ball is predo-
minant, as illustrated in Figure 2.
The considered model pertains to static loads, ne-
glecting the dynamic force component. This is justified
in cases when burnishing is performed at relatively low
DJ. VUKELIC et al.: BURNISHING PROCESS BASED ON THE OPTIMAL DEPTH OF A WORKPIECE PENETRATION
Materiali in tehnologije / Materials and technology 47 (2013) 1, 43–51 45
Figure 1: Theoretical analysis of the burnishing process based on a
stiff tool system
Slika 1: Teoreti~na analiza postopka gladilnega valjanja, zasnovanega
na sistemu togega orodja
speeds, which is the case in this experimental investi-
gation. The authors maintain that, given the discussed
restrictions, the developed model can prove useful in
clarifying the phenomenon of workpiece surface profile
variations at various penetration depths of a stiff tool
system.
Depending on the ball radius, the roughness pitch and
the roughness form, the contact between the ball and the
workpiece is established over one or more roughness
peaks (neighbourhood of points M1, M2, ... Mn). In the
course of the ball’s penetration into the roughness
profile, over the planes parallel to OO1 axis, the area of
the contact surface between the ball and the roughness
peaks increases. For a predefined tool displacement, yi,
(tool movement into roughness profile, from a reference
point) there follows: F(y1) 
 F(y2) 
 ... 
 F(yi), where
F(yi) is the force corresponding to the yi coordinate.
If the goal is to achieve the maximum surface quality,
it is essential to establish the depth, yopt, which represents
the optimal tool penetration into the roughness profile.
One of the possible models of material flow in the
contact zone during burnishing is presented in Figure 2.
Due to the high contact pressures that exceed the yield
stress, roughness peaks begin to flow and gradually fill
the valleys. It is well known that surface strengthening
should be the most pronounced in the layers closer to the
profile peaks. In other words, the intensity of the surface
strengthening drops towards the profile valleys. This
results in the fewer hard material layers being suppressed
towards the profile valleys by the layers of greater
hardness.
As the tool penetration depth, y, increases, the
roughness valleys are filled. We supposed in this work
that the tool penetration depth y should be determined
based on the recorded surface roughness profile prior to
burnishing, according to:
y  Rp (1)
where: Rp is the maximum height of the surface rough-
ness profile before burnishing. The resulting equality is:
Δ ΔP Pi
i
i n
i
i
I n
=
=
=
=
∑ ∑≅ ′
1 1
(2)
where Pi is the surface area of i-th peak relative to the
mean profile line, Δ ′Pi is the surface area of i-th valley
relative to mean profile line, that is:
f x x f x x
L
f x
L
f x
( ) ( )
( ) ( )
d d
0 0 0 0
∫ ∫
< >
≅ (3)
Evidently, the roughness profile curve of the pre-
viously machined surface (prior to burnishing) does not
have an analytical function, f(x). However, based on the
numerical data obtained by modern devices for surface-
roughness measurements it is possible to check whether
the condition of approximate equality between the
surface areas of the profile peaks and the valleys pro-
cessed by burnishing. This should require dedicated
software, which is one of the goals for future investi-
gation. The assumptions made in this work claim that the
optimal surface quality in burnishing of previously
machined surfaces is achieved for a tool penetration
depth that equals the maximum peak height, Rp, which
roughly conforms with the condition of equality between
the surface areas of the roughness peaks and valleys.
These assumptions are based on the following facts:
• The material which flows along the surface rough-
ness peaks should be allowed some space to deposit
(Figure 2). The condition of equality between the
surface areas of the peaks and valleys theoretically
allows a simultaneous decrease of the roughness
peaks, Rp, and the roughness valleys, Rv.
• When the tool (ball) is displaced for yopt relative to
some reference location into the roughness profile,
the material will flow from the peaks, filling the
valleys and leaving the profile without additional
peaks.
• Presumably, the flow of material in the surface peaks
predominantly occurs through the widening and
narrowing of peak profiles. This claim is supported
by assumption that, due to the large ratio of the ball
radius (R) and the relatively small feeds used in the
initial machining, the resulting contact force that
compresses the profile peaks probably assumes a
direction normal to the mean profile line (Figure 2).
Thus, regardless of the stochastic nature of the initial
surface roughness profile, it is realistic to expect that
the proposed burnishing method, based on a stiff tool
system, will yield a better surface quality compared
to the burnishing tools that operate with a constant
force (provided by spring mechanisms). Due to
variations in the material resistance, such tools
oscillate considerably, which represents the source of
the additional roughness and profile waviness.
To confirm the theoretical assumptions, Figure 3
presents the results of preliminary investigations that
DJ. VUKELIC et al.: BURNISHING PROCESS BASED ON THE OPTIMAL DEPTH OF A WORKPIECE PENETRATION
46 Materiali in tehnologije / Materials and technology 47 (2013) 1, 43–51
Figure 2: The contact zone during ball burnishing and the supposed
directions of material flow
Slika 2: Kontaktno podro~je med procesom gladilnega valjanja in
predvidena smer toka materiala
involved a stiff tool system with the ball penetration
depth limited to the maximum peak height. The same
scale was used to represent the superimposed surface
roughness profiles before and after the completion of the
burnishing process. The finished surface of very small
roughness, Ra = 0.061 μm (Figure 3), was obtained with
a 5 μm penetration depth, which approximately corres-
ponds to a maximum profile peak height, Rp = 4.3 μm, of
the same surface prior to burnishing.
3 EXPERIMENTAL INVESTIGATION
The experimental investigation included burnishing
of the cylindrical surface. The burnishing was performed
on a specimen (billet) of tempered steel, 36CrNiMo4,
hardness 42 HRC. The specimen dimensions were d = 50
mm and L = 400 mm. The burnishing was performed on
a universal lathe (Figure 4) with a specially designed
stiff tool system, a technical drawing of which is shown
in Figure 5. The system was designed with three roller
bearings that support the burnishing ball at the areas
defined by the neighbourhoods of three points, ensuring
that the ball rolls in a plane. The burnishing ball featured
a 7 mm diameter and was made of steel, A 295 52100
(USA/ASTM). The ball hardness was 65 HRC, while its
surface roughness was equal to Ra = 0.02 μm.
Considering the burnishing ball, it should be noted
that, due to the limited load capacity of the small roller
bearings (Figure 5), the maximum allowed force on the
ball is limited to F = 600 N. Thus, it is important to
emphasize that it was not possible to perform burnishing
with penetration depths above the maximum peak height
on the previously machined surface.
The burnishing was performed in a single pass, with
a feed rate f = 0.05 mm/r and the number of revolutions n
= 45 r/min. A small number of revolutions was selected
in order to eliminate the negative thermal and dynamic
effects. Considering the initial hypotheses and the goal
of this investigation, the tendency was towards avoiding
intensive heating of the burnishing ball, as well as any
significant vibrations of the tool system and the work-
piece. All this allowed a deeper insight into the very
process, as well as a more realistic comparison between
the proposed theoretical model and the experimental
results. Loads were applied using a cross slide on a
universal lathe. Mounted on the lathe was a dynamo-
meter, Kistler 9265A, which supported the tool system.
The level of loading was determined by this dynamo-
meter prior to the initiation of the burnishing process.
Also, during the burnishing process, the Kistler 9265A
dynamometer was used for continuous monitoring of the
burnishing force used in the process. Both workpiece
sections were previously lathed to different surface
roughness and subsequently burnished with six different
burnishing forces. Upon completion of the burnishing
process, a Talisurf 6 was used to measure the burnished
surface roughness parameters (Ra and Rp). Measurements
were taken in both sections (initial machining 1, and
initial machining 2), along the contour lines on the
workpiece, in three radial directions oriented at 120°
relative to the axis of the workpiece. Figure 6 shows a
DJ. VUKELIC et al.: BURNISHING PROCESS BASED ON THE OPTIMAL DEPTH OF A WORKPIECE PENETRATION
Materiali in tehnologije / Materials and technology 47 (2013) 1, 43–51 47
Figure 3: Superimposed recordings of the surface profiles
Slika 3: Prekrivanje meritev hrapavosti povr{ine
Figure 6: Schematic drawing of the fields in which the burnishing
forces were measured, and the locations of the contour lines along
which the roughness measurements were taken
Slika 6: Shemati~en prikaz polj, na katerih so bile izmerjene sile med
gladilnim valjanjem, in polo`aj linij, vzdol` katerih je bila izmerjena
hrapavost
Figure 5: Technical drawing of burnishing tool system
Slika 5: Tehni~na risba orodja
Figure 4: Photograph of burnishing tool in operation
Slika 4: Posnetek orodja med gladilnim valjanjem
drawing with the marked sections used to measure the
burnishing forces, showing the locations of the contour
lines along which the roughness measurements were
taken. The referential length for taking the roughness
measurements equalled 3 × 0.8 = 2.4 mm. Burnishing
was performed on two previously machined surfaces
with different roughness (initial machining 1, and initial
machining 2 – Figure 6). The first initial machining was
performed with the following parameters: feed rate f =
0.5 mm/r, depth of cut d = 1 mm, and number of
revolution n = 710 r/min. The second initial machining
was performed with f = 0.1 mm/r, while the rest of
parameters were identical to the first initial machining.
4 RESULTS
The experimental results encompass the data on the
burnishing forces that were used in the process as well as
the data on the roughness parameters of the burnished
surfaces.
Table 1 presents the mean values of the burnishing
force, the standard deviations, the minimum and maxi-
mum burnishing forces generated during the process, and
the surface-roughness parameters (arithmetic mean
surface roughness, Ra, maximum peak height, Rp, and
maximum valley height, Rv). Based on the maximum
peak height on the profile of the previously machined
surfaces, Rp(init), as well as the maximum peak height,
Rp(F), measured upon completion of the burnishing
process in a particular section and along a particular
contour line, the ball penetration depths were calculated
according to:
y = Rp(init) – Rp(F) (4)
Thus, the values of y represent the real ball-penetra-
tion depths into the roughness profile. These values were
measured after the completion of the burnishing process,
which eliminated the errors due to the elastic defor-
mations of the workpiece, tool system, and support.
We believe that measurements of the maximum peak
heights along the three profile lines helped mitigate the
negative influence of technological errors (deviation
from circularity during the workpiece rotation and the
workpiece elasticity). This influence is manifested
through a more or less pronounced dispersion of the
burnishing force within particular sections. It is not only
a logical assumption, but was actually observed in the
experiment, that each variation in the burnishing force
within a burnishing section, resulted in a roughness
DJ. VUKELIC et al.: BURNISHING PROCESS BASED ON THE OPTIMAL DEPTH OF A WORKPIECE PENETRATION
48 Materiali in tehnologije / Materials and technology 47 (2013) 1, 43–51
Table 1: Results of the experimental investigation
Tabela 1: Rezultati preizkusov
Forces and roughness
parameters
Initial machining 1 Initial machining 2
Ra [μm] Rp [μm] Rv [μm] Ra [μm] Rp [μm] Rv [μm]
7.3 ÷ 7.53 15 ÷ 16 –13 ÷ –14 4.5 ÷ 5.4 14 ÷ 17 –16 ÷ –19
Number of measurement
– force variation –
Number of measurement
– force variation –
1 2 3 4 5 6 1 2 3 4 5 6
Forces and
dispersion
of forces
Fysr [N] 47.98 67.41 32.30 55.53 213.28 320.70 81.75 166.08 194.86 322.37 305.21 444.28
Fn [N] 25.21 36.32 26.05 40.51 103.20 103.46 34.71 52.74 39.88 47.98 38.31 56.77
Fymin [N] 8.14 16.74 0 3.16 74.32 125.57 21.31 83.00 113.49 216.88 220.71 318.03
Fymax [N] 111.68 169.94 103.88 169.23 432.60 577.61 182.97 289.21 287.77 428.39 389.76 569.26
P
ro
fi
le
li
ne
on
w
or
kp
ie
ce
AB
Ra [μm] 6.6 6.9 6.6 6.9 4.1 1.04 0.84 1.33 1.05 0.49 0.85 0.774
Rp [μm] 13 15 13 15 9 1.9 3.6 3.5 2.5 1.3 2 2.27
y [μm] 3 1 3 1 7 14.1 13.7 13.8 14.8 16 15.3 15.03
CD
Ra [μm] 5.7 5.06 3.69 4.31 0.57 0.84 1.82 0.627 0.51 0.41 0.5 0.52
Rp [μm] 12 8.8 7.8 8.6 1.3 1.8 5 2.02 1.4 1.5 1.2 1.3
y [μm] 4 7.2 8.2 7.4 14.7 14.2 12 14.89 15.6 15.5 15.8 15.7
EF
Ra [μm] 6.59 6.85 6.1 6.35 2.75 1.2 1.89 0.88 0.76 0.54 1.01 0.55
Rp [μm] 13.1 14.3 12.3 12.6 4.3 2.3 6.1 2.4 2.4 1.5 2 2.8
y [μm] 1.9 0.7 2.7 2.4 10.7 12.7 7.9 11.6 11.6 12.5 12 11.2
Figure 7: Typical recording of the burnishing force signal
Slika 7: Zna~ilen zapis signala sile pri gladilnem valjanju
deviation (see the measurement results for Rp, along
profile lines AB, CD, and EF – Table 1).
A typical example of a burnishing-force signal
recording is shown in Figure 7. In the burnishing force
signal pattern one clearly distinguishes near-periodic
variations within a single revolution. Namely, the force
variations repeat every 1.33 s, which approximately
corresponds to n = 45 r/min. This indicates that
variations in the burnishing-force magnitude, and,
consequently, variations in the surface quality, occur due
to errors in the number of workpiece revolutions degrees
of its elastic deformations in particular sections, which
inevitably leads to variations in the contact pressure
between the tool and the workpiece surface. Since the
experiment was performed with an extremely stiff tool
system, minor errors in the circularity of rotation and
small deviations in the workpiece elastic deformations
were sufficient to generate a significant dispersion of the
burnishing force.
Through the processing of measurement signals
acquired with the Talisurf 6 profilometer, the profiles of
treated surfaces were superimposed for a visual analysis
of the changes in the surface roughness due to a varied
burnishing force. Figures 8 and 9 show the superim-
posed profiles generated by applying different burnishing
forces on two surfaces previously machined to various
surface-roughness values. These diagrams (Figures 8
and 9) pertain to the first initial machining, with the
surface roughness ranging within the Ra = 7.3–7.53 μm
interval. From the diagrams in Figures 8 and 9 we can
see that the increase of the burnishing force gradually
leads to a significant reduction of the surface roughness.
It is interesting to note that for both initial
machinings, the ball penetration depths that are close to
the maximum peak height of the previously machined
surface correspond to lowest Ra values (Table 1).
Based on the data from Table 1 the diagram in
Figure 10 shows, for both initial machinings, the
dependence between the roughness, Ra, and y, which
represents the ball penetration depth. For both initial
machinings, the diagram shows an obvious trend of the
surface roughness (Ra) variations as a function of the ball
penetration depth. Evidently, for both initial machinings,
the surface roughness drops sharply until it reaches the
maximum peak height of the previously machined
surface roughness profiles, i.e., y = Rp = 14–17 μm. After
that point, the surface roughness decreases very slightly.
5 DISCUSSION
Numerous authors have focused their investigations
on defining the optimal burnishing force. As previously
discussed, the widespread and efficient application of
burnishing should require an extensive database
containing optimal values for the burnishing forces that
are a function of a number of parameters (workpiece
material, burnishing feed, number of passes, and quality
of surface prior to burnishing). In this work the focus is
placed on defining the appropriate penetration depth of a
stiff burnishing tool system, i.e., the depth that will
DJ. VUKELIC et al.: BURNISHING PROCESS BASED ON THE OPTIMAL DEPTH OF A WORKPIECE PENETRATION
Materiali in tehnologije / Materials and technology 47 (2013) 1, 43–51 49
Figure 9: Superimposed profiles generated by burnishing after the
first initial machining, with F = 55 N, F = 213 N and F = 320 N
Slika 9: Prekrivanje profilov, dobljenih pri glajenju s silo F = 55 N, F
= 213 N in F = 320 N, na predhodno obdelani povr{ini
Figure 8: Superimposed profiles generated by burnishing after the
first initial machining, with F = 32 N, F = 47 N and F = 67 N
Slika 8: Prekrivanje profilov, dobljenih pri glajenju s silo F = 32 N, F
= 47 N in F = 67 N, na predhodno obdelani povr{ini
Figure 10: Dependence of the surface roughness Ra, on the depth of
the ball penetration into the roughness profile
Slika 10: Odvisnost hrapavosti povr{ine Ra od globine prodiranja
kroglic v profil hrapavosti
provide an optimal surface roughness, regardless of the
force magnitude and the other parameters of the
burnishing process. Discussed were the basic theoretical
considerations according to which the rolling of a ball
within a stiff tool system, and the penetration into the
roughness profile up to the predefined depth, are very
likely to provide near-minimal surface roughness, where
equalities (1) and (3) apply.
The basic assumptions in this paper were largely con-
firmed by the results of the experimental investigation.
The diagram in Figure 10 clearly indicates that the
displacement of the burnishing ball into the roughness
profile, y, which is close to the maximum peak height of
the previously machined surface-roughness profiles,
provides the lowest surface roughness, Ra. For the first
initial machining (Table 1) the maximum peak height is
in the range Rp = 15–16 μm. From the diagram in Figure
10 it is evident that y = 14.7 μm (measured along the
profile line CD – Table 1), which corresponds to Rp =
15–16 μm in the case of first initial machining, results in
the near-minimal surface roughness, Ra = 0.57 μm (see
Figure 10 and Table 1).
Considering the second initial machining (Table 1),
the maximum peak height is in the range Rp = 14–17 μm.
The diagram shown in Figure 10 clearly indicates that y
= 15.5 μm (measured along the profile line CD – Table
1), which corresponds to the maximum peak height, Rp =
14–17 μm, results in a near-minimal surface roughness,
Ra = 0.41 μm. It should be noted that in the case of the
second initial machining (Table 1) on the profile lines
AB and CD there are a number of y values that are
approximately equal to Rp = 14–17 μm, and which
correspond to a near-minimal surface roughness, as
shown in Figure 10. Based on the previous discussion,
and the diagram in Figure 10, it is logical to suppose
that the optimum ball-penetration depth equals the
maximum peak height, Rp, of the previously machined
surface. This claim is also supported by the trend of the
change of surface roughness Ra depending on the ball-
penetration depth (Figure 10). Based on the diagram
shown in Figure 10 it is obvious that, for both initial
machinings, the surface roughness, Ra, drops signifi-
cantly until the ball-penetration depth reaches the
maximum peak height achieved with the previous
machining. After that, the decrease of Ra is significantly
milder. With the first initial machining (higher surface
roughness) the percentage change of the surface
roughness, Ra, is significantly higher. Thus, regardless of
the surface quality obtained with the previous machin-
ing, the values of y that are close to the maximum peak
height, result in the lowest values of surface roughness.
This fact can be valuable when applying burnishing on
surfaces with a rough previous machining.
Therefore, a high surface quality can be achieved
with a tool displacement that corresponds to the maxi-
mum peak height, Rp, of the previously machined
workpiece surface. It is evident that the material flowing
from the profile peaks should be allowed allocation
space (Figure 2). Besides, the condition of equality of
the cross-section surface areas of the roughness peaks
and valleys theoretically allows the simultaneous
decrease of the peak, Rp, and the valley height, Rv. The
material which flows from the profile peaks fills up the
valleys leaving the profile without additional peaks. The
theoretical claim that material flow from the peaks
mostly manifests as their widening, is convincingly
illustrated by the experimental results (Figures 8 and 9).
Experimental investigations were conducted on a
universal lathe and it was not possible to precisely define
the depth of the ball penetration into the roughness
profile. In other words, the limited accuracy of the lathe
slide ways, the presence of clearances and the system
compliance prevented the burnishing ball from moving
exactly along the direction defined by the theoretical
considerations. For that reason, we determined the
required displacement in an indirect way, monitoring the
variation of the forces during burnishing. Thus, the
forces were periodically increased in order to achieve a
penetration depth that approximately equals the maxi-
mum peak height of the previously machined surface, Rp.
The diagrams in Figures 8 and 9 clearly show a gradual
decrease of the roughness over the 32–320 N force
interval. One of the basic goals was to visually identify
the oscillation of the profile curve of the burnished
surface, around the line that divides the profile of the
previously machined surface into two, approximately
equal, surface areas of peaks and valleys (Figure 9).
6 CONCLUSIONS
Theoretical assumptions pertaining to defining an
optimal tool trajectory that results in the best surface
quality were largely confirmed in this experiment. A ball
rolling within a high-stiffness tool system, according to a
predefined penetration depth, provides a near-optimal
condition, i.e., minimal surface roughness, regardless of
the quality of the initial machining. Based on the experi-
mental results, the ball should penetrate the surface
roughness profile up to a depth that approximately
equals the maximum peak height achieved by the initial
machining. The results of this study allow the assump-
tion that ball-penetration depths beyond Rp do not signi-
ficantly contribute to the surface quality, primarily
because the displaced material should create new,
probably higher profile peaks. According to the results of
the measurement, regardless of the surface quality on a
previously machined surface, the value of y that is close
to the maximum peak height results in the lowest Ra.
This could be especially valuable when burnishing rela-
tively rough surfaces.
An analysis of the surface-roughness measurement
results and the super positioning of the profiles generated
using various burnishing forces and tool displacements,
largely explained the phenomenon of the roughness
DJ. VUKELIC et al.: BURNISHING PROCESS BASED ON THE OPTIMAL DEPTH OF A WORKPIECE PENETRATION
50 Materiali in tehnologije / Materials and technology 47 (2013) 1, 43–51
peaks’ deformation. Having this and the theoretical
considerations in mind, we can conclude that the defined
penetration depth, y  Rp, satisfies the condition of
approximate equality of surface areas defined by the
roughness profile peaks and valleys. The proposed
burnishing methodology could be especially valuable
when dealing with roughly machined surfaces with
significant Rp.
The research reported in this paper opens up a
number of new, interesting directions for research, such
as the testing of a stiff tool system with various
workpiece materials, different burnishing regimes, and
various surface roughnesses as a result of the initial
machining. We believe that the proposed model casts
new light on the burnishing process. It opens up new
directions of research involving stiff tool systems and
penetration depths that yield a near-minimal surface
roughness, regardless of the workpiece material, the
burnishing parameters, and the initial surface roughness.
Future work should involve an investigation of surfaces
that drastically differ from the aspect of the surface
roughness achieved by previous machining. They should
be penetrated by a burnishing ball beyond the maximum
profile peak height. In this way it would be possible to
verify the results obtained in this study as well as open
up new directions for research on the burnishing process.
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DJ. VUKELIC et al.: BURNISHING PROCESS BASED ON THE OPTIMAL DEPTH OF A WORKPIECE PENETRATION
Materiali in tehnologije / Materials and technology 47 (2013) 1, 43–51 51

D. ]UR^IJA et al.: CALCULATION OF THE LUBRICANT LAYER FOR A COARSE SURFACE ...
CALCULATION OF THE LUBRICANT LAYER FOR A
COARSE SURFACE OF A BAND AND ROLLS
IZRA^UN SLOJA MAZIVA NA GROBI POVR[INI TRAKU IN
VALJEV
Du{an ]ur~ija1, Franc Vodopivec2, Ilija Mamuzi}1
1Croatian Metallurgical Society, Berislavi}eva 6, 10000 Zagreb, Croatia
2Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
mamuzic@simet.hr
Prejem rokopisa – received: 2012-07-03; sprejem za objavo – accepted for publication: 2012-09-26
The effect of the average roughness of a lubricated band caused by dressing processes is analysed by applying the Reynolds
differential equation for lubrication with the incorporated average roughness and evolution in the Fourier series to the third
member. The analysis has shown that the average roughness has two effects on the lubricant-layer thickness in the entering
section of the deformation zone. For a small surface roughness, the nominal lubricant-layer thickness decreases slowly (if the
process is treated as occurring on a smooth surface) and the thickness grows again with an increase in the roughness. The basis
for the analysis was the numerical Monte-Carlo method and the developed approximate analytical solution was in acceptable
agreement with the numerical method.
Keywords: surface roughness, lubricant-layer thickness, Reynolds equation, Monte-Carlo method, Fourier series
Analiziran je vpliv povpre~ne hrapavosti mazanega traku pri procesih dresiranja. Podlaga analize je Reynoldsova diferencialna
ena~ba za mazanje z vklju~eno povpre~no hrapavostjo in obravnavo s Fourierovo vrsto do tretjega ~lena. Analiza je pokazala, da
ima povpre~na hrapavost dva u~inka na debelino plasti maziva v vhodnem preseku podro~ja deformacije. Pri majhni za~etni
hrapavosti se nominalna debelina plasti maziva po~asi zmanj{uje (~e se proces obravnava, kot da poteka na gladki povr{ini) in
znova raste, ~e se pove~uje hrapavost. Podlaga za analizo je bila numeri~na metoda Monte Carlo, razvita pa je bila tudi pribli`na
analiti~na re{itev, ki se zadovoljivo ujema z numeri~no.
Klju~ne besede: hrapavost povr{ine, debelina plasti maziva, Reynoldsova ena~ba, metoda Monte Carlo, Fourierova vrsta
1 INTRODUCTION
This technology is strongly associated with the
quality of technological lubricants as it:
• diminishes the contact friction,
• removes the heat, cools the tool and diminishes the
wear,
• diminishes the deformation resistance and the defor-
mation work,
• diminishes the sticking to the tool and keeps the
surface of the product clean.
The basic groups examined in this work1–3 are:
• liquid emulsions,
• fats and compounds,
• consistent lubricants,
• transparent/glass lubricants,
• powder lubricants and
• metallic lubricants.
Technological lubricants must meet a series of
requirements, beginning with a high lubricity – the
ability to form a flat, firm layer separating the contact
surfaces – then there are thermal consistency and
stability that prevent the damaging effect of the product
corrosion, the properties not posing any health and envi-
ronmental risks, etc.
The liquid emulsions, whose compounds are
mixtures of vegetable and mineral oils, are especially
used in the cold rolling of 0.3–0.4 mm thick sheets and
strips.
In the cold rolling of sheets and strips, the dressing
process is also used with an application of liquid lubri-
cants to reduce undulation.
2 MATHEMATICAL MODELLING
Mathematical modelling is a requirement of today’s
metallurgy4,5 and it is also used in the field of plastic
deformation of metals. For an analysis of smooth
surfaces6,7 the following equation is used:
d
d
Rp
x
v v
x
Q
x
=
+
−
6 120
2 2






( )
( ) ( )
(1)
Q x udy
p
x
x
v v
x
x
( ) ( ) ( )
( )
= = +
+⎛
⎝
⎜ ⎞
⎠
⎟∫
0
3 01
12 2



 
d
d
R
(2)
The geometry of the lubricant contact8 and the length
of the lubricant wedge are described with the relations
(3), (4) and (5):
   ( ) cos sinx R
x
R
= + − − −⎛⎝
⎜ ⎞
⎠
⎟
⎡
⎣
⎢
⎤
⎦
⎥0
2
1 (3)
Materiali in tehnologije / Materials and technology 47 (2013) 1, 53–57 53
UDK 519.61/.64:621.77 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)53(2013)
a R
R R
a= − − +
⎛
⎝
⎜ ⎞
⎠
⎟ −
⎡
⎣
⎢
⎤
⎦
⎥1
0
2
cos sin
 
 (4)
  

( )x x
x
R
x
R
x
R
= − + − +0
2 3
2
4
32 2 8
(5)
For the average sheet roughness9, the mathematical
relation in accordance with Figure 1 is:
d
d
p
x
v v
x xR0
0 2
0
3
0
6
1
1
1
1
= + −
⎡
⎣
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥










( )
( ) ( )
⎥
(6)
  ( ) ( ) ( )x x x0 = + (7)
Reflexion of sheet roughness is added, as 0, to the
lubricant wedge (4). The calculation is possible only
with numerical mathematical methods and, in the pro-
gram MATHEMATICA, the numerical method Monte
Carlo was used. In the theoretical calculations regarding
the model of the average roughness, the following
function developed to the third term of Fourier series was
applied:
( ) (sin sin sin )x x x x Rz= + +
4 1
3
3
1
5
5
π
(8)
3 RESULTS AND DISCUSSION
In Table 1 the standard values of geometrical, rheo-
logical and kinematic characteristics of the processes of
theoretical investigations are given according to the
Russian-Ukrainian10,11 authors.
Table 1: Standard lubricant characteristics for theoretical calculations
Tabela 1: Standardne zna~ilnosti maziva za teoreti~ne izra~une
Parameter Value Unit
- piezo coefficient of viscosity 2.18E-7 Pa–1
p0- rolling pressure 20E6 Pa
vR- circumferential roll speed 10 m/s
v0- sheet speed 6 m/s
R- roll radius 0.35 (0.25) m
μ0- lubricant dynamic viscosity
μ = μ0 exp ( * p0) Barussa
equation
0.024–0.048 Pa s
- gripping angle 0–0.02 rad
a- lubricant thickness on sheet 0.001–0.00001 m
A- technological parameter 1965512
(3934525) m
–1
A = (1–exp(– * p0)/6μ0(v0+vR))
Rz ≈ 6 Rz = 1–10 μm
The parameters in Table 1 are of two groups:
1- lubricant rheological characteristics (μ0, )
2- geometrical characteristics of the technological pro-
cess (R, , Rz)
3- kinematics (v0, vR)
D. ]UR^IJA et al.: CALCULATION OF THE LUBRICANT LAYER FOR A COARSE SURFACE ...
54 Materiali in tehnologije / Materials and technology 47 (2013) 1, 53–57
Figure 1: Model of the tribomechanical system; 1- lubricant layer (x)
– nominal thickness for smooth surfaces, 2- band – in dressing pro-
cesses the adhering angle  is low, 3- average band roughness (x) –
casual sheet roughness, 4- roll defined by surface smoothness. In
Table 1 the roughness is Rz = 8 μm.
Slika 1: Model tribomehanskega sistema; 1- plast maziva (x) – nomi-
nalna debelina za gladke povr{ine, 2- trak – pri procesu dresiranja pri
majhnem kotu stika , 3- povpre~ma hrapavost traku (x) – slu~ajna
hrapavost traku, 4- valj z definirano gladkostjo povr{ine. V Tabeli 1 je
njegova hrapavost Rz = 8 μm.
Table 2: Lubricant-layer exit results (μm)
Tabela 2: Izhodni rezultati za plast maziva (μm)
4- compounds (A, roughness space angle)
The solutions of differential equation (6) are partially
given in Table 2.
The examined roughness is classified12,13 in 10 verti-
cal classes and the band profile roughness in 32 horizon-
tal classes.
In principle, with a decreasing band-lubricant thick-
ness (a in Figure 1) the lubricant thickness in the
entering section of the metal deformation zone is also
decreased (0). As shown in14, the lubricant wedge has
the ideal geometry and can give economic savings of the
lubricant in the metalworking technology.
The numerical integration of equation (6) was
checked with the approximate15–17 analytical solutions
possible in the case of practical interest, which is found
in equations (9), (10a)–(10e) and (11). Equation (9) is
the simplest analytical solution that does not consider the
thickness of the band lubricant layer, a >> 0. With a
clear complexity, equation (11) corrects this deficiency:
315 168 1824 0
05
1
6
3 7 2 4 2
0
2 0
AR R
R A
exp p
  
 



− − =
= =
− −
.
( )

( )v v R0 +
(9)
( )
W A
R
R
R
R R
R
1
2
0
2
4
3
4
3 1
= − −
− − +⎛
⎝
⎜ ⎞
⎠
⎟ −
⎡
⎣
⎢
⎢
⎤

 


cos sina
⎦
⎥
⎥
⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
⎥
⎥
⎥
3
(10a)
( )
− = −
− − +⎛
⎝
⎜ ⎞
⎠
⎟ −
W
R
R
R
R R
R
2 3
4
0
2
3
16
7
16
3 1



 



cos sina
⎡
⎣
⎢
⎢
⎤
⎦
⎥
⎥
⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
⎥
⎥
⎥
7
(10b)
W3
03
6
=
+
+
  
 

 

  (10c)
( )
W
R
R
R
R R
R
4 2
3
0
2
8
5
8
5 1
= −
− − +
⎛
⎝
⎜ ⎞
⎠
⎟ −
⎡
⎣
⎢
⎤
⎦
⎥


 


cos sina
5 (10d)
( )
− =
− − +⎛
⎝
⎜ ⎞
⎠
⎟ −
⎡
⎣
⎢
⎢
⎤
⎦
⎥
⎥
W
R
R
R R
R
5
5
0
2
932
5 1 cos sin
 
a
(10e)
W1 + W2 + W3 * (W4 + W5) =0 (11)
In Table 3 approximate numerical and analytical
solutions are compared. The approximate numerical
solutions can be compared with numerical integration
only for the entering roughness profile, thus, at the
entering section of the deformation zone with x = 0.
It is clear from Table 3 that the simple analytical
form of equation (9) with numerous approximations
describes well the lubricant layer for the case of a lubri-
cant excess on the sheet and the rolls.
Table 3: Comparison of approximate analytical and numerical
Monte-Carlo solutions for one point of the graph crossing from
Figure 2
Tabela 3: Primerjava pribli`nih analiti~nih in numeri~nih re{itev
Monte Carlo za eno to~ko prereza grafa na sliki 2
Case conditions
Approximate
analytical solutions,
eq. (11) and (9)
Monte-Carlo
method,
eq. (6)
x = 0 (initial
roughness profile)
Rz = 1 μm
Rz ≈ 6 
 = 0.00918759 rad
A = 1965512 m–1
R = 0.35 m
a = 0.001 m
0 = 14.721 μm (11)
0 = 14.771 μm (9)
a = 0.0001 m
0 = 13.834 μm (11)
a = 0.001 m
0 = 14.772 μm
a = 0.0001 m
0 = 13.761 μm
x = 0 (initial
roughness profile)
Rz = 10 μm
Rz ≈ 6 
 = 0.0092867 rad
A = 1965512 m–1
R = 0.35 m
a = 0.001 m
0 = 15.024 μm (11)
0 = 15.092 μm (9)
a = 0.0001 m
0 = 13.511 μm (11)
a = 0.001 m
0 = 15.077 μm
a = 0.0001 m
0 = 13.429 μm
x = 0 (initial
roughness profile)
Rz = 10 μm
Rz ≈ 6 
 = 0.00840867 rad
A = 3934525 m–1
R = 0.25 m
a = 0.001 m
0 = 8.776 μm (11)
0 = 8.838 μm (9)
a = 0.0001 m
0 = 8.464 μm (11)
a = 0.001 m
0 = 8.755 μm
a = 0.0001 m
0 = 8.429 μm
Table 4: Effect of the two-sided roughness of the sheet and rolls,
congruous for 0
Tabela 4: Vpliv dvostranske hrapavosti traku in valjev, kongruenten
za 0
x = 0 (initial roughness
profile)
Rz = 10 μm, average
roughness, horizontal
(transversal)
Rz = 8 μm, longitudinal
roll roughness
Rz ≈ 6  (GOST
2789-73)
 = 0.00840867 rad
A = 3934525 m–1
R = 0.25 m
Monte-Carlo method
a = 0.0001 m
0 = 9.299 μm
0 = 8.429 μm
One-sided
roughness of the
sheet
a = 0.0001 m
0 = 8.429 μm
Two-sided
roughness of the
sheet and roll
a = 0.0001 m
0 = 8.919 μm
Rz → 0
a = 0.0001 m
0 = 7.877 μm
The longitudinal band profile on abscissa in shown in
66 classes and on ordinate in 11 classes for roughness
(0–10 μm). It is useful to calculate the lubricant thick-
ness 0 in the range of 8.5–12.5 μm in the area of I-I. Q,
K and W designations connect the specific areas of the
network diagram with the contour plot (an aircraft
picture of the network diagram).
B and C are the left and right sides of the band
roughness defined as a sine evolution function in the
Fourier series: B in the range of (–2) rad and C in
(0–) rad.
Line P in Figure 2 represents the nominal lubricant-
layer thickness on side C, thus, by having the thickness
for Rz  6 μm, an equivalent to the lubricant-layer thick-
ness on a smooth surface is obtained. Side B does not
have this property.
Materiali in tehnologije / Materials and technology 47 (2013) 1, 53–57 55
D. ]UR^IJA et al.: CALCULATION OF THE LUBRICANT LAYER FOR A COARSE SURFACE ...
In Figure 3 both sides of the roll longitudinal rough-
ness C from Figure 2 are shown. The average roughness
conserves the same properties as in Figure 2. The longi-
tudinal roughness profile in the range of classes 33 to 66
gives a more stable hydrodynamic lubrication, while for
classes 1 to 33 the hydrodynamic lubrication is already
seriously impaired by the low roughness of the band and
rolls. The lubricant layer decreases rapidly and spreads
to fractal areas. A stable lubrication can be achieved on
small band segments and around class 4 of the longitudi-
nal sheet profile and around classes 10 and 30. The
complex shapes of the lubrication space are probably
determined by the band and roll roughness in the
entering section of the deformation zone that determines
a different lubrication layer than in the case of smooth-
sheet and roll surfaces.
4 CONCLUSIONS
Based on the results of theoretical analyses of the
effect of the band roughness on the lubrication dressing
processes, the following conclusions are proposed:
• The average band roughness has a critical value when
it starts to affect positively the lubricant layer with its
increase in comparison with a smooth surface. Up to
line P in Figure 2, the lubricant layer has a tendency
to increase and to decrease the formation of sunk
baskets in area Q. The theoretical explanation for this
is that the surface roughness determines the shape of
the lubricant layer for every value of Rz. This is the
range of a stable lubrication.
• If congruous roll roughness is added to the average
band roughness, forming a longitudinal roll rough-
ness with the positive side in the range of (0–), the
thickness of the lubricant layer in the entering section
of the band deformation zone will increase its
longitudinal profile from class 33 to 66 (Figure 3 and
Table 4) and will approach the boundary lubrication.
• The developed approximate analytical solutions agree
with the numerical integration of equation (6) and
ensure a reliable approach to the analysis.
• If the technological process was performed with a
nominal lubricant-layer thickness marked with line P
in Figure 2 the best roll rhythm would be obtained
without significant fluctuations of the lubricant
thickness, especially in the case of the boundary-
lubrication proximity. This includes the control of the
roll roughness.
5 SYMBOLS AND FIGURES
Symbol Unit Comment
0 m, (μm) Lubricant thickness in the entering
section of the deformation zone
(Figure 1)
(x) m Lubricant thickness in the range of [–a :
0], Figure 1, equations (3) and (5)
a m Lubricant thickness ahead of the
entering section of the deformation zone
a m Length of the lubricant wedge
(Figure 1), equation (4)
 rad Band dressing angle
vR m/s Circumferential roll speed
D. ]UR^IJA et al.: CALCULATION OF THE LUBRICANT LAYER FOR A COARSE SURFACE ...
56 Materiali in tehnologije / Materials and technology 47 (2013) 1, 53–57
Figure 2: Effect of the average sheet roughness and smooth rolls on 0
Slika 2: Vpliv povpre~ne hrapavosti traku in gladkih valjev na 0
Figure 3: Effect of the average sheet roughness and longitudinal roll
roughness on 0 (Table 4)
Slika 3: Vpliv povpre~ne hrapavosti traku in longitudinalne hrapavosti
valjev na 0 (tabela 4)
vT m/s Mandrel speed
R m Roll radius
Rz m Roughness of the band surface,
equation (8)
2 Dispersion roughness of the sheet and
rolls according to equation (9)
x Casual lubricant thickness depending on
the band roughness (and rolls)
< > Operative mathematical expectation
x, y Descartes coordinates
Q(x) – Volume use of lubricant (on the band
perimeter)
μ0 Pa s Lubricant dynamic viscosity by the
rolling pressure
μ Pa s Lubricant dynamic viscosity by the air
pressure
u m/s Lubricant rate on the abscissa
 m2/N Piezo coefficient of lubricant viscosity
p Pa Rolling pressure
Q m2/s Use of lubricant on the mandrel
perimeter – a one-dimensional model
dp/dx Pa/m Pressure gradient in the lubricant layer,
equation (1)
sin  rad Marking the trigonometric function for
the griping alpha angle
H m Enter band thickness
h m Exit band thickness
A m–1 Technological parameter:
A = [1– exp(–p)] / [6μ0(vR + v0) ]
exp, 
14
1 μm
2.718
1–1
10–6 m
Base of natural logarithm (3.141)
Reference
Micrometre
S μm Band- and roll-roughness classes
L μm Longitudinal holding-band profile
Q, K, W Markers for Figure 2
6 REFERENCES
1 I. Mamuzi}, V. M. Drujan, Teorija, Materijali, Tehnologija ~eli~nih
cijevi, Hrvatsko Metalur{ko Dru{tvo, Zagreb 1996, 428–435
2 D. ]ur~ija, Mater. Tehnol., 37 (2003) 5, 237–254
3 D. ]ur~ija, I. Mamuzi}, F. Vodopivec, Metalurgija, 45 (2006) 3, 250
(Summary)
4 A. I. Gubin, B. B. Veselovskiy, D. ]ur~ija, A. A. Kochubey, Mathe-
matical simulation and choice of optimum thermal models of con-
tinuous events, Metalurgija, 47 (2008) 3, 255 (Summary)
5 Iu. V. Brazaluk, O. O. Kochubey, D. ]ur~ija, M. V. Polyakov, D. V.
Yevdokymov, On a mathematical model of particle in liquid metal,
Metalurgija, 47 (2008) 3, 256 (Summary)
6 D. ]ur~ija, I. Mamuzi}, Mater. Tehnol., 43 (2009) 1, 23–30
7 D. ]ur~ija, I. Mamuzi}, Goriva i maziva, 48 (2009) 1, 3–28
8 D. ]ur~ija, I. Mamuzi}, Mater. Tehnol., 42 (2008) 2, 59–63
9 D. ]ur~ija, I. Mamuzi}, Metalurgija, 44 (2005) 4, 295–300
10 O. P. Maksimenko, N. P. Podberezniij, Izvestija ^ernaja metallurgija,
73 (2003) 10, 12–16
11 O. P. Maksimenko, A. A. Semen~a, Su~asni problemi metalurgii, 8
(2005), 99–103
12 O. P. Maksimenko, O. E. Lejko, Su~asni problemi metalurgii, 8
(2005), 93–99
13 P. I. Klimenko, Su~asni problemi metalurgii, 8 (2005), 44–49
14 S. M. Ionov, V. I. Kantorovi~, S. A. [epovalov, A. N. Krjukov,
Su~asni problemi metalurgii, 8 (2005), 224–228
15 D. M. Me, S. P. Liu, J. F. Zheng, Met. Form. Technol., 20 (2002) 5,
29–32
16 Y. T. Keun, B. H. Lee, R. H. Wagner, J. Mater. Process. Technol.,
130 (2002), 60–63
17 D. ]ur~ija, I. Mamuzi}, Goriva i maziva, 46 (2007) 1, 23–44
D. ]UR^IJA et al.: CALCULATION OF THE LUBRICANT LAYER FOR A COARSE SURFACE ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 53–57 57

Y. ÖZTÜRK et al.: STRUCTURAL AND MAGNETIC PROPERTIES OF CERIUM-DOPED YTTRIUM-IRON ...
STRUCTURAL AND MAGNETIC PROPERTIES OF
CERIUM-DOPED YTTRIUM-IRON GARNET THIN FILMS
PREPARED ON DIFFERENT SUBSTRATES USING THE
SOL-GEL PROCESS
STRUKTURNE IN MAGNETNE LASTNOSTI S CERIJEM
DOPIRANE ITRIJ-@ELEZOVE GARNETNE TANKE PLASTI,
IZDELANE S SOL-GEL POSTOPKOM NA RAZLI^NIH PODLAGAH
Yavuz Öztürk1, Mustafa Erol2, Erdal Çelik2,3, Ömer Mermer1,3, Gökalp Kahraman1,
Ibrahim Avgýn1
1Ege University, Dept. of Electrical and Electronics Engineering, Bornova, 35100 Izmir, Turkey
2Dokuz Eylul University, Dept. of Metallurgical and Materials Engineering, Buca, 35160 Izmir, Turkey
3Dokuz Eylul University, Center for Production and Applications of Electronic Materials (EMUM), Buca, 35160 Izmir, Turkey
yavuzoz@hotmail.com
Prejem rokopisa – received: 2012-07-13; sprejem za objavo – accepted for publication: 2012-08-27
The cerium-substituted yttrium-iron garnet (Ce-YIG) CexY3–xFe5O12 is considered as a promising material for applications in
high-performance magnetic and magneto-optic devices. In this work cerium-substituted yttrium-iron garnet films were produced
on fused silica and Si(100) substrates using the sol-gel technique from solutions with the yttrium/cerium molar ratio 2.8/0.2. A
heat treatment was applied to those as-deposited garnet films at temperatures ranging from 800 °C to 1000 °C for 2 h in air. The
as-deposited garnet films were characterized by X-ray diffraction (XRD) and scanning electron microscopy (SEM) to
investigate their structural properties. A vibrating-sample magnetometer was used at room temperature to characterize the
magnetic properties of the as-deposited garnet thin films.
Keywords: sol-gel, yttrium-iron garnet, magnetic properties
S cerijem nadome{~en itrij-`elezov garnet (Ce-YIG, CexY3–xFe5O12) je obetajo~ material za uporabo v visoko zmogljivih
magnetnih in magnetnoopti~nih napravah. V tem delu je bila izdelana s cerijem nadome{~ena itrij-`elezova garnetna tanka plast
na kremenovem steklu in podlagi Si(100) z uporabo sol-gel tehnike iz raztopine z molskim razmerjem itrij/cerij 2,8/0,2.
Nanesena garnetna plast je bila 2 h toplotno obdelana na zraku v temperaturnem obmo~ju od 800 °C do 1000 °C. Te plasti so
bile karakterizirane z rentgensko difrakcijo (XRD), njihove strukturne zna~ilnosti pa so bile pregledane z vrsti~nim
elektronskim mikroskopom (SEM). Vibracijski magnetometer je bil uporabljen pri sobni temperaturi za ovrednotenje magnetnih
lastnosti nanesenih tankih garnetnih plasti.
Klju~ne besede: sol-gel, itrij-`elezo garnet, magnetne lastnosti
1 INTRODUCTION
Pure YIG films and their substituted derivatives have
been researched for decades because of their wide range
of applications in the microwave, communication and
magnetic detection areas1–3. For instance, YIG is known
to be one of the important ferrites in 1–2 GHz micro-
wave applications owing to its small FMR line width4.
However, the integration of garnets into semiconductor
electronics is not straightforward as this process requires
the garnet materials to be in nano/micro-sized dimen-
sions. The magnetic and magneto-optical properties of
YIG thin films will be affected by some parameters,
namely, the type of the substituted material, the synthesis
methods, the substrate and the structure/microstructure
of YIG films5–9.
YIG is the most representative and well-known com-
pound among the rare-earth iron garnets. The definite
composition and the presence of only trivalent metal ions
make YIG particularly suitable for magnetic studies.
There are eight formula units, Y3Fe2(FeO4)3, in a unit cell
with a lattice constant a = (1.2376 ± 0.0004) nm. Some
magnetic properties, such as magnetization, remanence,
coercivity, Faraday rotation depend critically on the
structure and the microstructure of the materials. Also, it
is well known that deviations from stoichiometry have a
strong influence on the magnetic properties of ferrites.
There are three sub-lattices: tetrahedral (d), octahedral
(a) and dodecahedral (c) in YIG and they are surrounded
by four, six and eight oxygen ions, respectively. Among
the five iron ions, which represent a formula unit, three
are in 16 octahedral sites and two are in 24 tetrahedral
sites10. A magnetic moment of 4.64x10–24 J/T per formu-
la unit results from anti-ferromagnetic super-exchange
interaction between the Fe3+ ions in these two different
sites through the intervening O2– ions. This corresponds
to the moment of the one Fe3+ ion that is present at a
tetrahedral site in numbers greater than the Fe3+ ions at
an octahedral site. YIGs are also of scientific importance
because of the wide variety of magnetic properties that
we can obtain when substituting Y by rare-earth metals
or substituting Fe by other trivalent cations.
Materiali in tehnologije / Materials and technology 47 (2013) 1, 59–63 59
UDK 537.622:621.318.1 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)59(2013)
A variety of techniques have been applied to obtain
YIG thin films, such as radio-frequency (RF) sputtering,
liquid-phase epitaxial (LPE) growth and pulsed-laser
deposition (PLD). Most of these methods are generally
vacuum type and expensive. Aside from vacuum-type
expensive techniques, different wet chemical methods
have been used to obtain the YIG structure, such as
sol-gel, co-precipitation, micro-emulsion synthesis,
citrate gel routes, hydrothermal synthesis, etc.4–8. In
related researches, production processes usually focused
on obtaining powder and bulk materials instead of
producing thin films. Powder or bulk ceramic production
techniques require high annealing temperatures and long
processing duration, as mentioned in ref11. Materials that
were used as components of electronic, magneto-optic or
magnetic devices are all physically based on the
movement of ions, the interaction of light and the change
of the orientation in the structure. These structural
properties change negatively with increasing material
thickness or volume. In this way, producing thin films
instead of thick films (coatings) or bulk materials will
provide low annealing temperatures and a controlled
structure. To the best of our knowledge, there are few
studies on obtaining YIG thin films with chemical
methods4–8,12. Unlike obtaining powder, producing thin
films can be achieved at relatively low temperatures, like
700–1000 °C using the sol-gel method as a wet chemical
route8. Nevertheless, it has been proved that the sol-gel
process offers considerable advantages, such as better
mixing of the starting materials and excellent chemical
homogeneity in the final product. Moreover, the mole-
cular level mixing and the tendency of partially hydro-
lyzed species to form extended networks facilitate the
structure evolution, thereby lowering the crystallization
temperature13,14.
In this study, CexY3-xFe5O12 thin films were success-
fully deposited from solution with a Y/Ce molar ratio of
2.8/0.2 on Si(100) and fused silica substrates through the
sol-gel method from solutions that were synthesized with
ethylhexanoate and 2,4-pentanedionate based precursors.
2 EXPERIMENTAL DETAILS
The precursor materials Fe 2,4-pentanedionate
(0.1766 mg), Yttrium 2-Ethylhexanoate (0.1452 mg) and
Ce 2-Ethylehexonate (0.01140 mg), were dissolved in
methanol and glacial acetic acid (GAA) in order to form
a 0.23-M solution with the chemical composition Ce :
Y : Fe = 0.2 : 2.8 : 5. Three different solutions (A, B and
C) with different methanol and GAA ratios were
prepared as listed in Table 1. Table 1 also shows the pH
values of the solutions. In this study, GAA acts as a
chelating agent to involve a homogenous solution. A
higher GAA concentration leads to a poor interaction
with the substrate and a lower GAA concentration leads
to poorly dissolved precursors in a manner of the macro
view. As a result, solution B was determined to be
appropriate for deposition on substrates, since it has
good wetting and chelating properties. The prepared
optimal solutions were dip-coated on the fused silica and
Si(100) substrates at room temperature in air. The gel
films were dried at 300 °C for 10 min, and consequently
heat treated in the range 800–1000 °C for 2 h in air in an
electrical furnace. After this procedure, the specimens
were cooled down from the annealing temperatures.
Table 1: Solvent – Chelating agent contents and pH values of the
solutions A, B and C
Tabela 1: Vsebnost in pH vrednost raztopin topilo – kelat, A, B in C
Solution Methanol (ml) GAA (ml) pH
A 2 1.5 3.6
B 2.5 1 3.05
C 3 0.5 2.5
X-ray diffraction (XRD, Rigaku D/MAX-2200/PC)
patterns of the films were determined to identify the
phase structure. The surface properties and topographies
of the films were examined using scanning electron
microscopy (SEM, JEOL JSM 6060) with attached
energy-dispersive spectroscopy (EDS). The magnetic
properties of samples were determined with a vibrating-
sample magnetometer (VSM, Lakeshore 736, 7400) at
room temperature.
3 RESULTS AND DISCUSSION
XRD patterns of selected samples on Si(100) and
fused silica substrates are depicted in Figures 1 and 2
respectively. All the produced films contain cubic
Ce-substituted YIG films. As reported in the
literature15–17, cubic YIG formation with other impurity
phases such as Y2O3, FeYO3 and Fe2O3 was generally
observed at temperatures between 700 °C and 1100 °C.
In the present research, neither Y2O3 nor FeYO3 phase
formations were observed in the YIG film on either
60 Materiali in tehnologije / Materials and technology 47 (2013) 1, 59–63
Y. ÖZTÜRK et al.: STRUCTURAL AND MAGNETIC PROPERTIES OF CERIUM-DOPED YTTRIUM-IRON ...
Figure 1: XRD patterns of Ce-YIG films coated on Si(100), annealed
between 800 °C and 1000 °C for 2 h in air
Slika 1: XRD-posnetki Ce-YIG plasti na Si(100), `arjeni med 800 °C
in 1000 °C, 2 h na zraku
Si(100) or fused silica substrates. However, even though
the Fe2O3 phase was found in the films on fused silica, it
was not determined in the films on Si(100). At the same
time, two different crystal structured yttrium silicate
phases that transform from -Y2Si2O7 to -Y2Si2O7 at
1000 °C were found. The formation of these silicate
phases was considered to be substrate-film interactions.
This formation and the increased annealing temperature
affect the crystallization behaviour of the garnet in a
positive way.
It is well known that the activation energy for
nucleation is reduced by good lattice matching across the
interface. As mentioned ref.18 the nucleation mechanism
depends on the substrates and denser nucleation was
observed with an increasing lattice match. When the
XRD results of the films on both substrates were
compared, the Si substrates that were oriented exhibited
better crystallization features than the amorphous fused
silica. The different crystallization behavior observed for
two substrates can be explained by the formation of sili-
cate phases and differences in the nucleation mechanism.
Furthermore, fused silica has a high resistance to
chemicals, which reduces both interactions with the
solution and silicate phase formation.
Sol-gel deposition is a wet chemical route in which
the film quality is affected by various parameters such as
substrate interaction, pH, humidity, and temperature. In
order to tailor the magnetic and magneto-optical proper-
ties, film quality and homogeneity must be taken into
consideration, as reported in refs19,20. The microstructural
properties of Ce-YIG films were denoted in Figure 3. As
can be seen from these micrographs, we have
successfully obtained a coating of garnet structure with
minor cracks. Figure 3a provides a general view of the
structure of the films. In addition, some micro-size
cracks structure can be seen on the magnified image,
which is the result of the substrate film interaction, as
shown in Figure 3b. The optimum thicknesses of the
films were found to be around 300 nm using a profilo-
meter. As far as magnetic properties are concerned, the
magnetic hysteresis loop (M-H) of the Ce-YIG films on
different substrates annealed at different temperatures
was recorded with the VSM at room temperature. Fig-
ures 4 and 5 diagrammatically clarifies the M-H loop of
the Ce-YIG layers grown on the Si substrate annealed at
Y. ÖZTÜRK et al.: STRUCTURAL AND MAGNETIC PROPERTIES OF CERIUM-DOPED YTTRIUM-IRON ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 59–63 61
Figure 2: XRD patterns of Ce-YIG films coated on fused silica,
annealed between 800 °C and 1000 °C for 2 h in air
Slika 2: XRD-posnetki Ce-YIG plasti na taljenem kremenu, `arjene
med 800 °C in 1000 °C, 2 h na zraku
Figure 4: Magnetic hysteresis loops of Ce-YIG on Si(100) substrate
annealed at 1000 °C
Slika 4: Magnetna histerezna zanka Ce-YIG na podlagi Si(100),
`arjeno pri 1000 °C
Figure 3: XRD SEM micrographs of Ce-YIG on Si(100) prepared at
1000 °C
Slika 3: XRD SEM-posnetka Ce-YIG na Si(100), pripravljenem pri
1000 °C
1000 °C and 800 °C, respectively. The applied magnetic
field (Hex) is both in plane and perpendicular with respect
to the Si wafer. Different annealing temperatures were
also applied on the Ce-YIG film on Si and their M-H
loops were also measured, but are not shown in the
figure. All the parameters related to the magnetic proper-
ties of these films were summarized in Table 2. These
results indicate that strong magnetic anisotropy has been
detected for different annealing temperatures of the
Ce-YIG films. As the annealing temperature increases,
the coercivity values of the film decreases for a field
perpendicular to the Si wafer but this does not change
significantly for the in-plane field. The saturation mag-
netization values increase strongly as the annealing
temperature increases. This shows that an increase of the
annealing temperature changes the substrate–film
interaction21 and also increases the Ce-YIG phase with
respect to the Y2Si2O7 phases.
In order to compare the effect of different substrates
on the magnetic properties, the Ce-YIG film was coated
on fused silica. Again, all the films were annealed at the
same temperature as that of the Ce-YIG on the Si sub-
strate. Figure 5 shows the M-H loop of Ce-YIG on fused
silica annealed at 800 °C. The magnetic properties of
these films for different annealing temperatures were
also summarized in Table 3. As is evident from Table 3,
increasing the annealing temperature does not change
significantly the coercivity values for a field perpendicu-
lar to wafer; however, it decreases them for an in-plane
field. For fused-silica substrates the saturation magneti-
zation values decrease strongly when the annealing
temperature increases. In this case decreasing the
saturation magnetization is mainly due to the formation
of the Fe2O3 phase in the films, as shown by the XRD
pattern in Figure 2. This was also confirmed in our pre-
vious study12.
A different range of magnetization values as com-
pared to the bulk in garnet thin films have commonly
been observed21–23. Lower values of the magnetization
explained by parasitic phases in the structure or poorly
crystallized and magnetically disordered grain-boundary
materials24. A strong in-plane magnetic anisotropy was
observed for the Ce-YIG films grown on both substrates,
indicating that the substrate–film interaction has an
important role in the formation of the crystalline
structure as well as the magnetic properties. Because of
the in-plane anisotropy, these kinds of films are useful
for applications of planar waveguides and magnetic
biasing.
4 CONCLUSIONS
Ce-substituted YIG thin films were synthesized on
fused silica and Si(100) substrates with solutions pre-
pared from Ce, Y, and Fe-based precursors. The micro-
structural (SEM) results indicate good surface quality
with micro cracks. Also, two different crystallization
Y. ÖZTÜRK et al.: STRUCTURAL AND MAGNETIC PROPERTIES OF CERIUM-DOPED YTTRIUM-IRON ...
62 Materiali in tehnologije / Materials and technology 47 (2013) 1, 59–63
Figure 5: Magnetic hysteresis loops of Ce-YIG on fused-silica
substrate annealed at 800 °C
Slika 5: Magnetna histerezna zanka Ce-YIG na taljenem kremenu,
`arjeno pri 800 °C
Table 2: Magnetic properties of Ce-YIG films on Si substrate at different annealing temperatures
Tabela 2: Magnetne lastnosti plasti Ce-YIG na Si-podlagi pri razli~nih temperaturah `arjenja
Substrate Annealing temp.(°C)
Saturation
Magnetization
(kA/m)
Perpendicular In-plane
Remanence
(kA/m)
Coercivity
(kA/m)
Remanence
(kA/m)
Coercivity
(kA/m)
Si(100) 800 33 4 4.856 11 2.070
Si(100) 900 51 4 3.185 15 2.070
Si(100) 1000 78 7 2.866 22 2.468
Table 3: Magnetic properties of Ce-YIG films on fused-silica substrate at different annealing temperatures
Tabela 3: Magnetne lastnosti plasti Ce-YIG na kremenovem steklu pri razli~nih temperaturah `arjenja
Substrate Annealing temp.(°C)
Saturation
Magnetization
(kA/m)
Perpendicular In-plane
Remanence
(kA/m)
Coercivity
(kA/m)
Remanence
(kA/m)
Coercivity
(kA/m)
Fused silica 800 83 7 4.299 41 2.229
Fused silica 900 49 6 4.936 28 2.627
Fused silica 1000 39 5 4.459 23 2.946
dynamics were observed for the two substrates. The YIG
phase was obtained at 800 °C with the highest magneti-
zation value (83 kA/m) in all the samples deposited on
the fused-silica substrates. The Fe2O3 phase was obtained
at higher annealing temperatures. For the Si(100) a
strong substrate-layer interaction was observed, which
causes yttrium silicate phases. Higher annealing tempe-
ratures led to a substrate-film interaction with higher
crystallization and increased magnetization. All the films
show strong in-plane anisotropy that would be suitable
for waveguide- or magnetic-biasing-based devices.
Acknowledgement
This work was financially supported by the Scientific
and Technological Research Council of Turkey
(TUBITAK, project number 106T651).
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Y. ÖZTÜRK et al.: STRUCTURAL AND MAGNETIC PROPERTIES OF CERIUM-DOPED YTTRIUM-IRON ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 59–63 63

R. CHENG, X. Y. WU: STRUCTURAL STABILITY AND ELECTRONIC PROPERTIES OF AlCu3, AlCu2Zr AND AlZr3
STRUCTURAL STABILITY AND ELECTRONIC
PROPERTIES OF AlCu3, AlCu2Zr AND AlZr3
STABILNOST STRUKTURE IN ELEKTRONSKE LASTNOSTI
AlCu3, AlCu2Zr IN AlZr3
Rong Cheng, Xiao-Yu Wu
Shenzhen Key Laboratory of Advanced Manufacturing Technology for Mold & Die, College of Mechatronics and Control Engineering,
Shenzhen University, Shenzhen 518060, P. R. China
songxp12345@yeah.net
Prejem rokopisa – received: 2012-07-23; sprejem za objavo – accepted for publication: 2012-08-28
First-principles calculations were performed to study the alloying stability and electronic structure of the Al-based intermetallic
compounds AlCu3, AlCu2Zr and AlZr3. The results show that the lattice parameters obtained after the full relaxation of the
crystalline cells are consistent with the experimental data, and these intermetallics have a strong alloying ability and structural
stability due to their negative formation energies and their cohesive energies. A further analysis revealed that the single-crystal
elastic constants at zero-pressure satisfy the requirements for the mechanical stability of cubic crystals. The calculations on
Poisson’s ratio show that AlCu3 is much more anisotropic than the other two intermetallics. In addition, calculations on the
densities of states indicate that the valence bonds of these intermetallics are attributed to the valence electrons of Cu 3d states
for the AlCu3, Cu 3d and Zr 4d states for AlCu2Zr, and Al 3s, Zr 5s and 4d states for AlZr3, respectively. In particular, the
electronic structure of the AlZr3 shows the strongest hybridization.
Keywords: AlCu3, AlCu2Zr, first-principles, electronic structure
Narejeni so bili prvi na~elni izra~uni stabilnosti legiranja in elektronske strukture aluminijevih intermetalnih zlitin (AlCu3,
AlCu2Zr in AlZr3). Rezultati ka`ejo, da se mre`ni parametri po polni relaksaciji kristalnih celic ujemajo z eksperimentalnimi
podatki in da imajo te intermetalne zlitine veliko sposobnost legiranja ter stabilno strukturo zaradi negativne tvorbene energije
in kohezivnih energij. Nadaljnje analize so pokazale, da elasti~na konstanta monokristala pri ni~elnem tlaku ustreza zahtevam
mehanske stabilnosti kubi~nega kristala. Izra~un Poissonovega koli~nika poka`e, da je AlCu3 bolj anizotropen kot drugi dve
intermetalni zlitini. Dodatno izra~un gostote stanj poka`e, da so valen~ne vezi teh treh intermetalnih zlitin vezane na valen~ne
elektrone Cu 3d-stanja za AlCu3, Cu 3d in Zr 4d-stanja za AlCu2Zr, ter Al 3s, Zr 5s in 4d-stanja za AlZr3, posebno elektronska
struktura AlZr3 pa ka`e najmo~nej{o hibridizacijo.
Klju~ne besede: AlCu3, AlCu2Zr, na~elen izra~un, elektronska struktura
1 INTRODUCTION
Intermetallics involving aluminum and transition
metals (TMs) are known to have a high resistance to
oxidation and corrosion, elevated-temperature strength,
relatively low density, and high melting points, which
make them desirable candidates for high-temperature
structural applications1,2. In particular, zirconium can
effectively enhance the mechanical strength of the alloys
when copper and zinc elements exist in aluminum and
Al-based alloys3. Adding Zr to the Al-Mg alloys can
effectively remove or reduce hydrogen, grain refinement,
pinholes, porosity and hot cracking tendency, and so
improve the mechanical properties4. Many investigations
have focused on the constituent binary systems, such as
Al-Cu, Al-Zr, and Cu-Zr 5–10; however, there has been a
lack of systematic theoretical and experimental investi-
gations for binary and ternary systems, especially for
ternary alloy systems.
In recent years, first-principles calculations based on
the density-functional theory have become an important
tool for the accurate study of the crystalline and elec-
tronic structures and mechanical properties of solids11. In
the present study, we report on a systematic investigation
of the structural, elastic and electronic properties of
Al-based alloys (AlCu3, AlZr3 and AlCu2Zr) using
first-principles calculations, and the results are discussed
in comparison with the available experimental data and
other theoretical results.
2 COMPUTATIONAL METHOD
All the calculations were performed using the Vienna
ab-initio Simulation Package (VASP)12,13 based on the
density-functional theory (DFT)14. The exchange and
correlation energy was treated within the generalized
gradient approximation of Perdew–Wang 91 version
(GGA-PW91)15. The interaction between the valence
electrons and the ions was described by using potentials
generated with Blöchl’s projector augmented wave
(PAW) method16. The PAW potential used for Al treats
the 3s, 3p states as valence states, and the other elec-
tron-ion interaction was described by the 3d, 4s valence
states for Cu, 5s, 4d, 5p valence states for Zr. A plane-
wave energy cut-off was set at 450 eV for the AlCu3 and
AlCu2Zr, and at 350 eV for the AlZr3. Brillouin Zone
integrations were performed using the Monkhorst-Pack17
k-point meshes, e.g., the k-point meshes for AlCu3,
Materiali in tehnologije / Materials and technology 47 (2013) 1, 65–69 65
UDK 669.715:544.225 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)65(2013)
AlCu2Zr and AlZr3 were 15×15×15, 9×9×9 and
13×13×13 for optimizing the geometry and calculating
the elastic constants, and 25×25×25, 19×19×19and
23×23×23 for calculations of the density of states (DOS)
at the equilibrium volume, respectively. Optimizations of
the structural parameters (atomic positions and the lattice
constants) for each system were performed using the
conjugate gradient method, and the coordinates of the
internal atoms were allowed to relax until the total forces
on each ion were less than 0.01 eV/(10–1 nm). The total
energy and density of states (DOS) calculations were
performed with the linear tetrahedron method using
Blöchl corrections18. In order to avoid wrap-around
errors, all the calculations were performed using the
"accurate" setting within VASP.
3 RESULTS AND DISCUSSION
3.1 Equilibrium properties
The AlCu3 and AlZr3 alloys have the simple cubic
Cu3Au (L12 type, space group Pm-3m) structure19,20. The
AlCu2Zr alloy is a partially ordered Cu2MnAl-type fcc
structure with the Fm-3m space group21. Firstly, these
crystal structures were optimized with a relaxation of the
cell shape and the atomic positions. The equilibrium
volume V0, bulk modulus B0 and the pressure derivation
of the bulk modulus B’0 of the AlCu3, AlCu2Zr and AlZr3
were determined by fitting the total energy calculated at
different lattice-constant values to a Birch-Murnaghan
equation of state22. The results of the first-principles
calculations are listed in Table 1. From Table 1 it is
clear that the results of our calculations compare very
favorably with the experimental data. This shows that the
used parameters are reasonable.
It is known that the stability of a crystal structure is
correlated to its cohesive energy23, which is often defined
as the work that is needed when the crystal is decom-
posed into single atoms. Hence, the lower the cohesive
energy is, the more stable the crystal structure is23. In the
present study, the cohesive energies (Ecoh) of the AlCu3,
AlCu2Zr and AlZr3 crystal cells can be calculated by:
E
E N E N E N E
N N Ncoh
ABC tot A atom
A
B atom
B
C atom
C
A B C
=
− − −
+ +
( )
(1)
where Etot is the total energy of the compound at the
equilibrium lattice constant, and Eatom
A , Eatom
B , Eatom
C are
the energies of the isolated atoms A, B and C in the
freedom states. NA, NB and NC refer to the numbers of
A, B and C atoms in each unit cell. The energies of the
isolated Al, Cu and Zr atoms are –0.276 eV, –0.254 eV
and –2.054 eV, respectively. The cohesive energies
(Ecoh) per atom of all the crystal or primitive cells are
calculated from Eq. (1), and the results of the
calculations are listed in Table 2. From the calculated
values we find that the cohesive energy of AlZr3 is
2.237 eV and 1.413 eV per atom lower than that of
AlCu3 and AlCu2Zr, respectively. Therefore, of the three
phases, the AlZr3 phase has the highest structural
stability, followed by AlCu2Zr and finally the AlCu3.
This means that for the AlZr3, AlCu2Zr, and AlCu3
alloys the structural stability is higher with increasing
amounts of Zr in the crystal.
Table 2: Total energy Etot, cohesive energy Ecoh and formation energy
H of AlCu3, AlCu2Zr and AlZr3
Tabela 2: Celotna energija Etot, kohezivna energija Ecoh in tvorbena
energija H za AlCu3, AlCu2Zr in AlZr3
Compound Etot /eV peratom
Ecoh /eV per
atom
H/eV per
atom
AlCu3 –3.897 –3.637 –0.177
AlCu2Zr –5.261 –4.551 –0.359
AlZr3 –7.574 –5.964 –0.307
In order to compare the alloying abilities of the
present compounds, we calculate the formation energy
H, which can be given by:
ΔH
E N E N E N E
N N NABC
tot A solid
A
B solid
B
C solid
C
A B C
=
− − −
+ +
( )
(2)
where Esolid
A , Esolid
B , Esolid
C are the energies per atom of the
pure constituents A, B and C in the solid states, res-
pectively. And the other variables are as defined for Eq.
(1). If the formation energy is negative, the formation of
a compound from its elements is usually an exothermic
process. Furthermore, the lower the formation energy is,
the stronger the alloying ability is, and the more stable
the crystal structure is23. The calculated energies of Al,
Cu and Zr in their respective crystals are –3.696 eV,
–3.728 eV, –8.457 eV. The calculated results of these
compounds are also listed in Table 2. It is clear that all
R. CHENG, X. Y. WU: STRUCTURAL STABILITY AND ELECTRONIC PROPERTIES OF AlCu3, AlCu2Zr AND AlZr3
66 Materiali in tehnologije / Materials and technology 47 (2013) 1, 65–69
Table 1: Calculated and experimental lattice parameters a (nm), equilibrium volume V0 (nm3), bulk modulus B0 (GPa) and the pressure derivation
of the bulk modulus B’0 for AlCu3, AlCu2Zr, AlZr3
Tabela 1: Izra~unani in eksperimentalno dolo~eni mre`ni parametri a (nm), ravnote`ni volumen V0 (nm3), modul pri stiskanju B0 (GPa) in
izpeljava modula iz tlaka B’0 za AlCu3, AlCu2Zr, AlZr3
AlCu3 AlCu2Zr AlZr3
Present. Expt. Present. Expt. Present. Expt.
a/nm 0.3693 0.3607 19 0.6256 0.6216 21 0.4381 0.4392 20
V0/nm3 50.358 · 10–3 – 244.805 · 10–3 240.210 · 10–3 21 84.110 · 10–3 84.700 · 10–3 20
B0/GPa 131.010 – 128.600 – 100.800 101.4 7
B'0 4.47 – 4.280 – 3.48 3.33 7
the H is negative, which means that the structure of
these compounds can exist and be stable. A further
comparison and analysis showed that the alloying
abilities of AlCu2Zr were much stronger than AlCu3 and
AlZr3. It should be noticed that the alloying ability of
AlZr3 was higher than that of the AlCu3 alloy.
3.2 Elastic properties
The density-functional theory has become a powerful
tool for investigating the elastic properties of materials
(in the limit of zero temperature and in the absence of
zero-point motion). For a given crystal it is possible to
calculate the complete set of elastic constants by apply-
ing small strains to the equilibrium unit cell and deter-
mining the corresponding variations in the total energy.
The necessary number of strains is imposed by the
crystal symmetry24. For a material with cubic symmetry,
there are only three independent elastic constants, C11,
C11 and C11. The strain tensor is given by:

  
  
  
=
⎛
⎝
⎜
⎜⎜
⎞
⎠
⎟
⎟⎟
11 12 13
21 22 23
31 32 33
(3)
In the present study we applied three kinds of strains
(N) (N = 1, 2, 3) so as to obtain the elastic constants, and
they are listed in Table 3. The first strain is a volume-
conserving tetragonal deformation along the z axis, the
second refers to a uniform hydrostatic pressure, and the
last one corresponds to a volume-conserving orthorhom-
bic shear24. The elastic strain energy was given by:
U
E
V
C e eij i j
ji
= =
==
∑∑Δ
0 1
6
1
61
2
(4)
where E = Etotal(V0,) – Etotal(V0,0) is the total energy
variation between the deformed cell and the initial cell,
V0 is the equilibrium volume of the cell, Cij is the elastic
constant and  is the deformation added to the equili-
brium cell. The elastic strain energy is also listed in Tab-
le 3. For each kind of lattice deformation, the total ener-
gy has been calculated for different strains  = ±0.01n
(n = 0  2). By means of a polynomial fit, we extracted
three values of the second-order coefficients, corres-
ponding to 3(C11 – C12), 3(C11 + 2C12)/2 and 2C44,
respectively, the elastic constants C11, C12 and C44 were
obtained25, and the results are shown in Table 4. From
Table 4 we can see that our calculation results agree
well with the experimental data or other first-principle
calculations. These elastic constants satisfy the requi-
rement of mechanical stability for cubic crystals24: (C11
– C12) > 0, C11 > 0, C44 > 0, (C11 + 2 C12) > 0. This
shows that AlCu3, AlCu2Zr and AlZr3 have a stable
structure. The average bulk modulus is identical to the
single-crystal bulk modulus, i.e., B = (C11 + 2 C12)/3.
Interestingly, we noted that the bulk modulus calculated
from the values of the elastic constants is in good
agreement with the one obtained through fitting to the
Birch-Murnaghan equation of state (B0), giving a con-
sistent estimation of the compressibility for these com-
pounds26.
In order to further validate our results, the elastic
modulus, such as the shear modulus G(GPa), Young’s
modulus E(GPa), Poisson’s ratio v and anisotropy
constant A for a polycrystalline material were also cal-
culated with the single-crystal elastic constants Cij, all of
these elastic moduli are shown in Table 4. In the present
study we adopted Hershey’s averaging method27, which
has been known to give the most accurate relation bet-
ween single-crystal and polycrystalline values for a cubic
lattice28. According to this method, G is obtained by
solving the following equation:
G
C C
G
C C C
G
C C C C
3 11 12 2 44 11 12
44 11 12 11
5 4
8
7 4
8
+
+
−
−
−
−
−
( )
( )( +
=
C12
8
0
)
(5)
The calculated shear moduli G for AlZr3 are the
largest, while the quantities for AlCu2Zr are less than for
AlCu3.
Pugh29 found that the ratio of the bulk modulus to the
shear modulus (B/G) of polycrystalline phases can
predict the brittle and ductile behavior of the materials.
A high and low value of B/G are associated with ductility
and brittleness, respectively. The critical value which
separates ductility from brittleness is about 1.75. From
B/G calculated in Table 4 we can see that all the B/G
ratios are larger than 1.75. Therefore, AlCu3, AlCu2Zr
and AlZr3 have good ductility. In contrast, the biggest
B/G ratio for AlCu2Zr indicates that AlCu2Zr is of very
good ductility in these three Al-based alloys. AlCu3 has
an intermediate ductility, while AlZr3 has the worst
ductility.
Besides B/G, the Young’s modulus E and the
Poisson’s ratio v are important for technological and
engineering applications. The Young’s modulus is used
to provide a measure of the stiffness of the solid, i.e., the
larger the value of E, the stiffer the material24. According
to Hershey’s averaging method, the Young’s modulus is
defined as: E = 9GB/3(B+G). Based on the calculated
results, we find that AlZr3 has a Young’s modulus that is
18.806 GPa and 24.663 GPa larger than AlCu3 and
AlCu2Zr, respectively. This indicates that the AlZr3 phase
has the highest stiffness, followed by AlCu3 and finally
the AlCu2Zr. In addition, the Poisson’s ratio v has also
Materiali in tehnologije / Materials and technology 47 (2013) 1, 65–69 67
R. CHENG, X. Y. WU: STRUCTURAL STABILITY AND ELECTRONIC PROPERTIES OF AlCu3, AlCu2Zr AND AlZr3
Table 3: The strains used to calculate the elastic constants of AlCu3,
AlCu2Zr and AlZr3, with  = ±0.01n (n = 0  2)
Tabela 3: Napetosti, uporabljene za izra~un konstant elasti~nosti
AlCu3, AlCu2Zr in AlZr3, z  = ±0.01n (n = 0  2)
Strain Parameters (unlisted ij = 0) E/V0 to 0(2)
(1) 11 = 22 = , 33 = [(1+)–2 – 1] 3(C11 – C22)2
(2) 11 = 22 = 33 =  (3/2)(C11 + 2C12)2
(3) 12 = 21 = , 33 = [2(1 – 2)–1] 2C44 2
been used to measure the shear stability of the lattice,
which usually ranges from –1 to 0.5. The greater the
value of the Poisson’s ratio v, the better the plasticity of
the materials. So we can see that AlCu3, AlCu2Zr and
AlZr3 have a better plasticity.
The elastic anisotropy of the crystals has an
important application in engineering materials since it is
highly correlated with the possibility of inducing micro-
cracks24,30. For cubic symmetric structures31, the elastic
anisotropy is defined as A = (2C44 + C12)/C11. For a
completely isotropic material the value of will be 1,
while values smaller or bigger than 1 measuring the
degree of elastic anisotropy24. Interestingly, we note that
the values of A (Table 4) do not deviate far from unity,
suggesting that the present cubic-structure alloys also do
not deviate far from being isotropic. The calculated
results also indicate that AlCu3 is much more anisotropic
than the other two alloys.
3.3 Density of states
For a beter understanding of the electronic charac-
teristic and structural stability, the total density of states
(DOS) for AlCu3, AlCu2Zr and AlZr3 were calculated, as
shown in Figure 1, as well as the partial density of states
(PDOS) of Al, Cu and Zr atoms in these Al-based
intermetallic compounds. Figure 1 has evidence for the
metallic character of these considered AlCu3, AlCu2Zr
and AlZr3 structures because of the finite DOS at the
Fermi level. With regard to the total density of states
curve of AlCu3, it is clear from Figure 1a that the whole
valence band of AlCu3 is located between –7 eV and 9
eV, which is dominated by Cu 3d states and a small
contribution from the 3s and 3p states of Al. The valence
band of AlZr3 (Figure 1c) can be divided into three
areas. The first area is dominated by the valence electron
numbers of Al 3s and Zr 4d states are mostly located
between –7 eV and –5 eV, the second by the Zr 5s and 4d
states located between –4 eV and –3 eV, and the third by
Zr 4d states located between –2.8 eV and 3.0 eV. Both
below and above the Fermi level, the hybridization
between the Al-p states and Zr-d states is strong. Due to
the strong hybridization (or covalent interaction) the
entire DOS can be divided into bonding and anti-bond-
ing regions, and that a pseudogap resides in between.
The characteristic pseudogap around the Fermi level
indicates the presence of the directional covalent bond-
ing. The Fermi level located at a valley in the bonding
region implies the system has a pronounced stability. It is
R. CHENG, X. Y. WU: STRUCTURAL STABILITY AND ELECTRONIC PROPERTIES OF AlCu3, AlCu2Zr AND AlZr3
68 Materiali in tehnologije / Materials and technology 47 (2013) 1, 65–69
Table 4: Calculated elastic constants (GPa) and elastic modulus (bulk modulus B (GPa), shear modulus G (GPa), Young’s modulus E (GPa),
Poisson’s ratio v and anisotropy constant A) of AlCu3, AlCu2Zr and AlZr3
Tabela 4: Izra~unane konstante elasti~nosti (GPa) in elasti~ni moduli (modul pri stiskanju B (GPa), stri`ni moduli G (GPa), Youngovi moduli E
(GPa), Poissonov koli~nik v in konstante anizotropije A) za AlCu3, AlCu2Zr in AlZr3
Compound C11 C12 C44 B G B/G E v A reference
AlCu3 150.707 120.565 81.880 130.612 43.593 2.996 117.686 0.350 1.887 this study
176.000 117.400 92.400 136.900 49.600 132.800 0.340 5
AlCu2Zr 157.504 115.305 62.685 129.371 41.237 3.137 111.829 0.356 1.528 this study
AlZr3 148.653 79.387 70.834 102.476 53.400 1.919 136.492 0.278 1.487 this study
163.800 79.300 86.500 107.670 6
Figure 1: The total and partial density of states (DOS) of: a) AlCu3
crystal cell, b) AlCu2Zr crystal cell, c) AlZr3 crystal cell. The vertical
dot line indicates the Fermi level.
Slika 1: Skupna in parcialna gostota stanj (DOS) za: a) kristalno
celico AlCu3, b) kristalno celico AlCu2Zr, c) kristalno celico AlZr3.
Navpi~na pik~asta linija prikazuje Fermijev nivo.
also generally considered that the formation of covalent
bonding would enhance the strength of the material in
comparison with the pure metallic bonding32. According
to the covalent approach, the guiding principle is to
maximize the bonding. Therefore, for a series of com-
pounds having the same structure, the greater the
occupancy in the bonding region the higher the
stability33. It is indeed seen that the structural stability
increases from AlCu3 to AlZr3. For AlCu2Zr (see Figure
1b) it is clear that the main bonding peaks between
–6 eV and –2 eV are predominantly derived from the Cu
3d orbits, while the main bonding peaks between the
Fermi level and 3 eV predominantly derived from the Zr
4d orbits. It should be noted that the phase stability of
intermetallics depends on the location of the Fermi level
and the value of the DOS at the Fermi level, i.e.
N(EF) 34,35. A lower N(EF) corresponds to a more stable
structure. The value of the total DOS at the Fermi level is
3.64 states per eV for AlZr3, and the value of the total
DOS at the Fermi level is 5.74 states per eV for AlCu2Zr.
Therefore, AlZr3 has a more stable structure in these
three Al-based intermetallics. This is in accordance with
the calculation of cohesive energy.
4 CONCLUSIONS
In summary, using the first-principles method we
have calculated the alloying stability, the electronic
structure, and the mechanical properties of AlCu3,
AlCu2Zr and AlZr3. These intermetallics have a strong
alloying ability and structural stability due to the
negative formation energies and the cohesive energies. In
particular, AlCu3 is much more anisotropic than the other
two intermetallics. The valence bonds of these inter-
metallics are attributed to the valence electrons of the Cu
3d states for AlCu3, Cu 3d and Zr 4d states for AlCu2Zr,
and Al 3s, Zr 5s and 4d states for AlZr3, respectively, and
the electronic structure of the AlZr3 shows the strongest
hybridization, leading to the worst ductility.
Acknowledgements
This work was financially supported by the National
Natural Science Foundation of China (No. 51175348).
The authors are also grateful to colleagues for their
important contribution to the work.
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R. CHENG, X. Y. WU: STRUCTURAL STABILITY AND ELECTRONIC PROPERTIES OF AlCu3, AlCu2Zr AND AlZr3
Materiali in tehnologije / Materials and technology 47 (2013) 1, 65–69 69

S. KÜÇÜKDERMENCI et al.: SYNTHESIS OF A Fe3O4/PAA-BASED MAGNETIC FLUID ...
SYNTHESIS OF A Fe3O4/PAA-BASED MAGNETIC FLUID
FOR FARADAY-ROTATION MEASUREMENTS
SINTEZA MAGNETNE TEKO^INE NA OSNOVI Fe3O4/PAA ZA
MERITVE FARADAYEVE ROTACIJE
Serhat Küçükdermenci1,3, Deniz Kutluay1,2, Ýbrahim Avgýn1
1Department of Electrical and Electronics Engineering, Ege University, Bornova 35100, Izmir, Turkey
2 Departments of Electronics and Communications Engineering, Izmir University, Uckuyular 35290, Izmir, Turkey
3Department of Electrical and Electronics Engineering, Balikesir University, Campus of Cagis, 10145 Balikesir, Turkey
serhat.kucukdermenci@ege.edu.tr
Prejem rokopisa – received: 2012-07-24; sprejem za objavo – accepted for publication: 2012-08-27
Highly water-soluble Fe3O4/PAA (polyacrylic acid) nanoparticles (NPs) were synthesized with the high-temperature hydrolysis
method. We report the first demonstration of Faraday rotation (FR) for a magnetic fluid (MF) synthesized with this novel
method. The experiments were performed in the DC regime (0–6 · 10–2 T) at room temperature for 14 concentrations from 1.8
mg/ml to 5 mg/ml. The maximum rotation was recorded as 0.96° cm–1 for 3.33 mg/ml and this is called the critical
concentration (CCRITICAL). It was found that the rotation tends to decrease when the concentration is higher than CCRITICAL. The
MF behavior for FR is discussed with respect to substructure interactions (particle-particle, chain-chain). This work provides a
new insight for the FR investigations of MFs including highly water-soluble magnetic NPs.
Keywords: magnetic fluids, Fe3O4 nanoparticles, magneto-optic response, Faraday rotation
V vodi dobro topni nanodelci Fe3O4/PAA (poly acrylic acid) so bili sintetizirani po metodi visokotemperaturne hidrolize.
Poro~amo o prvi demonstraciji Faradayeve rotacije (FR) za magnetno raztopino (MF), sintetizirano po tej novi metodi.
Preizkusi so bili izvr{eni v DC-re`imu (0–6 · 10–2 T) pri sobni temperature za 14 razli~nih koncentracij od 1,8 mg/ml do 5
mg/ml. Najve~ja rotacija je bila ugotovljena kot 0,96° cm–1 pri 3,33 mg/ml, kar smo imenovali kriti~na koncentracija (CCRITICAL).
Ugotovljeno je bilo, da se rotacija zmanj{uje, ~e je koncentracija vi{ja od CCRITICAL. Lastnost MF za FR je bila obravnavana z
interakcijami podstrukture (delec-delec, veriga-veriga). To delo ponuja nov pogled v {tudij FR magnetnih raztopin (MFs),
vklju~no z v vodi dobro topnimi nanodelci (NPs).
Klju~ne besede: magnetna teko~ina, nanodelci Fe3O4, magnetnoopti~ni odgovor, Faradeyeva rotacija
1 INTRODUCTION
Magnetic NPs have unique colloidal, magnetic and
optical properties that differ from their bulk counter-
parts.1–8 Core-shell NPs have been a topic of great
interest due to their potential use in biology9,10, ima-
ging11, medicine12–14 and DNA separation.15,16 Colloidal
suspensions of magnetic NPs can self-assemble into
ordered structures. The ability to manipulate this
assembly with external tuning parameters such as the
field, the temperature and the concentration is essential
for developing new stimuli-responsive materials.
MFs, also named ferrofluids, are colloidal suspen-
sions of magnetic NPs that have both characteristics –
the fluidity of liquids and the magnetism of solid
magnetic materials. Several applications of MFs have
recently been introduced, such as a detection system
design for glucose concentration in addition to
optical-device applications.17 It is suitable for fabricating
optical devices such as optical attenuator, light modu-
lator, optical switch, etc., by using the magneto-optic
properties of MFs.18–20
In 2007, the Yin group synthesized novel superpara-
magnetic, magnetite colloidal NPs that can self-assemble
into one-dimensional (1D) particle chains and exhibit
excellent tunable photonic properties.21 A suspension of
these NPs displays tunable colors in the visible range of
the electromagnetic spectrum. The freedom to tune a
diffraction color not only depends on the particle size but
also varies with the strength of an applied external
magnetic field. Since then there has been a widespread
interest in these NPs and their applications. Despite their
tremendous potential in various applications, interesting
fundamental questions referring to their colloidal
crystallization with and without a magnetic field remain
unanswered. Therefore, we report on the first demonstra-
tion of FR for MFs based on these NPs.
FR has been demonstrated in the visible22,23 NIR24,25
and MIR26 regimes for different kinds of ferrofluids.
Experimental investigation on -Fe2O3 NPs FR was done
due to the particle-size dependence. Water-based ferro-
fluid samples are synthesized with the coprecipitation
method followed by a size-sorting process27. The
wavelength and concentration dependence of FR in MFs
was studied by Yusuf et al.28,29 Here we demonstrate that
long-term, stable MFs including highly water-soluble
NPs show FR in a DC regime (0–6 · 10–2 T) at room
temperature. Water-based ferrofluid samples are
synthesized with a novel high-temperature hydrolysis
method.
Materiali in tehnologije / Materials and technology 47 (2013) 1, 71–78 71
UDK 537.622 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)71(2013)
2 MATERIALS AND METHODS
2.1 Materials
Diethylene glycol (DEG, 99.9 %), anhydrous ferric
chloride (FeCl3, 97 %), sodium hydroxide (NaOH, 96 %),
and poly acrylic acid (PAA, Mw = 1800) were purchased
from the Sigma-Aldrich company. Distilled water was
used in all the experiments. All the chemicals were used
as received without further treatment and/or purification.
2.2 Synthesis of water-dispersible Fe3O4/PAA NPs
The polyol method based on the theory that NPs will
be yielded upon heating precursors in a high-boil-
ing-point alcohol at elevated temperature. In this method,
DEG is chosen as the solvent because it can easily
dissolve a variety of polar inorganic materials due to its
high permittivity ( = 32) and high boiling point (246
°C). DEG is not only a solvent but also a reducing agent
in an reaction. Hence, FeCl3 can be used as the only
precursor for synthesizing Fe3O4. PAA is used as the
capping agent, on which the carboxylate groups show a
strong coordination with Fe3+ on the Fe3O4 surface and
the uncoordinated carboxylate groups extend into the
water solution, rendering the particles with high water
dispersibility. A strong coordination of carboxylate
groups with the surface iron cations and the multiple
anchor points for every single polymer chain is an
important factor in creating a robust surface coating of
PAA on magnetite NPs. Therefore, we used PAA as the
capping agent in our synthesis to confer upon the part-
icles high water dispersibility. A mixture of ethanol and
water was used to wash the particles and remove the
unwanted leftover material from the particles.
For the synthesis of Fe3O4 NPs, a NaOH/DEG
solution was prepared by dissolving 100 mmol of NaOH
in 40 ml of DEG at 120 °C under nitrogen for 1h. Then
the light-yellow solution was cooled to 70 °C (the stock
solution A). In a 100-ml, three-necked flask equipped
with a nitrogen inlet, a stirrer and a condenser, 10 mmol
of FeCl3 and 20 mmol of PAA were dissolved in 41 ml of
DEG under vigorous stirring. The solution was purged
with bubbling nitrogen for 1 h and then heated to 220 °C
for 50 min (the stock solution B). Subsequently, 20 ml of
the NaOH/DEG solution was injected rapidly into the
above solution. The reaction was allowed to proceed for
2 h. The black color of the solution confirms the form-
ation of magnetite NPs. The resultant black product was
repetitively washed with a mixture of ethanol and water
and collected with the help of a magnet. The cycle of
washing and magnetic separation was performed five
times. A one-pot synthesis was done with the Fe3O4/PAA
particles, so no extra process was needed for the surface
modification. A flow chart of the synthesis is shown in
Figure 1.
3 CHARACTERIZATION
3.1 Structural and magnetic characterization of
Fe3O4 /PAA NPs
A powder X-ray diffraction (XRD) analysis was
performed on a Phillips EXPERT 1830 diffractometer
with Cu K radiation. The XRD data were collected over
the range of 10–80° (2 ) with a step interval of 0.02°
and a preset time of 1.6 s per step at room temperature.
Magnetic measurements were carried out using a Lake-
Shore 7400 (Lakeshore Cryotronic) vibration sample
magnetometer (VSM) at 300 K. Particle sizes of NPs
were measured using a Zetasizer 4 Nano S, dynamic
light scattering instrument (Malvern, Worcestershire,
UK). Light-scattering measurements were carried out
with a laser of the wavelength of 633 nm at the 90°
scattering angle. FTIR spectra were recorded on a KBr
disc on a Perkin Elmer 100 spectrometer.
For the FR experiments we used a Thorlabs model
HGR20 2.0 mW/nm laser source, a GMW Electro-
magnet-Systems model 5403 electromagnet, a Kepco
power-supply model BOP 20-5M, a LakeShore model
455 DSP Gaussmeter, a Stanford research systems model
SR830 DSP lock-in amplifier, a new focus model 2051
photo detector, an ILX Lightwave model OMM – 6810B
optical multimeter and an OMH – 6703B model silicon
power head.
3.2 Experimental setup for a magneto-optical charac-
terization
The magneto-optical-measurement setup is shown in
Figure 2. The measurements of FR were made using an
optical arrangement consisting of a He-Ne laser, a polar-
izer, MR3-2 magneto-optical glass (from Xi’an Aofa
Optoelectronics Technology Inc., China) placed in the
gap between the two poles of the electromagnet for the
S. KÜÇÜKDERMENCI et al.: SYNTHESIS OF A Fe3O4/PAA-BASED MAGNETIC FLUID ...
72 Materiali in tehnologije / Materials and technology 47 (2013) 1, 71–78
Figure 1: Flow chart of the synthesis
Slika 1: Potek sinteze
calibration process, an analyzer, collimating lens and a
power meter. The electromagnet generates a uniform
magnetic field in the sample region. The strength of the
magnetic field can be adjusted by tuning the magnitude
of the supply current and is monitored by a gauss meter.
According to Malus’ law, as the polarized light I0
passes through the transparent magneto-optical material,
the light intensity I can be expressed as:
l l l= = −0
2
0
2cos cos ( )   (1)
where  is the angle of the polarization axes of the
polarizer and analyzer and  is the rotation angle of the
polarized plane of the transmitted light. With respect to
MFs, this can be expressed as:
( )
( )
( )B C
M B
M
VBl B
S
= (2)
where l is the chain length at the magnetic field B in an
MF sample, C is a constant that can be found at a high
field assuming that the saturated chain length has no
change, M is the magnetization of the sample at the
magnetic field B, MS is the saturation magnetization of
the sample and V is the Verdet constant varying with the
wavelength and temperature. Generally, a positive
Verdet constant corresponds to the L-rotation (anti-
clockwise) when the direction of propagation is parallel
to the magnetic field and to the R-rotation (clockwise)
when the propagation direction is anti-parallel. To
obtain the maximum sensitivity of the transmitted
power, especially under very small magnetic fields, the
FR angle  can be identified as zero and the initial
polarization angle of the analyzer is set, with  = 45°,
according to equation (1). It can be described as:
d
d
(cos ( ))
( )
sin cos sin
2
2 2


  = = (3)
With respect to the FR experiments, the sensitivity is
maximum when  is 45° and dI/d = 1 from equation
(3). Therefore, the optical axis of the analyzer is aligned
at an angle of 45° to the optical axis of the polarizer.
4 RESULTS
4.1 Stability of Fe3O4 /PAA-based ferrofluids
NPs in MFs are usually coated with a surfactant
material to prevent agglomeration and provide stabil-
ity.30–33 The size of a magnetic NP, the material con-
centration, the carrier liquid and the surfactant are the
main variations for MFs.
PAA was selected as a surfactant because of its
strong coordination of carboxylate groups with iron
cations on a magnetite surface. An additional advantage
of PAA is that an extension of the uncoordinated
carboxylate groups on polymer chains into an aqueous
solution confers on the particles a high degree of
dispersibility in water.
A well-dispersed nanofluid was prepared as shown in
Figure 3. The particle content is 20 mg for various
concentrations with an addition of distilled water from
4 ml to 17 ml. NPs can still be dispersing well after the
Materiali in tehnologije / Materials and technology 47 (2013) 1, 71–78 73
S. KÜÇÜKDERMENCI et al.: SYNTHESIS OF A Fe3O4/PAA-BASED MAGNETIC FLUID ...
Figure 2: Magneto-optic-measurement setup
Slika 2: Magnetoopti~ni merilni sestav
Figure 3: Pictures of the nanofluids containing Fe3O4/PAA NPs with
various concentrations kept for different times: a) 1 h, b) 2 weeks, c) 4
weeks
Slika 3: Posnetki nanoteko~in z Fe3O4/PAA-nanodelci z razli~no kon-
centracijo po zadr`anju: a) 1 h, b) 2 tedna, c) 4 tedne
nanofluid has been kept standing still for more than
4 weeks and no sedimentation is observed for any of the
samples. Due to its long-term stability, this kind of MFs
is an ideal candidate for optical devices.
4.2 Physical and magnetic properties of Fe3O4 /PAA
NPs analyzed with XRD, DLS, TEM and FTIR
The crystal structure of the sample was confirmed
with an X-ray diffraction (XRD) analysis as shown in
Figure 4. The (220), (311), (400), (422), (511), and
(440) diffraction peaks observed on the curves can be
indexed to the cubic spinel structure, and all the peaks
were in good agreement with the Fe3O4 phase (JCPDS
card 19-0629).
The average diameter of the particles obtained with a
dynamic light scattering (DLS) analysis is ~10 nm (Fig-
ure 5). As seen in the typical transmission electron
microscopy (TEM) images of Fe3O4/PAA in Figure
6a-b, the average diameter of Fe3O4/PAA was about 8.16
nm Figure 6c. The average diameter was obtained by
measuring about 100 particles.
The magnetic properties of NPs are shown in Figure
7 as measured at 300 K with a vibrating sample mag-
S. KÜÇÜKDERMENCI et al.: SYNTHESIS OF A Fe3O4/PAA-BASED MAGNETIC FLUID ...
74 Materiali in tehnologije / Materials and technology 47 (2013) 1, 71–78
Figure 5: Size distributions of magnetic NPs obtained with DLS
Slika 5: Razporeditev velikosti magnetnih delcev (NPs), vzeto iz DLS
Figure 4: X-ray powder-diffraction pattern for PAA-coated Fe3O4
NPs. The peak positions and relative intensities recorded in the lite-
rature for bulk Fe3O4 samples are indicated by vertical bars.
Slika 4: Rentgenska difrakcija prahu za s PAA pokritimi Fe3O4-nano-
delci. Pozicija vrhov in v literaturi zapisane relativne intenzitete za
osnovne Fe3O4-delce so prikazane z navpi~nimi ~rtami.
Figure 7: Hysteresis loop of superparamagnetic particles at room
temperature
Slika 7: Histerezna zanka superparamagnetnih delcev pri sobni tem-
peraturi
Figure 6: Representative TEM images of magnetite CNCs with: a)
20-nm scale bar, b) 50-nm scale bar, c) the average diameter of
Fe3O4/PAA
Slika 6: Reprezentativni TEM-posnetki magnetitnih CNCs: a) merilo
20 nm, b) merilo 50 nm, c) povpre~ni premer Fe3O4/PAA
netometer (VSM). The saturation magnetization was
determined as 38.8 emu/g. The particles showed no
remanence or coercivity at 300 K, that is, superparamag-
netic behavior.
The stability of PAA capped on Fe3O4 was confirmed
by measuring the Fourier transform infrared spectro-
scopy (FTIR) spectrum on the sample as shown in Fig-
ure 8a. There is a very strong band at around 1718 cm–1
of pure PAA, which is characteristic of the C=O
stretching mode for protonated carboxylate groups. The
three peaks shown in Figure 8b and located at (1566,
1454 and 1406) cm–1 can be assigned to the characteristic
bands of the carboxylate (COO–) groups, corresponding
to the CH2 bending mode, asymmetric and symmetric
C–O stretching modes of the COO– group, respectively.
4.3 Magneto-optic properties of the Fe3O4/PAA-based
magnetic fluid
Figure 9 shows FR versus the magnetic field and
Table 1 indicates the maximum rotations of MFs with
different concentrations. The field and concentration
dependence of FR in MFs was investigated. A 10-mm-
thick cuvette was filled with the liquid of 14 various
concentrations from 1.8 mg/ml to 5 mg/ml. It can be
seen that the rotation increases rapidly with the field at
low fields.
The initial susceptibility  is determined by a linear
magnetic response M =  · H at the field strength H  0
and it depends on the particle concentration of fluids. At
low fields in Figure 9 the initial slope is:
m
C
B
M B
M
Vl Bi
S
= +
( )
( ) (4)
As long as the concentration is decreased, fewer NPs
are found in the medium and this situation causes a
lower magnetization and a shorter chain length at a
particular field. In other words, the volume fraction of
MF and the initial slope are decreased. Black curves
relate to samples 8–14 (Table 1) and their slopes are less
steep than the ones of the first seven samples, because
their magnetizations are lower and the chain lengths are
shorter than those of the first seven samples. It needs to
be pointed out that though the maximum FR of sample 1
(indicated as the yellow curve) is relatively low, its initial
susceptibility is the highest value due to the volume
fraction.
Table 1: Maximum FRs of the samples with different concentrations
Tabela 1: Maksimalna Faradayeva rotacija (FR) pri vzorcih z raz-
li~nimi koncentracijami
Sample Particle Water Concentration Faraday rot.
no mg ml mg/ml Max. degree
1 20 4 5.00 0.49
2 20 5 4.00 0.87
3 20 6 3.33(CCRITICAL)
0.96
4 20 7 2.86 0.83
5 20 8 2.50 0.78
6 20 9 2.22 0.70
7 20 10 2.00 0.65
8 20 11 1.82 0.56
9 20 12 1.67 0.48
10 20 13 1.54 0.44
11 20 14 1.43 0.37
12 20 15 1.33 0.35
13 20 16 1.25 0.30
14 20 17 1.18 0.28
As the volume fraction of NPs increases, dipole
interactions between the particles overcome the thermal
forces more easily. A graphical representation of the
maximum FR can be seen in Figure 10. It is important to
mention that most experimental investigations, based on
optical observations, of the chain formation in MFs are
usually carried out on the samples in the low concentra-
tion regime, where the chain-chain interaction is weak.
S. KÜÇÜKDERMENCI et al.: SYNTHESIS OF A Fe3O4/PAA-BASED MAGNETIC FLUID ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 71–78 75
Figure 9: Applied magnetic-field dependence and concentration
dependence of FR
Slika 9: Odvisnost FR od uporabljenega magnetnega polja in kon-
centracije
Figure 8: FTIR spectrum of: a) pure PAA, b) PAA with carbo-
xylate-capped Fe3O4 NPs
Slika 8: FTIR-spekter: a) ~isti PAA, b) PAA s karboksilatom omejeni
Fe3O4-nanodelci
However, we tried many samples with variable
concentrations to see the effect of higher concentrations.
When we analyze Figure 10, it can be seen that FR
increases with higher concentrations (from 1.18 mg/ml
to 2.26 mg/ml) up to CCRITICAL (3.33 mg/ml). But after
the CCRITICAL (sample 3) value, FR tends to decrease with
higher concentrations (4 mg/ml and 5 mg/ml). According
to Stokes’ law, there is a friction between the moving
particles and the carrier fluids, which depends on the
viscosity of the fluids.34 Thus, the viscosity of MF and
the friction in the medium were high for highly con-
centrated MFs and they affected the activity of the
magnetic particles. It is also important to point out that
the chain-chain interaction becomes much stronger once
the concentration of a sample is increased and may lead
to the closure of some chains or even the curling.35,36
Therefore, it can be said that FR tends to decrease after
CCRITICAL due to the substructure (particle-particle,
chain-chain) interactions.
Figure 11a shows the transmission of unpolarized
light as percentage and Figure 11b shows the trans-
mission difference of unpolarized light under magnetic
fields for each concentration in the 1–6 · 10–2 T region.
The transmitted light difference as percentage in Figure
12b is calculated from the difference between the inten-
sity of the light transmitted in a particular (1–6 · 10–2 T)
magnetic field and the intensity in the zero field for the
samples with different concentrations.
Low transmission responses to magnetic fields were
observed for samples 1 and 2. Except for these samples,
remarkable changes were observed for all the other
samples because of the weakening of the viscosity effect.
It can be said that samples 1 and 2 have a blocking pro-
perty and other samples have a channeling property due
to their concentrations. A decreasing percentage of NPs
in a unit volume indicates a lower magnetization and a
smaller chain length, which play important roles in the
Faraday effect.37
5 DISCUSSION
5.1 Model for Explaining the Experimental Results
There are two physical phenomena for the magneto-
optical effect in MFs in the presence of an external
magnetic field. One is the orientation theory (magnetic
orientation or physical orientation) based on the optical
anisotropy of magnetic particles or its aggregation, and
the other is the formation theory based on the chain for-
mation of the magnetic particles (Figure 12). In the zero
magnetic field the particles are distributed randomly with
S. KÜÇÜKDERMENCI et al.: SYNTHESIS OF A Fe3O4/PAA-BASED MAGNETIC FLUID ...
76 Materiali in tehnologije / Materials and technology 47 (2013) 1, 71–78
Figure 12: Magnetic NP alignment and the chain formation in the
direction of the external magnetic field
Slika 12: Magnetna ureditev nanodelcev in nastajanje verig v smeri
zunanjega magnetnega polja
Figure 10: Graph of the maximum FR of the samples
Slika 10: Graf maksimalne Faradayeve rotacije (FR) vzorcev
Figure 11: a) Transmissivity of Fe3O4 NPs, B = 0 T, b) transmitted
power difference of Fe3O4 NPs at particular fields and with different
concentrations
Slika 11: a) Transmisivnost Fe3O4-nanodelcev, B = 0 T, b) posre-
dovana razlika v mo~i Fe3O4-nano delcev pri dolo~enem polju z
razli~no koncentracijo
no coercivity and remanence forming an isotropic
material. Particles start to coagulate and form chain-like
structures in the direction of the field with the help of an
external magnetic field.38,39
Magnetic field induces NPs to line up or to behave
asymmetrically, introducing anisotropy and resulting in
birefringence. If lattice atoms of a crystal were not
completely symmetrically arrayed, the binding forces on
the electrons would be anisotropic, causing a material to
be circularly birefringent with different indices of refrac-
tion.40 When the applied magnetic field’s direction is
parallel to the light beam, the anisotropy is circular in a
longitudinal configuration. However, the rotation of
polarization does not seem to be linked to the particles’
anisotropy-axis orientation but to the orientation of the
magnetic moments of the particles in the applied field’s
(H) direction.41
When an external magnetic field is applied parallel to
the plane of MF, magnetic particles in the fluid agglome-
rate to form chain-like structures. As the field strength is
further increased, more particles contribute to agglome-
ration and the chains become longer under a higher field.
It has been found that the chain length varies with the
applied magnetic field and with the concentration of the
MF.42 In some experimental conditions, in which the
wavelength of the electromagnetic waves passed through
the sample is very small in comparison with the chain
length, FR is not only governed by magnetization of the
fluid but also affected by the chain formation.43
Additionally, MFs can be diluted magnetically by
passive liquid carriers such as glycerol, ethylene glycol,
diester, isopar M, ethanol or simply distilled water and it
was seen that FR was affected by a change in the con-
centration. Different concentrations of MFs can help us
describe various friction forces among the magnetic NPs.
Carrier fluids significantly influence the response of the
Faraday effect. Consequently, the chain formation is a
crucial parameter of the optical properties of MFs. The
chain lengths in MFs composed of magnetic NPs vary
with respect to concentration.44–46
The positional and magnetic field dependence of the
colloidal assembly of magnetite NPs arise from a sen-
sitive interplay between the local concentrations of the
particles, causing the effect of three types of forces
between the colloidal NPs. These forces are (1) the
hard-sphere repulsion between the particles in contact;
(2) a combination of electrostatic repulsion due to the
presence of the charges on the surface of NPs and
Van-der-Waals attraction; (3) the magnetic dipolar
attraction/repulsion due to the magnetite cores of the
particles. The interaction potential of the third force is
given by:
U r
rab a b
( ) ( )= ⋅
1
2
 

 
 (5)
The dipolar interaction between two magnetic
particles, a and b, depends on the magnitude and direc-
tion of their magnetic moments μi as well as on their
relative position rab. Depending on the particles configu-
ration, the dipolar energy may be repulsive or attractive.
Heinrich and coworkers showed that these forces play a
role in the assembly of magnetic nano-particles.47 When
a magnetic field is applied these particles initially form
chain-like structures. These chains are then arranged into
two-dimensional hexagonally packed sheets. This occurs
by shifting a neighboring chain by a distance of r corres-
ponding to the radius of NP. The chains are formed along
the direction of an external magnetic field.
6 CONCLUSION
In conclusion, high-quality Fe3O4/PAA-based nano-
structures for MF formation were synthesized success-
fully and FR investigations were made for many samples
with variable concentrations. We report the first demon-
stration of FR for MF synthesized with this novel
method. We have demonstrated the FR of a highly
water-soluble MF that was measured to be in the 0–6 ·
10–2 T range in the DC regime. The effects of both
viscosity and chain formation were observed on FR. We
found the maximum FR to be 0.96°/(10 mm) at room
temperature for 3.33 mg/ml. FR was on an increase with
the higher concentrations up to CCRITICAL. It was found
that the rotation begins to decrease again when the
concentration is higher than CCRITICAL. The reason for this
might be the blocking effect that arises from the par-
ticle-particle and chain-chain interactions. The experi-
ment results shed some light on the role of agglome-
ration and chain formation in FR. Taking into account
the flexibility of the liquid form (including the long-term
stability and no-sedimentation property) in order to
predict its optical behavior correctly, and the low-mag-
netic-field requirements, these fluids can be exploited for
the fabrication of a wide range of applications in
magneto-optics.
Acknowledgements
The authors wish to thank the Dokuz Eylul Uni-
versity, the Center for Fabrication and Application of
Electronic Materials (EMUM) for providing technical
support. We are grateful for the many helpful discussions
with Dr. Erdal Çelik and Dr. Ömer Mermer.
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S. KÜÇÜKDERMENCI et al.: SYNTHESIS OF A Fe3O4/PAA-BASED MAGNETIC FLUID ...
78 Materiali in tehnologije / Materials and technology 47 (2013) 1, 71–78
J. DOBROVSKÁ et al.: A NEW APPROACH TO EVALUATING THE CHEMICAL MICRO-HETEROGENEITY ...
A NEW APPROACH TO EVALUATING THE CHEMICAL
MICRO-HETEROGENEITY OF A CONTINUOUSLY CAST
STEEL SLAB
NOV NA^IN OCENJEVANJA KEMIJSKE MIKROHETEROGENOSTI
KONTINUIRNO ULITIH SLABOV
Jana Dobrovská1, Franti{ek Kavi~ka2, Vìra Dobrovská1, Karel Stránský2,
Hana Francová1
1Faculty of Metallurgy and Materials Engineering, VSB-Technical University of Ostrava, 17. listopadu 15, 708 33 Ostrava, Czech Republic
2Faculty of Mechanical Engineering, Brno University of Technology, Technicka 2, Brno 616 69, Czech Republic
jana.dobrovska@vsb.cz
Prejem rokopisa – received: 2012-08-01; sprejem za objavo – accepted for publication: 2012-09-14
The paper deals with a new approach to measuring and evaluating the chemical micro-heterogeneity of the elements in solidified
poly-component metallic systems. The original approach is based on experimental measurements made on the samples taken
from characteristic places in a casting and the subsequent application of an original mathematical model for determining the
element-distribution profile, characterizing the most probable distribution of an element concentration in the frame of a
dendrite, and an original mathematical model for determining the effective partition coefficients of these elements in the
structure of the analyzed alloy. The paper also describes an application of this method in the research of the chemical
heterogeneity on a cross-section of a CC steel slab and presents the selected results (indices of the heterogeneity and effective
partition coefficients of seven analyzed elements) characterizing the chemical micro-heterogeneity on one-half of the
cross-section of this CC steel slab. The following main results were obtained: (i) the dendritic heterogeneity of the
accompanying elements and impurities is comparatively high; (ii) all the analyzed elements segregate during the solidification
into an inter-dendritic melt, and their partition coefficient is smaller than one; (iii) the effective partition coefficients calculated
in this new way inherently include both the effect of segregation in the course of an alloy solidification, and the effect of the
homogenization occurring during the solidification as well as during the cooling of an alloy.
Keywords: micro-segregation, effective partition coefficient, continuous casting, steel
^lanek obravnava nov na~in merjenja in ocenjevanja kemijske mikroheterogenosti elementov v ve~komponentnih kovinskih
sistemih. Izviren na~in temelji na eksperimentalnih meritvah, narejenih na vzorcih, vzetih iz zna~ilnih mest ulitka, in uporabi
izvirnega matemati~nega modela za dolo~anje profila razporeditve elementov, kar omogo~a oceno najbolj verjetne razporeditve
koncentracije elementov v dendrite in dolo~anje efektivnega koeficienta razporeditve teh elementov v strukturi analizirane
zlitine. ^lanek opisuje tudi uporabo te metode pri raziskavi kemijske heterogenosti pre~nega prereza kontinuirno ulitega slaba in
predstavlja izbrane rezultate (indeks heterogenosti in koeficiente razporeditve sedmih analiziranih elementov) zna~ilne kemijske
mikroheterogenosti polovice prereza kontinuirno ulitega slaba. Glavni rezultati so: (i) heterogenost spremljajo~ih elementov in
ne~isto~ je razmeroma velika; (ii) vsi analizirani elementi izcejajo med strjevanjem v meddendritno talino, njihov koeficient
razporeditve je manj{i od ena; (iii) na nov na~in izra~unani koeficienti razporeditve vklju~ujejo v sebi oba u~inka, to je u~inek
izcejanja med potekom strjevanja in u~inek homogenizacije, ki se pojavi med strjevanjem, in nadaljnjim ohlajanjem zlitine.
Klju~ne besede: mikroizcejanje, koeficient porazdelitve, kontinuirno ulivanje, jeklo
1 INTRODUCTION
The structure of metallic alloys is one of the factors,
which significantly influence their physical and
mechanical properties. The formation of a structure is
strongly affected by production technology, casting and
solidification of these alloys. Chemical heterogeneities
are a common problem in castings and solidification
processes. A solute segregation either on the macro- or
micro-scale is sometimes the cause of unacceptable
products due to poor mechanical properties of the
resulting non-equilibrium phases.
Micro-segregation refers to a composition variation
within a dendritic solidification structure, which has a
length scale of the order of only a few micrometers. It is
usual to characterize the extent of micro-segregation
using a ranking scheme of randomly sampled electron
micro-analysis data.1–4 Thermodynamic quantities are
often calculated from the measurements of an as-cast
segregation profile, in particular, the partition coefficient.
A well-founded technique is thus imperative for
evaluating the compositional data from an X-ray
microanalysis.5,6
Micro-segregation is caused by the redistribution of a
solute during the solidification, as a solute is generally
injected into the liquid. Its fundamental cause is the
difference between the thermodynamic equilibrium
solubility of the alloy elements in different phases that
coexist in the mushy region during the solidification.
This is combined with the inability of the solid-state
diffusion to fully return the composition to its equili-
brium constant level after the solidification is complete,
due to the short times and small diffusion coefficients
involved.7
Materiali in tehnologije / Materials and technology 47 (2013) 1, 79–83 79
UDK 621.74.047:669.112.221 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)79(2013)
A micro-structure prediction is a very difficult task
requiring computationally intensive modeling methods,
such as the phase field8 and cellular automata.9
The presented paper describes a simple methodology
for determining the chemical heterogeneity in metallic
poly-component systems, which was developed on the
basis of a long-term investigation and mathematical mo-
deling of the segregation processes during the crystalli-
zation of metallic alloys at the authors’ workplaces.
An original approach to determining the chemical
heterogeneity in the structure of a poly-component
system is based on the experimental measurements made
on the samples taken from characteristic places of the
casting. In the selected sections of these samples the
concentration of solutes is determined in regular steps.
The length of a measured line depends on the dendrite-
arm spacing. The line segment, along which the concen-
tration of the elements is measured, should appropriately
intersect a number of these dendritic arms (at least five
and even more). Depending on the chemical hetero-
geneity and the structure of the casting the selected
sections usually have the lengths of 500 μm to 1000 μm,
and the total number of steps, in which the concentration
is determined, is set to 101. The method of the quanti-
tative-energy dispersion (ED) or wave dispersion (WD)
X-ray spectral micro-analysis is used for determining the
concentration of the elements.
After the termination correspondence of the mea-
sured concentrations of the elements and the structure of
the given alloy is documented, the average dendrite-arm
spacing is metallographically determined within the
frame of the measured section.
Further procedure is based on statistical processing of
the concentration-data sets and the application of the
original mathematical model for determining distribution
curves of the dendritic segregation of the elements,
characterizing the most probable element-concentration
distribution in the frame of a dendrite,10 and the original
mathematical model for determining the effective parti-
tion coefficients of these elements in the analyzed alloy.
2 EXPERIMENTAL WORK
A continuously cast steel slab (CC steel slab) with
the cross-section dimensions of 1530 mm × 250 mm was
chosen for the presentation of the results. The chemical
composition of the steel (in mass fractions, w/% ) was
the following: 0.14C; 0.75Mn; 0.23Si; 0.016P; 0.010S;
0.10Cr; 0.050Cu; 0.033Altotal.
After the solidification and cooling of the cast slab, a
transversal band was cut out, which was then axially
divided into halves. Nine samples were taken from one
half for determining the chemical heterogeneity as seen
on Figure 1. The samples had a form of a cube with an
edge of about 20 mm, with recorded orientation of its
original position in the CC slab.
All samples were prepared with the standard
metallographic techniques. On each sample a concen-
tration of seven elements (Al, Si, P, S, Ti, Cr and Mn)
were measured along the line of 1000 μm. The distance
between the measured points was 10 μm. An analytical,
complex JEOL JXA 8600/KEVEX Delta V Sesame and
an ED micro-analysis were used for determining the
concentration distribution of the elements. As an
example, Figure 2 presents the basic concentration
spectra of Mn and Si.
3 RESULTS AND DISCUSSION
The chemical micro-heterogeneity, i.e., the
segregation of individual elements at the distances, the
order of which is comparable to the dendrite-arm
spacing, can be quantitatively evaluated from the basic
statistical parameters of the measured concentrations of
the elements in individual samples: Cav average concen-
tration of an element in the measured section, C
standard deviation of the measured concentration of the
element, Cmin minimum and Cmax maximum concen-
trations of the element in the same measured section of
the sample.
Moreover, it is possible to calculate, from these data,
the indices of dendritic heterogeneity IH of the elements
in the measured section as the ratio between the standard
J. DOBROVSKÁ et al.: A NEW APPROACH TO EVALUATING THE CHEMICAL MICRO-HETEROGENEITY ...
80 Materiali in tehnologije / Materials and technology 47 (2013) 1, 79–83
Figure 2: Basic concentration spectra of Mn and Si (sample 21)
Slika 2: Spekter osnovnih koncentracij Mn in Si (vzorec 21)
Figure 1: Scheme of a sampling from a slab and marking of the sam-
ples
Slika 1: Na~rt vzor~enja in oznake vzorcev iz slaba
deviation C and the average concentration Cav of an
element.
Then the element-distribution profiles can be plotted
according to the Gungor’s method11 from the concen-
tration-data sets measured along the line segment with
the length of 1000 μm. The data plotted as the measured
weight-percent composition versus the number of data
(Figure 2) were put in an ascending or descending order
and the x-axis was converted to the fraction solid (fS) by
dividing each measured data number by the total
measured data number. The element composition versus
the fraction solid, i.e., the element-distribution profile
(the distribution curve of a dendritic segregation) was
then plotted; Figure 3 represents such dependences for
manganese and silicon. The slope of such a curve
(ascending or descending) depended on whether the
element in question enriched the dendrite core or the
inter-dendritic area in the course of the solidification.
From these statistical data it is also possible to
determine the values of effective partition coefficients kef
for each element analyzed on each sample. The original
mathematical model for an effective-partition-coefficient
calculation will be outlined here as follows:
The sequence of the arranged concentrations (Figure
3) was seen as a distribution of concentrations of a
measured element in the direction from the axis (fS = 0)
to the boundary (fS = 1) of an average dendrite. The
effective partition coefficient kef was, in this case, defined
with the relation:
kef (fS) = CS (fS)/CL (fS) (1)
where CS is the solute concentration in the solid and CL
is its concentration in the melt and the argument (fS)
expresses the dependence of both concentrations on the
fraction solid.
A perfect mixing of an element in an interdendritic
melt was then assumed (this assumption is the same as,
e.g., in the Scheil12 and Brody-Flemings1 model of
solidification). It was therefore possible to substitute the
equation (1) with the following formula:
kef (i) = Ci/CR (i) (2)
where Ci is the concentration in the i-th point of a
sequence (i.e., in the i-th point of the curve in Figure 3)
and CR(i) is the average concentration of the element i
in the residual part of the curve (i.e., for fS  i, 1),
expressed by the relation:
C i
n i
C j
j i
n
R ( ) = − +
⎛
⎝
⎜ ⎞
⎠
⎟ ⋅
=
∑1
1
(3)
where n was the number of the measured points. In this
way it was possible to determine the values of effective
partition coefficients for all i  1, n, i.e., for the entire
curve characterizing the segregation during the solidi-
fication. The effective partition coefficients of all the
analyzed elements were calculated with this original
method. The average values for the determined effective
partition coefficients are listed in Table 1. No segre-
gation occurs when kef = 1; the higher the deviation
from the number 1 is, the higher is the segregation
ability.
The data presented in Tables 1 and 2 make it possible
to evaluate the dendritic heterogeneity (micro-hetero-
geneity) of the elements in individual samples, and also
in the frame of the whole analyzed half of the slab’s
cross-section. It is obvious that the dendritic hetero-
geneity of the elements is comparatively high. This is
demonstrated with the index of dendritic heterogeneity
IH. It follows from Table 1 that distinct differences exist
between micro-heterogeneities of individual elements.
The average value of this coefficient for all the analyzed
elements and the whole set of nine samples is given in
J. DOBROVSKÁ et al.: A NEW APPROACH TO EVALUATING THE CHEMICAL MICRO-HETEROGENEITY ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 79–83 81
Figure 3: Distribution profiles of Mn and Si constructed from the
experimentally obtained data according to the Gungor’s method11
(sample 21)
Slika 3: Profil razporeditve Mn in Si, dobljen iz eksperimentalno
izmerjenih podatkov skladno z Gungorjevo metodo11 (vzorec 21)
Table 1: Average values of the heterogeneity index IH and the
effective partition coefficient kef of the elements in individual samples
Tabela 1: Povpre~na vrednost indeksa heterogenosti IH in koeficienta
porazdelitve kef elementov v posameznem vzorcu
Sample
Element
Al Si P S Ti Cr Mn
11 IHkef
1.24
0.32
0.28
0.78
1.22
0.33
1.45
0.26
0.30
0.76
0.22
0.83
0.14
0.88
12 IHkef
1.54
0.24
0.30
0.77
1.12
0.36
1.74
0.20
0.29
0.78
0.27
0.79
0.15
0.88
13 IHkef
1.44
0.27
0.30
0.78
1.25
0.32
1.48
0.26
0.30
0.77
0.29
0.78
0.15
0.88
21 IHkef
1.33
0.29
0.29
0.78
1.58
0.24
1.49
0.25
0.31
0.76
0.24
0.81
0.13
0.89
22 IHkef
1.14
0.35
0.28
0.78
1.31
0.30
1.41
0.27
0.30
0.77
0.26
0.80
0.14
0.88
23 IHkef
1.56
0.24
0.29
0.78
1.34
0.29
1.86
0.18
0.26
0.80
0.28
0.78
0.13
0.89
31 IHkef
1.11
0.37
0.28
0.78
1.22
0.33
2.34
0.18
0.31
0.76
0.23
0.82
0.16
0.87
32 IHkef
1.44
0.27
0.27
0.79
1.16
0.34
1.49
0.25
0.34
0.74
0.25
0.80
0.14
0.88
33 IHkef
1.32
0.30
0.29
0.78
1.24
0.32
1.64
0.22
0.35
0.74
0.26
0.80
0.13
0.89
Table 2. It follows from this table that the dendritic
heterogeneity of the slab decreases in this order of
elements: sulphur, aluminium, phosphor, titanium,
silicon, chromium and manganese, which has the lowest
index of heterogeneity.
The dendritic heterogeneities of the analyzed
elements are also expressed with the values of their
effective partition coefficients for the individual samples
as listed in Table 1 and for the set of samples in Table 2.
It is obvious that pair values of the index of dendritic
heterogeneity and the effective distribution coefficient
for the same element do mutually correspond. The fact is
that the higher the value of a heterogeneity index, the
lower the value of an effective partition coefficient and
vice versa. The lowest value of the effective partition
coefficient is found in sulphur and the highest value is
found in manganese. It follows from Table 2 that an
effective distribution coefficient increases in this order of
elements: S, Al, P, Ti, Si, Cr and Mn. All the analyzed
elements segregate during the solidification into an
inter-dendritic melt, and their partition coefficient is
smaller than one.
For comparison, Table 2 contains also the values of
the partition coefficients found in literature.13–15 It is
obvious that our values of the effective partition
coefficients, calculated according to the original model,
are in good agreement with the data from the literature,
with the exception of sulphur (and titanium). The reason
for this difference is probably the method of calculating
an effective partition coefficient – the value of this
parameter is calculated from the concentration-data set
measured on a solidified and cooled casting. Conse-
quently, the effective partition coefficients calculated in
this way inherently include both the effect of segregation
in the course of alloy solidification and the effect of
homogenization, occurring during the solidification as
well as during the cooling of an alloy.
The presented methodology of an investigation into
the chemical micro-heterogeneity makes it possible to
study and to describe the micro-segregation behavior of
the selected elements in the representative areas of a
steel slab. Since a microprobe was used for the experi-
mental investigation, the results have a high accuracy
(even though an assessment with a microprobe is time-
intensive and costly). The results, acquired in this way,
can also provide the standards for another, faster and
cheaper method for investigating the (micro)hete-
rogeneity of a CC steel slab.16
4 CONCLUSIONS
The following main findings and results were
obtained during an investigation into the chemical micro-
heterogeneity of a CC steel slab:
– the dendritic heterogeneity of the accompanying
elements and impurities is comparatively high;
– all the analyzed elements segregate during the
solidification into an inter-dendritic melt, and their
partition coefficient is smaller than one (concrete
values of the partition coefficients for the analyzed
elements and individual samples are given in Table
1, the average values for all the samples are given in
Table 2);
– the dendritic heterogeneity decreases in the follow-
ing order of elements: S, Al, P, Ti, Si, Cr and Mn;
– the effective partition coefficients calculated in this
new way inherently include both the effect of
segregation in the course of an alloy’s solidification,
and the effect of homogenization, occurring during
the solidification, as well as during the cooling of an
alloy.
Acknowledgements
The authors acknowledge the support of the Czech
Science Foundation, projects No. P107/11/1566.
5 REFERENCES
1 M. C. Flemings, D. R. Poirier, R. V. Barone, H. D. Brody, Micro-
segregation in iron-base alloys, J. Iron Steel Int., 208 (1970), 371–81
2 S. Tin, T. M. Pollock, Phase instabilities and carbon additions in
single-crystal nickel-base superalloys, Mater. Sci. Eng. A, 348
(2003), 111–21
3 M. Ganesan, D. Dye, P. D. Lee, A technique for characterizing
microsegregation in multicomponent alloys and its application to
single-crystal superalloy castings, Met. Mat. Trans A, 36 A (2005),
2191–2204
4 K. Stransky, F. Kavicka, B. Sekanina, J. Dobrovska, V. Gontarev,
Numerical and experimental analyses of the chemical heterogeneity
of a solidifying heavy ductile-cast-iron roller, Mater. Tehnol, 46
(2012) 4, 389–392
5 A. Howe, J. Lacaze, P. Benigni, Some issues concerning experiments
and models for alloy microsegregation, Adv. Eng. Mater., 5 (2003),
37–46
6 F. Xie, X. Yan, L. Ding, F. Zhang, S. Chen, M. G. Chu, Y. A. Chang,
A study of microstructure and microsegregation of aluminum 7050
alloy, Mater. Sci. Eng. A, 355A (2003), 144–53
82 Materiali in tehnologije / Materials and technology 47 (2013) 1, 79–83
J. DOBROVSKÁ et al.: A NEW APPROACH TO EVALUATING THE CHEMICAL MICRO-HETEROGENEITY ...
Table 2: Average values of the measured and calculated quantities in
the set of all the samples
Tabela 2: Povpre~ne vrednosti izmerjenih in izra~unanih koli~in za
sklop vseh vzorcev
Cav ± C IH ± I kef ± k k(ref) 13–15
Al 0.01360.0029
1.352
0.162
0.294
0.046 0.12–0.92
Si 0.19100.0068
0.285
0.011
0.781
0.005 0.66–0.91
P 0.01410.0023
1.270
0.133
0.314
0.035 0.06–0.50
S 0.01360.0030
1.657
0.297
0.232
0.035 0.02–0.10
Ti 0.09510.0032
0.306
0.027
0.765
0.019 0.05–0.60
Cr 0.17580.0076
0.255
0.023
0.799
0.017 0.30–0.97
Mn 0.82320.0169
0.143
0.009
0.873
0.033 0.72–0.90
7 Y. M. Won, B. G. Thomas, Simple Model of Microsegregation
During Solidification of Steels, Met. Mat. Trans A, 32A (2001),
1755–1761
8 J. Rappaz, J. F. Scheid, Existence of solutions to a phase-field model
for the isothermal solidification process of a binary alloy, Math.
Methods Appl. Sci., 23 (2000), 491–513
9 Y. M. Won, K. H. Kim, T. Yeo, K. H. Oh, Effect of Cooling Rate on
ZST, LIT and ZDT of Carbon Steels Near Melting Point, Iron Steel
Inst. Jpn. Int., 38 (1998), 1093–99
10 J. Dobrovská, V. Dobrovská, A. Rek, K. Stránský, Possible ways of
prediction of the distribution curves of dendritic segregation of
alloying elements in steels, Scripta Mat., 38 (1998), 1583–1588
11 M. N. Gungor, A statistically significant experimental technique for
investigating micro-segregation in cast alloys, Metall. Trans. A, 20A
(1989), 2529–2538
12 E. Scheil, Uber die eutektische kristallisation, Z. Metallkd., 34
(1942), 70–80
13 L. [mrha, Solidification and crystallisation of steel ingots, SNTL,
Praha 1983 (in Czech)
14 P. Leví~ek, K. Stránský, Metallurgical defects of steel castings
(causes and their elimination), SNTL, Praha 1984, 257 (in Czech)
15 H. J. Eckstein, Wärmebehandlung von Stahl, VEB, DVG, Leipzig
1971, 139
16 H. Presslinger, S. Illie, A. Pissenberger, E. Parteder, C. Bernhard,
Methods for assessment of slab centre segregation as a tool to control
slab continuous casting with soft reduction, ISIJ Int., 46 (2006),
1845–1851
J. DOBROVSKÁ et al.: A NEW APPROACH TO EVALUATING THE CHEMICAL MICRO-HETEROGENEITY ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 79–83 83

J. ONDROUSKOVA et al.: HEAT-FLUX COMPUTATION FROM MEASURED-TEMPERATURE HISTORIES ...
HEAT-FLUX COMPUTATION FROM
MEASURED-TEMPERATURE HISTORIES DURING HOT
ROLLING
RA^UNANJE TOPLOTNEGA TOKA NA OSNOVI PRETEKLIH
MERITEV TEMPERATURE MED VRO^IM VALJANJEM
Jana Ondrouskova1, Michal Pohanka1, Bart Vervaet2
1Heat Transfer and Fluid Flow Laboratory, Faculty of Mechanical Engineering, Brno University of Technology, Technická 2, 616 69 Brno,
Czech Republic
2Centre for Research in Metallurgy, Technologiepark 903c, 9052 Zwijnaarde, Belgium
ondrouskova@lptap.fme.vutbr.cz
Prejem rokopisa – received: 2012-08-31; sprejem za objavo – accepted for publication: 2012-09-18
Due to the long service life of work rolls it is very important to follow the thermal load, but it is very difficult to measure it. One
option for computing this thermal load is to measure the temperature and to study the thermal load through the heat flux. A
unique work roll was made for testing different process conditions, such as rolling velocity, roll cooling, skin cooling and
reduction. This work roll was tested on a real, hot-rolling, continuous pilot line. Two types of temperature sensors were
embedded in the work roll in order to measure the temperature and these gave very detailed information about the development
of the temperature inside the work roll. A time-dependent heat flux was computed using an inverse heat-conduction task with a
detailed numerical model. The surface-temperature history was also obtained from this computational model. These boundary
conditions give detailed information about the influence of different process conditions and allow a computation of the
temperature field in the work roll. The paper describes the measuring equipment, details of the used temperature sensors, the
inverse heat-conduction task for computing the thermal-surface boundary conditions and the results obtained from hot-rolling
conditions.
Keywords: heat flux, hot rolling, roll cooling, inverse heat-conduction problem, surface temperature
Zaradi dolge dobe trajanja delovnih valjev je pomembno sledenje njihovih toplotnih obremenitev, ki jih je te`ko izmeriti. Ena od
opcij izra~una termi~nih obremenitev je merjenje temperature in {tudij toplotne obremenitve s toplotnim tokom. Izdelan je bil
valj za preizku{anje razli~nih procesnih pogojev, kot so hitrost valja, ohlajanje valja, ohlajanje skorje valja in odvzem. Ta
delovni valj je bil preizku{en na realni kontinuirni vro~i valjarni{ki pilotni liniji. V delovni valj sta bili vgrajeni dve vrsti
temperaturnih senzorjev, da bi izmerili temperature, in obe sta dali zelo podrobno informacijo o poteku temperature v njem.
Izra~unana je bila ~asovna odvisnost toplotnega toka z natan~nim matemati~nim modelom inverznega toka toplote. Iz tega
ra~unskega modela je bila dobljena tudi zgodovina temperature povr{ine. Ti mejni pogoji so dali natan~no informacijo o u~inku
razli~nih parametrov procesa in omogo~ili izra~un temperaturnega polja v delovnem valju. V ~lanku je predstavljena merilna
oprema, detajli uporabljenih senzorjev temperature, inverzno prevajanje toplote za izra~un mejnih termi~nih razmer na povr{ini
in rezultati, dobljeni pri valjanju.
Klju~ne besede: toplotni tok, vro~e valjanje, hlajenje valja, problem inverznega prevajanja toplote, temperatura povr{ine
1 INTRODUCTION
The work roll is repeatedly heated and cooled during
the hot-rolling process. A good knowledge of the tem-
perature field of the work roll and a stress-strain analysis
help us improve the cooling of the work roll and, thus, to
extend its service life. The measurement of temperature
histories inside the work roll during a hot-rolling trial
was reported by Raudensky et al.1 This article describes
the results of the measurements in the pilot mill of CRM,
Gent. The temperature sensors were embedded into the
roll using inserts (as seen on the top of Figure 1). There
were five thermocouples in each roll. An example of the
data from the real measurement is on the bottom of
Figure 1. The thermocouples were placed at two
different depths. Two thermocouples were soldered at a
depth of 0.4 mm (Figure 1) and three were drilled at a
depth of 0.8 mm. A detailed description of the plug with
a drilled shielded thermocouple and the discretization
used for the computational model is given in2.
2 INVERSE HEAT-CONDUCTION PROBLEM
A complex 2D-axis symmetric model was used for
the numerical computation. The model includes the
shielded thermocouple with all its parts and the used
solder. The thermocouple must be taken into account
because the homogeneity of the material is disturbed by
the inserted thermocouple, and, thus, the temperature
profile is also disturbed. One-dimensional sequential
Beck’s approach3 is used for the computing of the heat
fluxes and the surface temperatures. The main feature of
this method is the sequential estimation of the time-
varying heat fluxes and surface temperatures using the
future time-step data to stabilize an ill-posed problem.
The measured temperature history is used as input T* to
minimize the equation:
SSE T Ti i
i m
m f
= −
= +
+
∑ ( )* 2
1
(1)
Materiali in tehnologije / Materials and technology 47 (2013) 1, 85–87 85
UDK 621.77.014.2 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)85(2013)
where m is the current time, f is the number of future
time steps and Ti indicates the computed temperatures
from the forward solver4. SSE denotes the sum of square
errors. The value of surface heat flux q at time m is:
( )
q
T T T
q
m
i i q i
i m
m f
i
i m
m f i
i
m
m
=
−
=
== +
+
= +
+
∑
∑
  
 
 
*
;
0
1
2
1
∂
∂
(2)
where  i is a sensitivity coefficient of the temperature
sensor at the time index i to the heat-flux pulse at time
m. The temperatures Ti q m =0 at the sensor location,
computed from the forward solver, use all the previously
computed heat fluxes without the current one, qm. When
the heat flux is found for time m, the corresponding
surface temperature Tm
surf is computed from the forward
solver. Using this procedure, the whole heat-flux history
and surface-temperature history are computed.
3 RESULTS
The heat fluxes and surface temperatures computed
using an inverse heat-conduction problem are presented
here. The measured data were used for computing.
Figure 2 shows a comparison of the results obtained
from the soldered and drilled sensor. It is obvious that
faster changes can be studied with a thermocouple being
implemented closer to the surface. During the measure-
ments various parameters such as reductions (10 %,
30 % and 50 %, Figure 3), velocities (0.1 m/s, 0.3 m/s
and 0.5 m/s), and roll cooling (on/off) were tested.
J. ONDROUSKOVA et al.: HEAT-FLUX COMPUTATION FROM MEASURED-TEMPERATURE HISTORIES ...
86 Materiali in tehnologije / Materials and technology 47 (2013) 1, 85–87
Figure 1: Cross-section of the rolls with the inserts from HiCr and
HSS materials and the holes for the thermocouple wires; a detail of the
plug with a soldered shielded thermocouple; temperature data of the
two roll cycles (depth of 0.4 mm, speed of 0.1 m/s, reduction of 10 %)
Slika 1: Pre~ni prerez valja z vlo`ki iz materialov HiCr in HSS in
luknje za `ico termoelementov; detajl ~epa s prispajkanim termo-
elementom; podatki o temperaturah pri dveh ciklih valjanja (globina
0,4 mm, hitrost 0,1 m/s, odvzem 10-odstoten)
Figure 2: Comparison of the results obtained from the drilled and
soldered thermocouple
Slika 2: Primerjava rezultatov, dobljenih iz izvrtanega oziroma pri-
spajkanega termoelementa
Figure 3: Influence of the slab reductions on the heat flux computed
from the thermocouple situated 0.8 mm under the roll surface. The
contact with the hot strip is on the top and the cooling of the roll is on
the bottom.
Slika 3: Vpliv redukcije slaba na izra~unan tok toplote, izra~unan z
uporabo termoelementa, name{~enega 0,8 mm pod povr{ino valja.
Zgoraj je prikazan na stiku z vro~im trakom, spodaj pa ohlajanje valja.
4 CONCLUSION
An implementation of shielded thermocouples in a
work roll used in a pilot mill of CRM, Ghent, was
presented. Two types of plugs were used. One type had a
soldered thermocouple very close to the surface and the
other type had a thermocouple placed in a drilled hole
that was 0.8 mm under the surface. An inverse algorithm,
which enables a computation of the surface heat fluxes
and surface temperatures from the measured temperature
history inside a roll, was described. The presented results
show the influence of the distance of the thermocouple
from the investigated surface on the calculated surface-
temperature and heat-flux accuracy. The closer the
thermocouple is to the surface the faster changes can be
investigated. The sensor with a soldered thermocouple
gives a more accurate result as it is closer to the surface;
however, the durability of this sensor is very low and it is
destroyed after several contacts with the hot strip. On the
other hand, the durability of a sensor with a drilled hole
is much longer, but its results are not so accurate because
the distance of the thermocouple from the surface is
longer.
Acknowledgement
The research has been supported within the project
No. CZ.1.07/2.3.00/20.0188, HEATEAM–Multidisci-
plinary Team for Research and Development of Heat
Processes, and partially investigated within the RFCS
project CHILLUB, Advanced method to improve
work-roll life time and surface quality of hot rolled strip
by new coupled oil-free lubrication and chilling, No.
RFSR-CT-2008-00012.
5 REFERENCES
1 M. Raudenský, J. Horsky, A. A. Tseng, J. Weng, Heat Transfer Eva-
luation of impingement cooling in hot rolling of shaped steels, Steel
Research, 65 (1994) 9, 375–381
2 M. Pohanka, M. Raudenský, Determination of heat resistances bet-
ween installed thermocouple and body used for computing heat
transfer coefficients. In: Engineering mechanics, Svratka, Czech
Republic, 2002, 227–228
3 J. Beck, B. Blackwell, C. R. Clair, Inverse heat conduction: ill-posed
problems, Wiley, New York 1985, 308
4 W. J. Minkowycz, E. M. Sparrow, J. Y. Murthy, Handbook of Nume-
rical Heat Transfer, 2nd edition, John Willey & Sons, New Jersey
2006, 968
J. ONDROUSKOVA et al.: HEAT-FLUX COMPUTATION FROM MEASURED-TEMPERATURE HISTORIES ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 85–87 87

M. MOZETI^, A. VESEL: THERMAL DEFUNCTIONALIZATION OF AN OXYGEN-PLASMA-TREATED POLYETHERSULFONE
THERMAL DEFUNCTIONALIZATION OF AN
OXYGEN-PLASMA-TREATED POLYETHERSULFONE
TERMI^NA DEFUNKCIONALIZACIJA POLIMERA
POLIETERSULFONA, OBDELANEGA V KISIKOVI PLAZMI
Miran Mozeti~, Alenka Vesel
Jo`ef Stefan Institute, Jamova cesta 39, 1000 Ljubljana, Slovenia
miran.mozetic@ijs.si
Prejem rokopisa – received: 2012-09-03; sprejem za objavo – accepted for publication: 2012-09-14
The phenomenon of spontaneous decay of oxygen-rich functional groups on the surface of a polymer was studied with
high-resolution X-ray photoelectron spectroscopy (XPS). Samples of smooth poly(ether-sulfone) PES foils were exposed to a
highly non-equilibrium oxygen plasma created in an electrodeless RF discharge. After receiving a dose of about 1 × 1024 m–2
oxygen atoms the surface of the polymer became saturated with oxygen-rich functional groups such as C–O, C=O, COO–. The
XPS survey spectra showed an increase in the oxygen concentration from the original mole fractions 21 % to more than 42 %,
while the high-resolution C1s spectra showed that the carboxyl group prevailed. The samples were heated in the XPS chamber
for different periods and characterized in-situ right after the heat treatment. It is shown that the functional groups decay
spontaneously with time and their concentration decreases below the detection limit after annealing at 200 °C for 8 min.
Keywords: oxygen plasma, polymer, functionalization, defunctionalization, thermal stability, activation, deactivation
Z metodo rentgenske fotoelektronske spektroskopije (XPS) smo raziskovali proces spontanega razpada kisikovih funkcionalnih
skupin na povr{ini polimera. Vzorce folije PES smo izpostavili visoko neravnovesni kisikovi plazmi, ustvarjeni v brezelektrodni
radiofrekven~ni razelektritvi. Po izpostavitvi vzorcev kisikovi plazmi, kjer je bila prejeta doza kisikovih atomov okoli 1 × 1024
m–2, se je povr{ina vzorca nasitila z razli~nimi funkcionalnimi skupinami, kot so C–O, C=O, COO–. Iz preglednih XPS-spektrov
smo ugotovili, da je koncentracija kisika na povr{ini narasla iz prvotnih molskih dele`ev 21 % na 42 %, medtem ko smo iz
visokolo~ljivih spektrov ogljika ugotovili, da prevladuje karboksilna skupina. Plazemsko obdelane vzorce smo nato po kemijski
analizi v posodi XPS in-situ segrevali razli~no dolgo, jih ohladili in {e enkrat pomerili povr{insko sestavo. Ugotovili smo, da
funkcionalne skupine s ~asom spontano razpadajo. Po 8 min gretja pri 200 °C pade njihova koncentracija pod detekcijsko mejo.
Klju~ne besede: kisikova plazma, polimer, funkcionalizacija, defunkcionalizacija, termi~na stabilnost, aktivacija, deaktivacija
1 INTRODUCTION
Polymer materials are nowadays widely used in
practical life and science. The surface properties of
selected polymers are often not adequate so they should
be modified prior to specific applications. The basic pro-
perty of a polymer is its hydrophilicity. Most polymers
are moderately hydrophobic. If it is necessary to make
them more hydrophobic, the surface is functionalized
with nonpolar functional groups, typically fluorine-rich
functional groups.1–4 Together with the high surface
roughness, the polymers can be made superhydrophobic
which means that the contact angle of a water drop
approaches 180°. On the other hand, it is sometimes
necessary to make polymers more hydrophilic. A stan-
dard procedure is the functionalization of a surface with
polar functional groups – usually oxygen-rich functional
groups such as carboxyl, carbonyl, hydroxyl, etc. Surpri-
singly enough, a combination of a high concentration of
these functional groups and a very rough surface, which
is due to an intense etching of a polymer surface,5–8 often
leads to a superhydrophilic character of the polymer.9
Namely, the contact angle of the water drop becomes
immeasurably low. In practice it means that the contact
angle is below a few degrees.
There are many methods for the surface func-
tionalization, but it seems that the nowadays treatment
with non-equilibrium gaseous plasma prevails. Plasma
treatment often assures for rapid functionalization.10
While atmospheric plasmas are increasingly popular,
low-pressure plasmas are still more applicable because
they assure for a rather uniform treatment of the arbitrary
samples with a complex shape. Plasma treatment is
nowadays widely used also on the industrial scale. A
phenomenon that is known to many researchers, but
often neglected in relevant literature, is a slow loss of
hydrophilicity.11–13
Highly hydrophilic properties of polymer materials
are usually obtained by applying weakly ionized, highly
reactive oxygen plasma. There are numerous reports on
the functionalization of different materials with an
oxygen-plasma treatment, including polyethylene
terephthalate9,14,15, polyethersulphone16, polystyrene17,
polyvinylchloride18, polymethyl methacrylate19,
cellulose20,21, etc. While many authors use plasma as a
black box (i.e., they do not measure the plasma parame-
ters but present the results on the functionalization as a
function of discharge parameters), there are also reports
in the literature about a precise determination of both
plasma parameters and the surfaces of plasma-treated
Materiali in tehnologije / Materials and technology 47 (2013) 1, 89–92 89
UDK 533.9:678.7 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)89(2013)
materials.22,23 A decent scientific paper would typically
include a flux of positively charged oxygen as well as
neutral oxygen atoms on the surface of a polymer, and a
characterization with XPS and/or SIMS and AFM and/or
SEM. In fact, quite a few scientific papers include all
these data. There are many papers studying the ageing
effects (especially with the water-contact-angle measure-
ments)9 but, on the other hand, very few papers address
the ageing phenomenon, i.e., a spontaneous decrease in
the concentration of functional groups induced with
thermal effects. In the present paper we address these
phenomena. The decay of the functional groups was
studied in-situ in the XPS chamber in order to minimize
the number of possible effects.
2 EXPERIMENTAL WORK
Samples of the commercially available PES foils with
pretty smooth surfaces were prepared as follows. The
freshly packed foils were unpacked, cut into small pieces
and mounted to a plasma system to be functionalized.
The plasma system has been described to details
elsewhere.9 For the sake of the completeness of the paper
let us just summarize the properties. Plasma is excited
with an electrodeless RF discharge at variable power and
oxygen pressure. In the case of our latest experiment we
used a rather low RF power of about 100 W. At the
pressure of 75 Pa the density of the charged particles was
about 8 × 1015 m–3, while the density of the neutral
oxygen atoms was orders of magnitude larger at a value
of about 4 × 1021 m–3. The samples of PES foil were
exposed to the plasma with such parameters for 3 s. The
resultant flux of the neutral oxygen atoms was about 6 ×
1023 m–2 s–1 and the received dose of the atoms in 3 s was
close to 2 × 1024 m–2. Such a large dose of the neutral
oxygen atoms assures for a saturation of the polymer
surface with the oxygen-rich functional groups as has
been already shown in another paper9.
The plasma-treated samples were exposed to air for a
short time and mounted into the XPS instrument (TFA
XPS Physical Electronics). The XPS (X-ray photo-
electron spectroscopy) measurement was performed as
follows: a plasma-treated sample was first analyzed to
see its chemical composition after the plasma treatment.
Then the sample holder in the XPS chamber was heated,
in a short time, to 200 °C to study the thermal stability of
the functional groups formed during the plasma treat-
ment. The sample was kept at this temperature for a
certain time (1 min and 8 min) and the holder was cooled
down to room temperature before the XPS characte-
rization of the treated sample.
During the XPS measurement the sample was excited
with the X-rays with a monochromatic Al K1,2 radiation
at 1486.6 eV. The photoelectrons were detected with a
hemispherical analyzer positioned at an angle of 45°
with respect to the normal to the sample surface. The
survey-scan spectra were made at a pass energy of
187.85 eV and with a 0.4 eV energy step. The
high-resolution spectra of C1s were made at a pass
energy of 23.5 eV and with a 0.1 eV energy step. The
concentration of the elements was determined using
MultiPak v7.3.1 software from Physical Electronics,
which was supplied together with the spectrometer. The
high-resolution spectra were fitted using the same
software. The curves were fitted with the symmetrical
Gauss-Lorentz functions. A Shirley-type background
subtraction was used. Both the relative peak positions
and the relative peak widths (FWHM) were fixed in the
curve-fitting process.
3 RESULTS
Small pieces of foils were treated with the oxygen
plasma in order to get the surface saturated with the
oxygen-rich functional groups. Several samples were
tested, but the differences between particular samples
treated under the same conditions were minimal. The
XPS survey spectra of the selected samples are shown in
Figure 1. The lowest curve refers to the untreated
sample and the other curves refer to the sample treated
by plasma, the sample treated by plasma and heated at
200 °C for 1 min and the sample treated by plasma and
heated for 8 min at the same temperature. From Figure 1
it can be estimated that the plasma treatment caused an
increase in the oxygen concentration, while the thermal
treatment caused a decrease in the oxygen concentration
on the plasma-treated samples. The quantitative values of
the surface composition for all four cases are presented
in Table 1. As expected, the concentration of oxygen on
the plasma-treated sample is much larger than on the
original sample. The O/C ratio is increased by almost a
factor of 3. The values of the oxygen concentration on
the samples treated by plasma and then heated in the
M. MOZETI^, A. VESEL: THERMAL DEFUNCTIONALIZATION OF AN OXYGEN-PLASMA-TREATED POLYETHERSULFONE
90 Materiali in tehnologije / Materials and technology 47 (2013) 1, 89–92
Figure 1: Comparison of the XPS survey spectra of an untreated
polymer, a polymer treated in oxygen plasma for 3 s, a polymer
treated in plasma and heated at 200 °C for 1 min and the one heated at
200 °C for 8 min
Slika 1: Primerjava preglednih spektrov XPS neobdelanega polimera,
polimera obdelanega v plazmi 3 s in plazemsko obdelanega polimera,
ki je bil za 1 min ter 8 min segret na 200 °C
XPS chamber are more interesting. Table 1 shows that
the oxygen concentration drops substantially even after
1 min of heat treatment. After 8 min of heat treatment
the oxygen concentration becomes so low that any spe-
culation about the ratio between O and C for these
samples and the untreated sample would not be fair.
Figure 2 represents the high-resolution C1s peaks for
the selected samples. From this figure one can conclude
that any difference between the untreated sample and the
sample treated with oxygen plasma and heated for 8 min
is below the detection limit of our device. On the other
hand, the high-resolution C1s peak for the sample treated
in the oxygen plasma for 3 s is completely different. The
quantification of these results is summarized in Table 2.
From this table we can see that the majority of the
oxygen is incorporated into the surface layer of PES in
the form of carboxyl groups. It is interesting that the
decay of a particular functional group during the heat
treatment is not very selective. Within the experimental
error one cannot conclude weather any of the newly
formed functional groups decays faster than the other
groups. It is only possible to conclude that the thermal
decay is pretty effective.
Finally, it is worth mentioning that the original con-
centration of sulphur is preserved after the oxygen-
plasma treatment and also after the heating experiments.
This is an indication that the original structure of the
polymer is preserved and oxygen is bonded practically
only on the carbon atoms of the phenyl ring.
4 CONCLUSIONS
The experiments on the ageing of a plasma-func-
tionalized PES showed a spontaneous decrease in the
functional groups during the heat treatment. The tem-
perature of 200 °C was found suitable for the effective
decay of the functional groups. Since the experiments
were performed in an ultra-high vacuum in the XPS
chamber the decay of the groups cannot be explained
with the environmental effects. The two possible
explanations (either a diffusion of oxygen inside the
polymer or a thermal desorption from the surface) should
be taken into account. However, it is not possible to draw
the final conclusion from the experimental results pre-
sented in this paper.
Acknowledgements
The work was financially supported by the Slovenian
Research Agency within the research project L7-4009
"Functionalization of biomedical samples with thermo-
dynamical non-equilibrium gaseous plasma".
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Polym., 2 (2005), 493–500
2 R. di Mundo, F. Palumbo, R. d’Agostino, Langmuir, 24 (2008),
5044–5051
3 S. Marais, Y. Hirata, C. Cabot, S. Morin - Grognet, M. R. Garda, H.
Atmani, F. Poncin - Epaillard, Surf. Coat. Technol., 201 (2006),
868–879
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252–256
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M. Thiebaut, J. Loureiro, Surf. Coat. Technol., 200 (2005), 26–30
6 M. Mafra, T. Belmonte, F. Poncin - Epaillard, A. S. D. S. Sobrinho,
A. Maliska, Plasma Chem. Plasma Process., 28 (2008), 495–509
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M. MOZETI^, A. VESEL: THERMAL DEFUNCTIONALIZATION OF AN OXYGEN-PLASMA-TREATED POLYETHERSULFONE
Materiali in tehnologije / Materials and technology 47 (2013) 1, 89–92 91
Figure 2: Comparison of the high-resolution C1s spectra of the
untreated and treated PES polymers
Slika 2: Primerjava visokolo~ljivih spektrov ogljika C1s neobdelanega
in obdelanega polimera PES
Table 1: Surface compositions of the untreated and treated PES
polymers
Tabela 1: Povr{inska sestava neobdelanega in obdelanega polimera
PES
Sample C O S O/C
untreated 73.8 21.1 5.1 0.29
treated for 3 s in O2 52.0 42.3 5.7 0.81
after 1 min at 200 °C 61.3 33.4 5.4 0.54
after 8 min at 200 °C 69.8 24.6 5.6 0.35
Table 2: Concentrations of different chemical states of carbon atoms
on the untreated and treated PES polymers
Tabela 2: Dele`i razli~nih kemijskih stanj atomov ogljika na
neobdelanem in obdelanem polimeru PES
Sample
C1
C-C
C-S
C2
C-O
C3
C=O
C4
COO-
untreated 82.7 17.3
treated for 3 s in O2 55.5 24.5 7.3 12.7
after 1 min at 200 °C 71.1 21.3 2.1 5.5
after 8 min at 200 °C 84.2 15.8
11 P. Jannasch, Macromolecules, 31 (1998) 4, 1341–1347
12 J. Larrieu, B. Held, H. Martinez, Y. Tison, Surf. Coat. Technol., 200
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Pichelin, Vacuum, 84 (2010), 90–93
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(2011), 022129-1–022129-11
M. MOZETI^, A. VESEL: THERMAL DEFUNCTIONALIZATION OF AN OXYGEN-PLASMA-TREATED POLYETHERSULFONE
92 Materiali in tehnologije / Materials and technology 47 (2013) 1, 89–92
J. SOBOTOVÁ et al.: DIAGNOSTICS OF THE MICROSTRUCTURAL CHANGES IN SUB-ZERO-PROCESSED VANADIS 6 P/M ...
DIAGNOSTICS OF THE MICROSTRUCTURAL CHANGES
IN SUB-ZERO-PROCESSED VANADIS 6 P/M
LEDEBURITIC TOOL STEEL
DIAGNOSTIKA SPREMEMB MIKROSTRUKTURE PRI KRIOGENI OBDELAVI
P/M LEDEBURITNEGA ORODNEGA JEKLA VANADIS 6
Jana Sobotová1, Peter Jur~i2, Jan Adámek3, Petra Salabová4, Otakar Prikner4,
Borivoj [u{tar{i~5, Darja Jenko5
1Czech Technical University in Prague, Faculty of Mechanical Engineering, Karlovo nám. 13, 121 35 Prague 2, Czech Republic
2Faculty of Material Sciences and Technology in Trnava, Paulínská 16, 917 24 Trnava, Slovakia
3Czech Technical University in Prague, Faculty of Nuclear Sciences and Physical Engineering, Bøehová 7, 115 19 Prague 1, Czech Republic
4Prikner - tepelné zpracování kovù, Martínkovice 279, 550 01, Czech Republic
5Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
jana.sobotova@fs.cvut.cz
Prejem rokopisa – received: 2012-09-11; sprejem za objavo – accepted for publication: 2012-09-19
Specimens made from P/M Vanadis 6 cold-work steel were austenitized, quenched and tempered for various combinations of
the parameters. The selected sets of samples, also in the sub-zero range, were treated at a temperature of –196 °C after
quenching. The microstructure was investigated as a function of the austenitizing temperature and the parameters of the
sub-zero processing using transmission electron microscopy (TEM), high-resolution transmission electron microscopy
(HRTEM) and X-ray diffraction. It was found that the as-quenched microstructure is composed of martensite, retained austenite
and carbides. The sub-zero processing reduced the amount of retained austenite and led to an increase in the tetragonality of the
martensitic lattice. As a result, the hardness of the material was higher by 2 HRc before the tempering of the samples after the
sub-zero processing, but the hardness of the sub-zero-processed material after tempering is about 2.5 HRc lower than that of the
non-sub-zero-processed steel. Based on the facts that the sub-zero-processed steel contained less retained austenite and
an unknown amount of the expected nano-precipitates, we expected it to have a lower capability to manifest the
secondary-hardening effect.
Keywords: P/M cold-work steel, heat treatment, sub-zero treatment, precipitates
Vzorci, izdelani iz jekla za delo v hladnem P/M Vanadis 6, so bili avstenitizirani, kaljeni in popu{~ani pri razli~nih
kombinacijah parametrov. Izbrana skupina vzorcev je bila obdelana v kriogenem podro~ju pri temperaturi –196 °C po kaljenju.
Preiskana je bila mikrostruktura v odvisnosti od temperature avstenitizacije in parametrov kriogene obdelave, s presevno
elektronsko mikroskopijo (TEM), visoko lo~ljivo presevno elektronsko mikroskopijo (HRTEM) in z rentgensko difrakcijo.
Ugotovljeno je, da kaljeno mikrostrukturo sestavljajo martenzit, zaostali avstenit in karbidi. Kriogena obdelava zmanj{a dele`
zaostalega avstenita in povzro~i pove~ano tetragonalnost martenzitne re{etke. Posledica je pove~anje trdote za 2 HRc pred
popu{~anjem kriogeno obdelanega materiala, po popu{~anju tega materiala pa je bila trdota za okrog 2,5 HRc ni`ja v primerjavi
z jeklom brez kriogene obdelave. Na osnovi dejstva, da kriogeno obdelano jeklo vsebuje manj zaostalega avstenita in nepoznano
koli~ino nanoizlo~kov, pri~akujemo njegovo manj{o zmo`nost za pojav sekundarnega utrjevanja.
Klju~ne besede: P/M jeklo za delo v hladnem, toplotna obdelava, kriogena obdelava, izlo~ki
1 INTRODUCTION
Vanadis 6 is a powder-metallurgy cold-work ledebu-
ritic tool steel. It is well known that the materials
manufactured with powder metallurgy exhibit an
extremely refined and homogenous microstructure
compared with that of the conventionally produced
steels. According to these facts and, also due to its
nominal chemical composition, Vanadis 6 has a very
high wear resistance, good toughness and high com-
pressive strength. Based on these characteristics, Vanadis
6 is suitable for cutting blades, rolling planes or tools for
forming operations.
The structure and properties of ledeburitic steels are
determined by the character of the matrix and the type,
quantity, size and distribution of carbides and, therefore,
the demanded tool life is determined by the conditions of
heat treatment. During austenitizing the eutectoid and
a part of the secondary carbides are dissolved in auste-
nite, which results in a high hardness of the material
after quenching. Other carbides, which are not subjected
to dissolution, inhibit the coarsening of the austenite
grains and make the steel wear resistant. Bílek et al.1
reported that there are two types of carbides in Vanadis 6
after austenitizing. The first phase of M7C3 carbides
underwent the dissolution in the austenite to a greater
extent than the second phase of MC carbides, which is
stable up to the temperature of 1150 °C. A previous
publication2 states that a higher austenitizing temperature
results in an increase in the hardness of Vanadis 6 and a
reduction of the three-point bending strength, since the
increased austenitizing temperature results in a grain
coarsening. After quenching, Cr-V ledeburitic steels con-
tain martensite, retained austenite and undissolved carbi-
des. For a more complete martensitic transformation,
Materiali in tehnologije / Materials and technology 47 (2013) 1, 93–98 93
UDK 669.14:621.78 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(1)93(2013)
sub-zero treatment can be inserted between quenching
and tempering.
Recent investigations3–5 have established that sub-
zero treatment does not only substantially reduce the
content of retained austenite but also significantly
modifies the precipitation behaviour of the carbides in
the processed materials. These changes are expected to
have a considerable influence on the mechanical pro-
perties and, subsequently, on the wear resistance of tool
steels. Opinions on the effect of sub-zero processing on
the properties of ledeburitic steels differ. Debdulal et al.3
reported that sub-zero treatment reduces the fracture
toughness of the cold-work tool steel AISI D2 that is
lower than that of the conventionally heat-treated ones.
However, the degree of reduction in fracture toughness
varies with the types of sub-zero treatment; it is the
lowest for deep, cryogenically treated DCT (–196 °C to
–160 °C) specimens and the highest for shallow, cryoge-
nically treated (–160 °C to –80 °C) ones. Fracture-
toughness values decrease, expectedly, with an increase
in the hardness, a reduction of the retained-austenite
content, an increase in the amount of secondary carbides
and their size refinement. According to Oppenkowski4,
further investigations on D2 cold-work tool steel show
that !-carbides can be observed at higher tempering
temperatures, but in some investigations it was
impossible to detect any !-carbides. Opposing results
have also been obtained with respect to the precipitation
of !-carbides. Some investigations found that carbides
precipitated after DCT during tempering, others report
about the precipitation during the DCT process.
Shaohong5 reported that DCT decreases the retained
austenite and increases the hardness and wear resistance
of cold-work die steel. Carbon atoms were segregated
close to dislocations in the DCT process, -carbides
precipitated in the twinned martensite during the process
of tempering.
The aim of this research is to assess the microstruc-
tural changes during the DCT process made to powder-
metallurgical tool steel Vanadis 6.
2 EXPERIMENTAL WORK
The nominal chemical composition of the experi-
mental material Vanadis 6, manufactured with P/M is
shown in Table 1. Details of the experimental heat-treat-
ment program were published elsewhere6. Two types of
the heat-treatment process were performed. The con-
ventional heat treatment involved the following steps:
vacuum austenitizing up to a temperature of 1000 °C or
1075 °C, respectively, and nitrogen gas quenching at a
pressure of 5 bar. For the novel heat-treatment process, a
sub-zero period at a soaking temperature of –196 °C
lasting for 4 h was inserted after quenching and
tempering. No tempering was carried out in order to
highlight the possible changes in the microstructure of
the processed material due to the sub-zero period. A
microstructural analysis was performed using a trans-
mission electron microscope TEM JEOL 200CX at an
acceleration voltage of 200 kV. Thin foils for the TEM
were prepared with an electrolytic jet-polisher (Tenupol)
or with an ion slicer. The content of the retained
austenite was measured with X-ray diffraction. X-ray
patterns were recorded using a Phillips PW 1710 device
with Fe-filtered Co1,2 characteristic radiation. The
detector arm was equipped with a monochromator. The
data were recorded in the range of 20–144° of the
two-theta angle with a 0.05° step and a counting time per
step of 5 s.
Table 1: Chemical composition of the investigated steel in mass
fractions (w/%)
Tabela 1: Kemijska sestava preiskovanega jekla v masni dele`ih
(w/%)
Steel C Si Mn Cr Mo V Fe
Vanadis 6 2.1 1.0 0.4 6.8 1.5 5.4 balance
3 RESULTS AND DISCUSSION
The mechanical properties obtained after different
steps of the heat treatment are summarized in Table 2 2,6.
Although the hardness is approximately the same after
quenching and after quenching and sub-zero processing
at both austenitizing temperatures, the hardness of the
sub-zero-processed material after tempering is by about
2.5 HRc lower than that of the non-sub-zero-processed
steel.
The tempering of the material induced different
behaviours of the non-sub-zero- and sub-zero-processed
steels. Generally, the non-sub-zero-processed steel had a
higher hardness than the sub-zero-processed one. We
assume that, firstly, the occurrence of a high number of
nano-particles and, secondly, a lower share of the
retained austenite are responsible for the deterioration of
the secondary hardenability of the material. This seems
to be logical because the secondary hardenability can be
considered as a comprehensive effect of the tempering of
martensite (hardness decrease), transformation of the
retained austenite to martensite (hardness increase) and
precipitation of carbides (hardness increase). Based on
the facts that the sub-zero-processed steel contained a
lower share of the retained austenite and an expected
large number of nano-precipitates, we anticipated its
capability to manifest the secondary-hardening effect to
be lower.
Figure 1 shows a bright-field image of the micro-
structure of the specimen quenched from 1000 °C with-
out DCT and tempering. The matrix of the material
apparently has a two-phase structure. The main structural
constituent is the twinned martensite, which is typical for
high-carbon steels. The second structural feature of the
matrix is the retained austenite as shown in the dark-field
image of the area, Figure 2. The fact that the material
after quenching, without DCT, contains a large amount
J. SOBOTOVÁ et al.: DIAGNOSTICS OF THE MICROSTRUCTURAL CHANGES IN SUB-ZERO-PROCESSED VANADIS 6 P/M ...
94 Materiali in tehnologije / Materials and technology 47 (2013) 1, 93–98
of retained austenite could have been expected. During
austenitizing some of the carbides are dissolved and the
austenite becomes saturated with carbon and alloying
elements. This induces a lowering of both the Ms and Mf
temperatures of Cr-V ledeburitic steels, whereas Mf lies
well below the freezing point. As a result, the martensitic
transformation is not fully completed. The diffraction
patterns from both structural constituents are shown in
Figures 3 and 4.
After quenching from the austenitizing temperature
of 1000 °C and DCT at –196°C/4 h, the matrix again has
a two-phase structure as seen in Figures 5 and 6. This
means that even the sub-zero processing does not com-
pletely remove the retained austenite from the structure.
This could be evidently expected. From the physical-
J. SOBOTOVÁ et al.: DIAGNOSTICS OF THE MICROSTRUCTURAL CHANGES IN SUB-ZERO-PROCESSED VANADIS 6 P/M ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 93–98 95
Table 2: Hardness and bending strength of differently heat-treated specimens2,6
Tabela 2: Trdota in upogibna trdnost razli~no toplotno obdelanih vzorcev2,6
Austenitizing
temperature
Hardness A
HRc
Sub-zero
processing
Hardness B
HRc
Hardness C
HRc
Bending strenght
MPa
1000 °C 66.0 no 59.0 3564
1075 °C 66.0 no 63.0 2768
1000 °C 65.0 –196 °C/4 h 66.0 56.5 3272
1075 °C 65.5 –196 °C/4 h 67.0 60.5 3087
Hardness A hardness after quenching
Hardness B hardness after quenching and sub-zero processing
Hardness C hardness after complete heat treatment with a double tempering at 530 °C/2 h
Figure 4: Electron-diffraction patterns from retained austenite
Slika 4: Uklonska slika zaostalega avstenita
Figure 2: Microstructure of Vanadis 6 after quenching from 1000 °C
without DCT and tempering – TEM dark-field image
Slika 2: Mikrostruktura Vanadis 6 po kaljenju iz 1000 °C, brez DCT
in popu{~anja (TEM – temno polje)
Figure 1: Microstructure of Vanadis 6 after quenching from 1000 °C
without DCT and tempering – TEM bright-field image
Slika 1: Mikrostruktura Vanadis 6 po kaljenju iz 1000 °C, brez DTC
in popu{~anja (TEM – svetlo polje)
Figure 3: Electron-diffraction patterns from martensite
Slika 3: Uklonska slika martenzita
metallurgy viewpoint the martensitic transformation
cannot be fully completed and after quenching a certain
share of the retained austenite will always remain in the
structure. The problem is how to minimize the content of
the retained austenite. Transmission microscopy, how-
ever, will never precisely quantify the content of the
retained austenite.
A quantification of the amount of the retained auste-
nite was performed with an X-ray diffraction analysis,
the results are in Table 3.
Besides the martensite and the retained austenite, MC
(or M4C3) and also M7C3 carbides were found in the
structure. The contents of these carbides were found to
be independent of the use (or no use) of DCT in the
heat-treatment cycle. This is logical since it is the
austenitizing temperature that plays a substantial role in
the changes in the volume fraction of these carbides. On
the contrary, a large difference was found in the contents
of the retained austenite and martensite. The content of
the retained austenite was almost three times lower for
the DCT material than for the non-DCT material. It is
thus evident that DCT really contributes to a marked
decrease in the amount of the retained austenite in the
structure of the material after the heat treatment with
DCT. At the same time, a higher supersaturation of the
martensite occurred due to DCT – the degree of tetra-
gonality c/a was found to be much higher for the DCT
material.
Besides deformation twins, a dense dislocation
network can also be found in the martensite needles.
This was well apparent at high magnification in Figures 7
and 8. The dislocation network prevents the identifi-
cation of the potential minority phases, so that, at this
stage of the research, the assumed precipitation of
carbide nanoparticles during the DCT, often published in
literature, could not be confirmed. On the other hand, the
presence of such a high number of dislocations is a sign
of a high plastic deformation and an apparently high
stress in the crystal lattice. Plastic deformation can
induce various processes, e.g., self-tempering of marten-
site accompanied by a formation of carbide nano-
particles.
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96 Materiali in tehnologije / Materials and technology 47 (2013) 1, 93–98
Figure 7: Microstructure of Vanadis 6 after quenching from 1000 °C
and DCT without tempering – TEM bright-field image – a detail
Slika 7: Detajl mikrostrukture Vanadis 6 po kaljenju iz 1000 °C in
DCT, brez popu{~anja (TEM – svetlo polje)
Figure 6: Microstructure of Vanadis 6 after quenching from 1000 °C
and DCT without tempering – TEM dark-field image
Slika 6: Mikrostruktura Vanadis 6 po kaljenju iz 1000 °C in DCT,
brez popu{~anja (TEM – temno polje)
Figure 5: Microstructure of Vanadis 6 after quenching from 1000 °C
and DCT without tempering – TEM bright-field image
Slika 5: Mikrostruktura Vanadis 6 po kaljenju iz 1000 °C in DCT,
brez popu{~anja (TEM – svetlo polje)
Table 3: Results of an X-ray diffraction of Vanadis 6 after quenching from the temperature of 1075 °C and sub-zero processing at –196 °C/4 h
Tabela 3: Rezultati rentgenske difrakcije Vanadis 6 po kaljenju iz temperature 1075 °C in kriogeni obdelavi pri –196 °C/4 h
Heat treatment Content of martensite,w/%
Content of retained
austenite, w/%
Lattice parameter of martensite
(nm) degree of tetragonality c/a Other phases
Quenching 49.45 17.76 a = 0.28623, c = 0.29154c/a = 0.101186
MC (M4C3)
M7C3
Quenching + DCT 63.1 6.1 a = 0.28596, c = 0.29234c/a = 0.10223
MC (M4C3)
M7C3
This could explain the assumed higher stresses in the
lattice and the similarly high dislocation density in the
martensite obtained with DCT, Figures 7 and 8.
It was shown that the magnification of a standard
TEM is not sufficient for the identification of the
assumed precipitation of carbide nanoparticles. Accord-
ing to this, we tried to investigate the assumed preci-
pitation of carbides, often recorded in literature, with
HRTEM, Figures 9 and 10.
On the left-hand side of Figure 9 there is a typical
contrast between an amorphous layer and some faint
underlying fringes of the crystal planes. To the right the
contrast passes to Moire fringes due to the super-
imposed, twisted crystal lattices. A Moire pattern is also
in the bottom right-hand corner of the micrograph. In the
central part of the lattice the fringes are not clear,
perhaps due to some internal strains. Their clearest
contrast is in the top right-hand corner. However, the
atomic columns are not parallel to the electron beam and
so no dots, but only fringes, are visible.
In the central part of Figure 10 there is a dot contrast
between the atomic columns, probably in the 111-Fe
projection. The regular pattern in the upper central part
could be a nanotwinned region or a Moire pattern caused
due to the superimposed, twisted crystal lattices or a
faulted nanoparticle. However, the contrast is not clear
enough to allow an unambiguous interpretation.
The expected nano-precipitates could not even be
confirmed by using HRTEM. Neither did the results of
EDS analyses show the differences between the mate-
rials after quenching and after quenching and DCT.
In the next investigation we will turn our attention to
the assessment of microstructural changes after temper-
ing the sub-zero-processed cold-work tool steel Vanadis
6. Such a research might answer the question as to
whether it is possible to identify the nanoparticles in the
structure after a complete heat treatment including DCT,
often addressed in literature.
4 CONCLUSIONS
• The content of the retained austenite was almost three
times lower for the DCT Vanadis 6 than that without
DCT.
• At the same time, a higher supersaturation of marten-
site occurred due to DCT – the degree of tetrago-
nality c/a was found to be much higher for the DCT
material.
• Using TEM it was proved that the dislocation
network prevents the identification of potential
minority phases in the microstructure of the DCT
material.
• The expected nano-precipitates in the microstructure
of the DCT material could not be confirmed, even by
using HRTEM.
• For the next investigation we recommend an assess-
ment of the microstructural changes after tempering
the DCT cold-work tool steel Vanadis 6.
J. SOBOTOVÁ et al.: DIAGNOSTICS OF THE MICROSTRUCTURAL CHANGES IN SUB-ZERO-PROCESSED VANADIS 6 P/M ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 93–98 97
Figure 10: Microstructure of Vanadis 6 after quenching from 1075 °C
and DCT without tempering – HRTEM
Slika 10: Mikrostruktura Vanadis 6 po kaljenju iz 1075 °C in DCT,
brez popu{~anja (HRTEM)
Figure 8: Microstructure of Vanadis 6 after quenching from 1000 °C
and DCT without tempering – TEM dark-field image – a detail
Slika 8: Detajl mikrostrukture Vanadis 6 po kaljenju iz 1000 °C in
DCT, brez popu{~anja (TEM – temno polje)
Figure 9: Microstructure of Vanadis 6 after quenching from 1000 °C
and DCT without tempering – HRTEM
Slika 9: Mikrostruktura Vanadis 6 po kaljenju iz 1000 °C in DCT,
brez popu{~anja (HRTEM)
Acknowledgements
The authors wish to thank the Ministry of Industry
and Trade of the Czech Republic for the financial
support given to the Project TIP FR-TI1/003. The paper
was created within the research MSM6840770021
"Diagnostic of materials" and project OPPA
CZ.2.17/1.1.00/32213 "Qualifications of the academic
workers from CTU in Prague in science and research".
5 REFERENCES
1 P. Bílek, J. Sobotová, P. Jur~i, Evaulation of Structural changes in
Cr-V Ledeburitic Tool Steels Depending on Temperature Austeniti-
zation, Mater. Tehnol., 44 (2011) 4, 33–37
2 P. Jur~i et al., Effect of Sub-zero Treatment on Mechanical Properties
of Vanadis 6 PM Ledeburitic Tool Steel, in Metal 2010: 19th Inter-
national Conference on Metalurgy and Materiále, Roznov pod Rad-
hostem
3 D. Debdulal et al., Influence of sub-zero treatments on fracture
toughness of AISI D2 steel, Materials Science and Engineering A,
528 (2010), 589–603
4 A. Oppenkowski, S. Weber, W. Theisen, Evaluation of factors
influencing deep cryogenic treatment that affect the properties of tool
steels, Journal of Materials Processing Technology, 210 (2010),
1949–1955
5 L. Shaohong et al., Influence of deep cryogenic treatment on micro-
structure on microstructure and evaluation by internal friction of tool
steel, Cryogenics, 50 (2010), 754–758
6 J. Sobotova et al., Structure and properties of sub-zero processed
Vanadis P/M ledeburitic tool steel, in Metal 2011: 20th Anniversary
International Conference on Metalurgy and Materilas, Brno
J. SOBOTOVÁ et al.: DIAGNOSTICS OF THE MICROSTRUCTURAL CHANGES IN SUB-ZERO-PROCESSED VANADIS 6 P/M ...
98 Materiali in tehnologije / Materials and technology 47 (2013) 1, 93–98
M. VUK^EVI] et al.: UTILIZATION OF GEOPOLYMERIZATION FOR OBTAINING CONSTRUCTION MATERIALS ...
UTILIZATION OF GEOPOLYMERIZATION FOR
OBTAINING CONSTRUCTION MATERIALS BASED ON
RED MUD
UPORABA GEOPOLIMERIZACIJE ZA PRIDOBIVANJE
GRADBENEGA MATERIALA NA OSNOVI RDE^EGA BLATA
Mira Vuk~evi}1, Danka Turovi}1, Milun Krgovi}1, Ivana Bo{kovi}1,
Mileta Ivanovi}1, Radomir Zejak2
1University of Montenegro, Faculty of Metallurgy and Technology, 81000 Podgorica, Montenegro
2University of Montenegro, Faculty of Civil Engineering, D`ord`a Va{ingtona bb, 81000 Podgorica, Montenegro
mirav@ac.me
Prejem rokopisa – received: 2012-05-08; sprejem za objavo – accepted for publication: 2012-08-29
In this study the geopolymerization process for obtaining construction materials based on red mud was used. The aim of this
study was to define the most favorable conditions enabling the utilization of the geopolymerization process in the production of
construction materials based on red mud as a by-product of alumina production. For this purpose, the physicochemical and
mechanical properties of the obtained construction (geopolymers) materials were tested. On the basis of the results the optimal
conditions for geopolymerization and the effect of the main synthesis parameters were determined with repect to the satisfactory
mechanical and other properties of the obtained materials. The inorganic polymeric materials produced by the
geopolymerization of red mud developed satisfactory compressive strength, which leads to the conclusion that these materials
may be used in the sector of construction materials.
Keywords: geopolymerization, red mud, compressive strength, metakaolin
V tej {tudiji je bil uporabljen postopek geopolimerizacije za pridobivanje gradbenega materiala na osnovi rde~ega blata. Namen
{tudije je bil opredeliti najugodnej{e razmere, ki bi omogo~ile uporabo geopolimerizacijskega postopka za pridobivanje
gradbenega materiala na osnovi rde~ega blata, ki je stranski proizvod pri proizvodnji aluminija. Za ta namen so bile preizku{ene
fizikalno-kemijske in mehanske lastnosti materiala (geopolimera). Na osnovi dobljenih rezultatov so bile dolo~ene optimalne
razmere za geopolimerizacijo in u~inki glavnih parametrov sinteze s stali{~a zadovoljivih mehanskih in drugih lastnosti.
Anorganski polimerni material, izdelan z geopolimerizacijo rde~ega blata, je imel zadovoljivo tla~no trdnost, kar omogo~a
sklep, da bi bil ta material uporaben kot gradbeni material.
Klju~ne besede: geopolimerizacija, rde~e blato, tla~na trdnost, metakaolin
1 INTRODUCTION
The present work investigates the possibility of using
the red mud from the Bayer process for the production of
construction elements made with the process of geo-
polymerization. The geopolymerization process is based
on the heterogeneous chemical reaction that occurs
between the solid, aluminosilicate-rich materials, and the
highly alkaline silicate solution. The basic part of this
process is the hardening of geopolymers which is based
on the polycondensation reactions of the alkali pre-
cursors formed from a dissolution of active silicates and
aluminosilicate solid materials in an alkali-hydroxide
solution. The polymeric network as a result of the
polycondensation process hardens rapidly acting as a
gluing component1. The geopolymerization is an exo-
thermic reaction that takes place at an atmospheric
pressure and a temperature bellow 100 °C leading to a
formation of compact, solid materials, typical for their
three-dimensional polymer structure. Such materials are
called geopolymers2–4. The first stage in this reaction is
the formation of hydroxyl complexes of silicon and
aluminum with the polymer-bond types of Si-O-Si and
Si-O-Al, followed by the formation of three-dimensional
aluminosilicate networks containing SiO4 and/or AlO4
tetrahedral, alternatively linked through a common oxy-
gen ion. The last stage in the process is the wrapping of
the insoluble solid particles in the geopolymer5–10.
The aim of this investigation is to define the optimal
combination of the relevant parameters that would
enable the use of red mud from the alumina-production
process to be the dominant raw material in combination
with the activator and the binder for the production of
geopolymers. Geopolymerization creates favorable con-
ditions in promoting red mud as the basis for the deve-
lopment of a new class of construction materials,
inorganic polymers – geopolymers. The most important
expectations regarding the construction materials are
good physicochemical and mechanical characteristics,
dimensional stability, a good fire resistance and an
aggressive-environment resistance. However, the pre-
sence of hydroxyl Fe oxides in bauxite (goethite) com-
promises their use in the conventional construction
materials because of their dehydroxylation-hydroxy-
lation activities generating a dimensional instability.
Geopolymerization lowers the level of water absorption
because of the amorphous or semi-crystal structure,
Materiali in tehnologije / Materials and technology 47 (2013) 1, 99–104 99
UDK 66.095.26:691:539.411 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(1)99(2013)
lowers the micro porosity, enables higher values of the
specific mass, compressive strength, etc. The final objec-
tive of this investigation was to define the influence of
the relevant parameters affecting the geopolymerization.
The compressive strength, the apparent density and the
microstructure of polymeric materials were investigated
to define the optimal conditions for a polymeric-material
synthesis.
2 EXPERIMENTAL WORK
2.1 Materials
For the production of construction materials the
following raw materials were used:
• red mud obtained as a byproduct of the Bayer process
of obtaining alumina (Podgorica Aluminum Factory),
• sodium hydroxide of analytical grade (Merck,
anhydrous pellets),
• metakaolin, which initially provides a geopolymeric
system with soluble silicon and aluminum that are
essential for an aluminosilicate-oligomer formation
and the progress of geopolymeriztion,
• sodium-silicate solution (Merck: m(Na2O) : m(SiO2)
= 3.4, w(Na2O) = 7.5–8.5 %, w(SiO2) = 25.5–28.5 %
and d = 1,347 g cm–3),
• deionized water for the synthesis of the polymeric
material.
Red mud originated from the aluminum metallurgical
plant in Podgorica, Montenegro, as a by-product of the
alumina production known for the presence of hydro-
xylation Fe oxides, dried to a constant mass at a tempe-
rature of 105 °C, and then sifted through a sieve with a
hole diameter of  = 1 mm.
Metakaolin is a dehydroxylation product of the
industrial mineral kaolin in the temperature range
between 650 °C and 850 °C. The thermal dehydro-
xylation of kaolin increases its solubility in an alkaline
media and it was performed at 750 °C. The basic
material was mineral kaolin from the Bijele Poljane site
in Montenegro. Metakaolin is a predominantly amor-
phous material with minor crystalline constitutes.
As an alkaline activator of the process of geo-
polymerization, a combination of sodium water glass and
sodium hydroxide was used. The activator solution was
prepared by mixing the previously mentioned compo-
nents 48 h before the geopolymer production. Different
concentrations of NaOH (CNaOH = (3, 7 and 10) mol dm–3)
and a concentration of Si in Na-silica (1, 1.5 and 3.5)
mol dm–3 were used. The levels of substitution of red
mud with metakaolin in the solid phase were in mass
fractions w = (4, 8 and 15) %.
2.2 Experimental procedure
The process of the sample production was performed
as follows:
• mixing the solid and liquid phases (the solid-to-
liquid-phase ratio was 2.5 g cm–3) until a fine, thick
pulp was obtained,
• mass transfer to a rectangular mould with a cover,
• setting the mould to the shaking mode for 10 min to
displace the residual air,
• keeping the samples at the room temperature for 48 h
and
• keeping the samples in a dryer at the temperature of
100 °C for 72 h,
• aging the samples for 14 d.
3 RESULTS AND DISCUSSION
The chemical content of red mud is shown in Table
1, while the chemical content of metakaolin is shown in
Table 2. The proportion of SiO2 in kaolin was 59.87 %,
the proportion of Fe2O3 was 3.12 %, the proportion of
Al2O3 was 19.45 % and the rest was water.
Table 1: Chemical content of the red mud from the Podgorica Alu-
minum Factory in mass fractions
Tabela 1: Kemijska sestava rde~ega blata iz Tovarne aluminija Pod-
gorica v masnih dele`ih
oxide w/%
Fe2O3 40.78
Al2O3 17.91
SiO2 11.28
TiO2 10.20
Na2O 6.9
Table 2: Chemical content of metakaolin in mass fractions
Tabela 2: Kemijska sestava metakaolina v masnih dele`ih
oxide w/%
SiO2 52.26
Al2O3 42.83
Fe2O3 1.01
CaO 0.02
MgO 0.09
Na2O 0.02
K2O 1.56
TiO2 0.13
ZnO <0.01
The XRD analysis of red mud (Figure 1) shows the
presence of hematite Fe2O3, gibbsite Al(OH)3, akdalaite
4Al2O3 · H2O, lapidocrocte FeO(OH) and calcite CaCO3.
The value of the specific mass of red mud was CM =
2.7773 g cm–3, and for metakaolin this value was MK =
2.4738 g cm–3.
The investigation of the chemical composition, mine-
ralogical content and thermal characteristics was per-
formed on several geopolymeric samples obtained under
different synthesis conditions (Table 3).
According to the X-ray diffraction analysis, the
geopolymer samples (specimens under numbers 1 and 8
from Table 3) showed that the dominating minerals are
M. VUK^EVI] et al.: UTILIZATION OF GEOPOLYMERIZATION FOR OBTAINING CONSTRUCTION MATERIALS ...
100 Materiali in tehnologije / Materials and technology 47 (2013) 1, 99–104
quartz (Q) and kaolinite (K) as shown in Figures 2 and
3. The feldspar appears in trace amounts. The X-ray
diagrams indicate that the treatment is characterized by
dissolution of the starting material and a formation of
amorphous and crystalline aluminosilicate phases as well
as the stable phases of leucite and kalsilite. The existence
of non-dissolved solid particles of red mud is also
indicated. The unidentified peaks in XRD diagrams
represent the residual unreacted kaolinite or sodium-
aluminosilicate phase. The selected diagrams show the
existence of an amorphous phase in the system (baseline
noise) of the aluminosilicate material. It is also clear that
M. VUK^EVI] et al.: UTILIZATION OF GEOPOLYMERIZATION FOR OBTAINING CONSTRUCTION MATERIALS ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 99–104 101
Figure 4: Infrared spectroscopy of geopolymer samples
Slika 4: Infrarde~a spektroskopija vzorca geopolimera
Figure 2: X-ray diffractogram of sample 1
Slika 2: Rentgenski difraktogram vzorca 1
Table 3: Geopolymer samples under different synthesis conditions
Tabela 3: Vzorci geopolimerov iz razli~nih razmer pri sintezi
Number of
sample
CNaOH/
(mol dm–3)
CSi/
(mol dm–3)
Content of
metakaolin
(w/%)
1 3 1 2
2 7 1 2
3 10 1 2
4 3 1.5 2
5 7 1.5 2
6 10 1.5 2
7 3 3.5 2
8 7 3.5 2
9 10 3.5 2
10 3 1 8
11 7 1 8
12 10 1 8
13 3 1.5 8
14 7 1.5 8
15 10 1.5 8
16 3 3.5 8
17 3 3.5 8
18 10 3.5 8
19 3 1 15
20 7 1 15
21 10 1 15
22 3 1.5 15
23 7 1.5 15
24 10 1.5 15
25 3 3.5 15
26 7 3.5 15
27 10 3.5 15
Figure 1: X-ray diffractogram of red mud
Slika 1: Rentgenski difraktogram rde~ega blata
Figure 3: X-ray diffractogram of sample 8
Slika 3: Rentgenski difraktogram vzorca 8
increasing concentrations of NaOH as well as an
increasing participation of the binder in the solid phase
lead to a formation of a more pronounced peak of the
new phase, i.e., sodium aluminosilicate.
The FTIR spectroscopy was used to determine the
changes in the structure during the treatment of the
starting material using the concentrated solution of
NaOH and Na-silica (Figure 4).
The characteristic wavenumbers for kaolinite are:
OH- at (3700, 3650, 3620) cm–1;
Al–OH at 913 cm–1;
Si–O at (1032, 1008, 469) cm–1;
Si–O–Al at 538 cm–1.
The absence of the Al-O-H bands at 913 cm–1 as well
as the band duplication at 3700 cm–1 and 3620 cm–1 are
evident. The absence of the bands at 539 cm–1 and 913
cm–1 and the presence of the new bands at 800 cm–1 can
be explained with the transformation of the octahedral
structure of Al3+ into a tetrahedral one under the influence
of the agents. The bands at 1100 cm–1 and 1200 cm–1 are
related to the amorphous SiO2. The characteristic peaks
of metakaolin were reduced under the influence of the
agents, but they did not disappear after introducing
1 mol dm–3 NaOH solution. When using 7 mol dm–3
NaOH solution, these peaks almost disappear and create
a new band within the wavenumber range of 1200–850
cm–1.
The TG-analysis was performed within the tempera-
ture range of 20–1200 °C.
The TGA results (Figures 5 and 6 refer to the speci-
mens under the numbers 1 and 8 in Table 3, respec-
tively) show that the mass loss occurs in two steps. In the
first step, at a temperature below 150 °C, the absorbed
water is released into the pores and on the surface. In the
temperature range of 150–600 °C, the weight loss is
associated with the pre-dehydration process, where there
is a reorganization of the octahedral lattice. In the second
step, the dehydroxylation of the starting material occurs
within the temperature range of 350–800 °C.
The density values of the obtained samples are shown
in Table 4. The density values for the geopolymer
samples are in the range between 2.2443 g cm–3 and
2.2816 g cm–3.The highest density values were obtained
with the lowest percentage of metakaolin as a binder in
the raw mixture. The explanation for this can be found in
the fact that the participation of red mud as a very dense
component in the raw mixture is the highest in the
samples with the lowest percentage of the added binder.
The results also show that an important parameter is the
M. VUK^EVI] et al.: UTILIZATION OF GEOPOLYMERIZATION FOR OBTAINING CONSTRUCTION MATERIALS ...
102 Materiali in tehnologije / Materials and technology 47 (2013) 1, 99–104
Figure 6: TG analysis of geopolymer, sample 8
Slika 6: TG-analiza geopolimera, vzorec 8
Table 4: Values of the sample density
Tabela 4: Gostota vzorcev
Number of
sample
CNaOH/
(mol dm–3)
CSi/
(mol dm–3)
Content of
metakaolin
(w/%)
Density
(g cm –3)
1 3 1 2 2.2743
2 7 1 2 2.2760
3 10 1 2 2.2750
4 3 1.5 2 2.2791
5 7 1.5 2 2.2808
6 10 1.5 2 2.2820
7 3 3.5 2 2.2747
8 7 3.5 2 2.2806
9 10 3.5 2 2.2816
10 3 1 8 2.2673
11 7 1 8 2.2705
12 10 1 8 2.2710
13 3 1.5 8 2.2650
14 7 1.5 8 2.2670
15 10 1.5 8 2.2690
16 3 3.5 8 2.2655
17 3 3.5 8 2.2659
18 10 3.5 8 2.2690
19 3 1 15 2.2520
20 7 1 15 2.2530
21 10 1 15 2.2550
22 3 1.5 15 2.2443
23 7 1.5 15 2.2480
24 10 1.5 15 2.2550
25 3 3.5 15 2.2510
26 7 3.5 15 2.2520
27 10 3.5 15 2.2540
Figure 5: TG analysis of geopolymer, sample 1
Slika 5: TG-analiza geopolimera, vzorec 1
ratio of the components in an alkaline activator, so the
combination of a high hydroxide concentration and a low
silicate concentration generates lower density values.
The explanation can be found in the fact that a high
NaOH concentration influences the deficit of the silicon
needed for the reaction of geopolymerization.2,4
The values of compressive strength of geopolymer in
dependence of the concentration of the activator (NaOH
-Na-silicate) and the presence of a binder are shown in
Figures 7a and b, respectively.
The factors affecting an increase in compressive
strength are different concentrations of alkaline activa-
tors. The results show that with the increasing concentra-
tions of NaOH (3, 7, 10) mol dm–3 the compressive strength
increases up to a NaOH concentration of 7 mol dm–3.
Higher NaOH concentrations (10 mol dm–3) cause a
reduction in the compressive-strength value. The expla-
nation for this lies in the fact that the initial increase in
the NaOH concentration leads to an increase in the
silicon and aluminum dissolution from the solid phase2.
The increased Si and Al contents in the aqueous phase
are essential for initiating the oligomer formation and
polycondensation. The decrease in the values that occurs
under higher NaOH concentrations is the consequence of
the fact that dissolved silicon and aluminum remain
almost constant while the free NaOH increases, resulting
in a lower SiO2/Na2O ratio in the aqueous phase. There-
fore, the monosilicates and oligomeric species are
predominantly in favor of polymer and consequently the
polycondensation is slower2.
Change in the silicon concentration in Na-silicate (1,
1.5, 3.5) mol dm–3 causes an increase in the compressive
strength of the geopolymer samples. With the increasing
concentration of the alkaline activator, the amount of
dissolved silicon in the reaction mixture increases as
well. Silica originating from sodium silicate has an
M. VUK^EVI] et al.: UTILIZATION OF GEOPOLYMERIZATION FOR OBTAINING CONSTRUCTION MATERIALS ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 99–104 103
Figure 8: SEM microphotographs of inorganic polymer materials: a)
sample (CNaOH = 7 mol dm–3; CSi = 1.5 mol dm–3; S/L = 2.5 g cm–3),
magnification 500-times; b) sample (CNaOH = 7 mol dm–3; CSi =
1.5 mol dm–3; S/L = 2.5 g cm–3), magnification 5000-times
Slika 8: SEM-posnetek anorganskega polimera: a) vzorec (CNaOH =
7 mol dm–3; CSi = 1,5 mol dm–3; S/L = 2,5 g cm–3), pove~ava
500-kratna; b) vzorec (CNaOH = 7 mol dm–3; CSi = 1,5 mol dm–3; S/L =
2,5 g cm–3), pove~ava 5000-kratna
Figure 7: Values of compressive strength of geopolymer samples: a)
influence of NaOH concentration on compressive strength as a func-
tion of binder percentage; b) influence of Si concentration on com-
pressive strength as a function of binder percentage
Slika 7: Vrednosti tla~ne trdnosti vzorcev geopolimera: a) vpliv
koncentracije NaOH na tla~no trdnost v odvisnosti od dele`a veziva;
b) vpliv koncentracije Si na tla~no trdnost v odvisnosti od dele`a
veziva
important role because it starts the reaction of geopoly-
merization by allowing a faster and more complete
dissolution from the raw material2. A higher silicon
concentration leads to the formation of silicate species
with a complex polymeric structure, thus allowing the
three-dimensional polymeric framework to rise. Soluble
silica fosters the polycondensation. Under the higher
initial silica concentrations, the surface cracks were
noticed. This might be because of the entrapped free
water of the aqueous phase (the water for the Na-silicate
dissolution).
The level of substitution of red mud with metakaolin
w = (4, 8 and 15) % causes an increase in the com-
pressive-strength value up to w = 15 %. The lower level
of the compressive strength with 15 % of metakaolin, or
higher, can be explained with a lack of NaOH for
dissolving such a quantity of metakaolin, or with the fact
that a high level of polycondensation (because of an
excessive amount of metakaolin) can create a surface
non-permeable membrane entrapping water from the
liquid phase (water for the dissolution of Na-silicate)4.
The microstructure of the synthesized inorganic
polymer material was investigated by scanning electron
microscopy and it is shown in Figure 8.
SEM microphotographs show that the obtained
materials are compact, with no discontinuity, which is
confirmed by the mechanical properties. The isolated
pores noticed inside the material are in the range of up to
200 μm. The presence of a new amorphous phase can be
seen in Figure 8a. In Figure 8b a gelatinous phase
around the particles of the starting material is identified
under high magnification.
4 CONCLUSION
This investigation shows that the red mud obtained as
a by-product of the Bayer process for obtaining alu-
minum in the Podgorica Aluminum Factory is, according
to its physicochemical and mechanical properties, a
good-quality aluminum-silicate material appropriate for
geopolymer formation. Under the optimum synthesis
conditions (the S/L ratio of 2.5 g cm–3, the NaOH
concentration of 7 mol dm–3, the metakaolin percentage
of 10 %), the red-mud/metakaolin-based polymeric
materials develop satisfactory compressive strength.
These mechanical properties were attributed to the
formation of the amorphous phase that bonded the
non-dissolved particles of the raw solid materials in a
good manner. The presence of this phase was also
revealed by the XRD, TG and FTIR analyses as well as
SEM analysis.
The results of the investigations of the basic
properties of the developed geopolymer show that it can
be used in civil engineering (as a substitute for brick
products and road foundations). Depending on the
purpose of a product, an additional research is required
(resistance to different external effects, durability in
exploitation conditions, etc.) as well as an analysis of the
economic feasibility.
Apart from that, the described process of geopoly-
merization can contribute significantly to environmental
preservation, since there is a possibility of conserving
large quantities of industrial waste.
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2 I. Giannopoulou, D. Dimas, I. Maragkos, D.Panias, Utilization of
Metallurgical solid by-products for the development of inorganic
polymeric construction materials, Global NEST Journal, 11 (2009),
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for synthesys of inorganic polymeric materials, Mineral processing
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5 M. Izquierdo, X. Querol, C. Phillipart, D. Antenucci, World of coal
ash (WOAC) Conference, Lexington, KY, USA, 2008
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GC 4, Engineering Faculty, Curtin University of Technology, Perth,
Australia, 2008
7 P. De Silva, K. Sagoe-Crenstil, V. Sirivivatnanon, Kinetics of geo-
polymerisation: Role of Al2O3 and SiO2, Cement and Concrete
Research, 37 (2007), 512–518
8 K. Sagoe-Crenstil, L. Wang, Dissolution processes, hydrolysis and
condensation reaction during geopolymer synthesis: Part 2. High
Si/Al ratio systems, Journal of Materials Science, 42 (2007),
3007–3014
9 J. S. J. van Deventer, J. L. Provis, P. Duxon, G. C. Lukey, Reaction
mechanisms in the geopolymeric conversion of inorganic waste to
usefull products, Journal of hazardous materials, A139 (2007),
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M. VUK^EVI] et al.: UTILIZATION OF GEOPOLYMERIZATION FOR OBTAINING CONSTRUCTION MATERIALS ...
104 Materiali in tehnologije / Materials and technology 47 (2013) 1, 99–104
M. K. KÜLEKCÝ: ANALYSIS OF PROCESS PARAMETERS FOR A SURFACE-GRINDING PROCESS ...
ANALYSIS OF PROCESS PARAMETERS FOR A
SURFACE-GRINDING PROCESS BASED ON THE
TAGUCHI METHOD
ANALIZA PROCESNIH PARAMETROV POSTOPKA BRU[ENJA
POVR[INE S TAGUCHIJEVO METODO
Mustafa Kemal Külekcý
Mersin University, Tarsus Technical Education Faculty, Department of Mechanical Education, 33140, Tarsus-Mersin, Turkey
mkkulekci@gmail.com
Prejem rokopisa – received: 2012-06-18; sprejem za objavo – accepted for publication: 2012-08-27
In this study, the Taguchi method that is a powerful tool to design optimization for quality is used to find the optimum surface
roughness in grinding operations. An orthogonal array, a signal-to-noise (S/N) ratio, and an analysis of variance (ANOVA) are
employed to investigate the surface-roughness characteristics of AISI 1040 steel plates using EKR46K grinding wheels.
Through this study, not only can the optimum surface roughness for grinding operations be obtained, but the main grinding
parameters affecting the performance of grinding operations can also be found. Experimental results are provided to confirm the
effectiveness of this approach. The results of this study showed that the depth of cut and the wheel speed have significant effects
on the surface roughness, while the rate of feed has a lower effect on it.
Keywords: grinding, surface integrity, surface topography, surface roughness
V tej {tudiji je bila uporabljena Taguchijeva metoda kot mo~no orodje za dolo~anje pogojev za doseganje optimalne hrapavosti
pri bru{enju. Za preiskavo zna~ilnosti hrapavosti povr{ine plo{~ iz jekla AISI 1040 pri uporabi brusnih kolutov EKR46K so bile
uporabljene ortogonalna matrika, signal hrupa (S/N) in analiza variance (ANOVA). S to {tudijo je bilo mogo~e dobiti optimalno
hrapavost povr{ine ter glavne parametre bru{enja, ki vplivajo na zmogljivost bru{enja. Na voljo so eksperimentalni rezultati, ki
potrjujejo ta na~in. Rezultati {tudije ka`ejo, da imata globina bru{enja in hitrost vrtenja brusnega koluta pomemben u~inek na
hrapavost povr{ine. Hitrost podajanja pa ima manj{i vpliv na hrapavost povr{ine.
Klju~ne besede: bru{enje, povr{inska integriteta, topografija povr{ine, hrapavost povr{ine
1 INTRODUCTION
Grinding is a manufacturing process with an un-
steady process behavior, whose complex characteristics
determine the technological output and the quality. An
assessment of the grinding-process quality usually inclu-
des the micro-geometric quantities of the component. In
order to predict the component behavior during the use
or to control the grinding process, it is necessary to
quantify surface roughness, which is one of the most
critical quality constraints for the selection of grinding
factors in a process planning. The process set-up often
depends on the operator competence1.
The quality of a surface generated by grinding deter-
mines many workpiece characteristics such as the
minimum tolerances, the lubrication effectiveness and
the component life, among others. A typical surface is
characterized by clean cutting paths and plowed material
to the sideway of some grooves. However, many other
marks can be found, such as cracks produced by the
thermal impact, back-transferred material and craters
produced by a grain fracture2,3.
The surface quality produced in surface grinding is
influenced by various parameters such as4,5: i. wheel
parameters – abrasives, grain size, grade, structure,
binder, shape and dimension; ii. workpiece parameters –
fracture mode, mechanical properties and chemical
composition; iii. process parameters – wheel speed,
depth of cut, table speed and dressing condition; iv.
machine parameters – static and dynamic characteristics,
spindle system, and table system.
Kwak4 evaluated the effect of grinding parameters on
the geometric error and optimum grinding conditions.
Krajinik et al.1 developed a second-order surface-rough-
ness model using the central composite design technique.
Hecker et al.2 presented a prediction of the arithmetic
mean surface roughness based on a probabilistic,
undeformed, chip-thickness model for surface roughness
in a grinding process. Gupta et al.6 optimized grinding-
process parameters using a numerical method. Tawakoli
et al.7 investigated the effects of a workpiece and
grinding parameters on minimum quantity lubrication
(MQL) and they compared the results with dry lubri-
cation. Silva et al.7–9 investigated the effects of grinding
parameters on the ABNT 4340 steel using the MQL
technique. Shaji et al.5 analyzed the effects of process
parameters and the mode of dressing on the force com-
ponents and surface using the Taguchi experimental
design method. Taguchi is the developer of the Taguchi
method. He proposed that an engineering optimization of
a process or product should be carried out with a three-
step approach: i. system design, ii. parameter design, and
iii. tolerance design10,11.
Materiali in tehnologije / Materials and technology 47 (2013) 1, 105–109 105
UDK 621.923:620.179.11:519.233.4 ISSN 1580-2949
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In this study the wheel speed (V), the rate of feed (F)
and the depth of cut (D) were selected as variable
parameters. Other process parameters were fixed. The
above parameters of grinding were analyzed and opti-
mized with the Taguchi method using the experimentally
obtained surface-roughness (Ra) results. Confirmation
experiments were conducted at the optimum level to
verify the effectiveness of the Taguchi approach.
2 EXPERIMENTAL STUDIES
Commercial AISI 1040 steel plates with the dimen-
sions of 10 mm × 60 mm × 100 mm were used as the
workpiece material. The grinding experiments were
carried out on a grinding machine as shown in Figure 1
using EKR46K grinding wheels.
The intense heat generated during the grinding due to
a relatively high friction impairs the workpiece quality
by inducing thermal damage. Therefore, the cooling and
lubrication play a decisive role in grinding. In this study,
a cutting fluid was used as a coolant in order to cool the
workpiece and, hence, to prevent the thermal damage of
the workpiece.
The initial grinding parameters were selected as
follows: the wheel speed (V) of 1000 r/min; the feed rate
(F) of 25 m/min and the depth of cut (D) of 0.10 mm. A
feasible space for the grinding parameters was defined
on the basis of the reference12. These parameters were
further defined as follows: the wheel speed in the range
of 1000–2 000 r/min, the feed rate in the range of 20–30
m/min, and the depth of cut in the range of 0.05–0.15
mm. Also, during the grinding operation the workpieces
were cooled using cutting fluids. In the grinding
parameter design, three levels of grinding parameters
were selected as shown in Table 1. The surface rough-
ness of grinded workpices was recorded as the Ra value
with the ISO class-3-type, surface-roughness tester.
3 RESULTS AND DISCUSSION
To select an appropriate orthogonal array for the
experiments, the total degrees of freedom need to be
computed. The degrees of freedom are defined as the
number of comparisons between the design parameters
that need to be made to determine which level is better
and, specifically, how much better it is. For example, a
four-level design parameter counts for three degrees of
freedom11. The degrees of freedom (Vf) associated with
the interaction between two design parameters are given
by the product of the degrees of freedom for the two
design parameters. In this study, the interaction between
the grinding parameters is neglected. Therefore, there are
eight degrees of freedom (Vf = 9 –1= 8) owing to there
being three grinding parameters (such as the wheel
speed, the rate of feed and the depth of cut) of the grind-
ing operations. Once the required degrees of freedom are
known, the next step is to select an appropriate ortho-
gonal array to fit the specific task. Basically, the degrees
of freedom for the orthogonal array should be greater
than, or at least equal to, those for the design parameters.
In this study, an L9 orthogonal array with four columns
and nine rows was used. This array has eight degrees of
freedom and it can handle three-level design parameters.
Each grinding parameter is assigned to a column, having
nine grinding-parameter combinations available. There-
fore, only nine experiments are required to study the
entire parameter space using the L9 orthogonal array. The
experimental layout for the three grinding parameters
using the L9 orthogonal array is shown in Table 2. Since
the L9 orthogonal array has four columns, one column of
the array is left empty for the error of the experiments;
orthogonality is not lost by keeping one column of the
array empty11. To increase the sensitivity of the results
M. K. KÜLEKCÝ: ANALYSIS OF PROCESS PARAMETERS FOR A SURFACE-GRINDING PROCESS ...
106 Materiali in tehnologije / Materials and technology 47 (2013) 1, 105–109
Figure 1: Experimental grinding machine
Slika 1: Eksperimentalni brusilni stroj
Table 1: Grinding parameters and their levels
Tabela 1: Parametri pri bru{enju in njihove ravni
Symbol Grindingparameter Unit Level 1 Level 2 Level 3
V Wheel speed r/min 1000 1500a 2000
F Rate of feed m/min 20 25a 30
D Depth of cut mm 0.05 0.10a 0.15
a: initial grinding parameters
Table 2: Experimental layout of the L9 orthogonal array
Tabela 2: Eksperimentalna postavitev ortogonalne matrike L9
Experiment
number
Grinding parameter level
V/(r/min)
Wheel speed
F/(m/min)
Rate of feed
D/mm
Depth of cut
1 1 1 1
2 1 2 2
3 1 3 3
4 2 1 2
5 2 2 3
6 2 3 1
7 3 1 3
8 3 2 1
9 3 3 2
each experiment in the L9 orthogonal array was repeated
three times.
Experimental results were transformed into a
signal-to-noise (S/N) ratio. There are three categories of
quality characteristics, i.e., the-lower-the-better, the
higher-the-better, and the-nominal-the-better. In the
present study the-lower-the-better quality characteristic
was selected for the surface roughness. The loss function
of the-lower-the-better quality characteristics can be
expressed as13:
L
n
yj i
k
n
=
=
∑1 2
1
(1)
! j jL= −10 lg (2)
where n is the number of tests, yi is the experimental
value of the ith quality characteristic, Lj is the overall
loss function and !j is the S/N ratio. Table 3 shows the
average experimental results for the surface roughness
and the corresponding S/N ratio using equations (1) and
(2).
Table 3: Experimental results for the surface roughness and S/N ratio
Tabela 3: Eksperimentalno izmerjena hrapavost povr{ine in razmerje
S/N
Experi-
ment
number
Process parameters
Ra
(μm)
S/N ratio
(dB)V
(r/min)
F
(m/min)
D
(mm)
1 1000 20 0.05 0.22 13.15
2 1000 25 0.10 0.32 9.90
3 1000 30 0.15 0.56 5.04
4 1500 20 0.10 0.28 11.06
5 1500 25 0.15 0.30 10.46
6 1500 30 0.05 0.20 13.98
7 2000 20 0.15 0.48 6.38
8 2000 25 0.05 0.38 8.40
9 2000 30 0.10 0.40 7.96
Since the experimental design is orthogonal, it is then
possible to separate out the effect of each grinding
parameter at different levels11,13. The mean S/N ratio for
each level of the grinding parameters is summarized in
an S/N response table for the surface roughness. In addi-
tion, the total mean S/N ratio for the nine experiments is
also calculated and listed in Table 4.
As shown in equations (1) and (2), from the view
point of the-lower-the-better quality characteristic, a
greater S/N ratio corresponds to a smaller variance of the
output characteristic around the desired value. As
indicated in Figure 2, the optimum parameters for the
grinding of the AISI 1040 steel become the wheel speed
at level 2, the feed rate at level 1 and the depth of cut at
level 1 and therefore the combination of parameters is
V2F1D1.
The purpose of the analysis of variance (ANOVA) is
to investigate which design parameters significantly
affect the quality characteristic. This is accomplished by
separating the total variability of the S/N ratios, which is
measured with the sum of the squared deviations from
the total mean S/N ratio, into contributions by each
design parameter and the error. First, the total sum of
squared deviations SSTotal from the total mean S/N ratio
!m can be calculated as11,13:
SS Total i m
i
n
= −
=
∑ ( )! ! 2
1
(3)
where !i is the mean S/N ratio for the ith experiment.
Statistically, there is a tool called an F test named after
Fisher14 enabling us to see which design parameters
have a significant effect on the quality characteristic.
Normally, the change of the grinding parameter has a
significant effect on the surface roughness when the F
value is large. In performing the F test, the mean of
squared deviations for each design parameter needs to
be calculated. The mean of squared deviations is equal
to the sum of squared deviations divided by the number
of degrees of freedom associated with the design
parameter. Then the F value for each design parameter
is simply the ratio of the mean of squared deviations to
M. K. KÜLEKCÝ: ANALYSIS OF PROCESS PARAMETERS FOR A SURFACE-GRINDING PROCESS ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 105–109 107
Table 4: S/N response table for grinding parameters
Tabela 4: Tabela odzivov S/N za razli~ne parametre bru{enja
Symbol Grindingparameters
Mean S/N ratio (dB) Total mean S/N
(dB)
Maximum-
MinimumLevel 1 Level 2 Level 3
V/(r/min) Wheel speed 9.36 11.83 7.58 9.60 4.25
F/(m/min) Rate of feed 10.19 9.59 8.99 1.20
D/mm Depth of cut 11.85 9.64 7.29 4.56
Figure 2: S/N graph for the surface roughness
Slika 2: S/N-graf za hrapavost povr{ine
the mean of the squared error. Table 5 shows the results
of ANOVA for the grinding parameters.
It can be found that the depth of cut and the wheel
speed have significant effects on the surface roughness.
Increasing wheel speed and depth of cut increase the
surface roughness. However, the change in the rate of
feed has a lower effect on the surface roughness. In the
present study, the contribution order of the studied
grinding parameters for the surface roughness of the
AISI 1040 steel is found to be as follows: the depth of
cut (50 %), the wheel speed (40 %) and the rate of feed
(9 %).
Once the optimum level of design parameters has
been selected, the final step is to predict and verify the
improvement of the quality characteristic using the
optimum level of design parameters. The estimated S/N
ratio

! using the optimum level of design parameters can
be calculated as:

! ! ! != + −−
=
∑m i m
i
j
( ) 2
1
(4)
where !m is the total mean S/N ratio, ! i
− is the mean
S/N ratio at the optimum level, and j is the number of
the main design parameters that affect the quality cha-
racteristic. The estimated S/N ratio using the optimum
grinding parameters for surface roughness can then be
obtained and the corresponding surface roughness can
also be calculated by using equations (1) and (2). Table
6 shows a comparison of the predicted surface rough-
ness with the actual surface roughness using the opti-
mum grinding parameters. A close agreement between
the predicted and actual surface roughness is observed.
The increase in the S/N ratios from the initial surface
roughness to the optimum surface roughness is calcul-
ated as 0.68 dB.
The values given in Table 6 were calculated with the
Taghuchi method and experimentally verified.
4 CONCLUSIONS
In this study, an application and adaptation of the
Taguchi optimization and quality-control method were
established for the optimization of the surface roughness
in a grinding process. The Taguchi method provides a
systematic and efficient methodology with fewer
experiments and trials. The experimental results obtained
in this study showed that the depth of cut and the wheel
speed have significant effects on the surface roughness.
The rate of feed has a lower effect on the surface
roughness. The contribution order of the grinding para-
meters including the depth of cut, the wheel speed and
the rate of feed is 50 %, 40 % and 10 %, respectively. A
change made to all the grinding parameters significantly
changes the surface roughness. The optimum grinding-
parameter combination for the AISI 1040 steel includes
the wheel speed (V) of 1 500 r/min and the depth of cut
(D) of 0.05mm.
5 REFERENCES
1 P. Krajnik, J. Kopac, A. Sluga, Design of grinding factors based on
response surface methodology, Journal of Materials Processing
Technology, 162 (2005), 629–636
2 L. R. Hecker, Y. L. Steven, Predictive modeling of surface roughness
in grinding, International Journal of Machine Tools & Manufacture,
43 (2003), 755–761
3 C. P. Bhateja, Grinding: theory, techniques and troubleshooting,
Society of Manufacturing Engineers Society of Manufacturing
Engineers (SME), Dearborn, 1982
4 J. Kwak, Application of Taguchi and response surface methodologies
for geometric error in surface grinding process, International Journal
of Machine Tools & Manufacture, 45 (2005), 327–334
5 S. Shaji, V. Radhakrishnan, Analysis of process parameters in
surface grinding with graphite as lubricant based on the Taguchi
method, Journal of Materials Processing Technology 141 (2003) 1,
51–59
6 R. Gupta, K. S. Shishodia, G. S. Sekhon, Optimization of grinding
process parameters using enumeration method, Journal of Materials
Processing Technology, 112 (2001), 63–67
M. K. KÜLEKCÝ: ANALYSIS OF PROCESS PARAMETERS FOR A SURFACE-GRINDING PROCESS ...
108 Materiali in tehnologije / Materials and technology 47 (2013) 1, 105–109
Table 5: Results of ANOVA for the grinding parameters
Tabel 5: Rezultati za parametre bru{enja ANOVA
Symbol Grindingparameters
Degree of
freedom Sum of square Mean square F Contribution (%)
V Wheel speed 2 0.040 0.020 2.48 40
F Rate of feed 2 0.009 0.003 0.40 9
D Depth of cut 2 0.050 0.025 3.09 50
Error 2 0.001 0.003 1
Total 8 0.1 100
Table 6: Results of the confirmation experiment
Tabela 6: Rezultati potrditvenih preizkusov
Initial grinding parameters
Optimum grinding parameters Improvement of the S/N
ratio (dB)Prediction Experiment
Level V1F1D1 V2F1D1 V2F1D1 0.68
Ra/μm 0.22 0.19 0.20
S/N (dB) 13.15 14.69 13.83
7 T. Tawakoli, M. J. Hadad, M. H. Sadeghi, A. Daneshi, S. Stockert, A.
Rasifard, An experimental investigation of the effects of workpiece
and grinding parameters on minimum quantity lubrication-MQL
grinding, International Journal of Machine Tools & Manufacture, 49
(2009), 924–932
8 L. R. Silva, E. C. Bianchi, R. E. Catai, R. Y. Fusse, T. V. França, P.
R. Aguiar, Study on the behavior of the minimum quantity lubri-
cant-MQL technique under different lubrication and cooling
conditions when grinding ABNT 4340 steel, J. Braz. Soc. Mech. Sci.
Eng., 27 (2005) 2, 192–199
9 L. R. Silva, E. C. Bianchi, R. Y. Fusse, R. E. Catai, T. V. França, P.
R. Aguiar, Analysis of surface integrity for minimum quantity
lubricant-MQL in grinding, Int. J. Mach. Tools Manuf., 47 (2007) 2,
412–418
10 G. Taguchi, Introduction to Quality Engineering, Asian Productivity
Organization, Tokyo 1990
11 W. H. Yang, Y. S. Tarng, Design Optimization of cutting parameters
for turning operations based on the Taguchi method, Journal of
Materials Processing Technology, 84 (1998), 122–129
12 M. P. Groover, Fundamentals of Modern Manufacturing, John Wiley
& Sons, Inc., Danvers 2010
13 D. C. Montgomery, Design and Analysis of Experiments, Wiley,
Singapore 1991
14 R. A. Fisher, Statistical Methods for Research Workers, Oliver and
Boyd, London 1925
M. K. KÜLEKCÝ: ANALYSIS OF PROCESS PARAMETERS FOR A SURFACE-GRINDING PROCESS ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 105–109 109

M. IPEK: CHARACTERIZATION OF AN Cu-Cr-Zr ALLOY SYNTHESIZED WITH THE POWDER-METALLURGY TECHNIQUE
CHARACTERIZATION OF AN Cu-Cr-Zr ALLOY
SYNTHESIZED WITH THE POWDER-METALLURGY
TECHNIQUE
KARAKTERIZACIJA ZLITINE Cu-Cr-Zr, SINTETIZIRANE S
TEHNIKO PRA[NE METALURGIJE
Mediha Ipek
Sakarya University, Engineering Faculty, Department of Metallurgy and Materials Engineering, Esentepe Campus, 54187 Sakarya-Turkey
mipek@sakarya.edu.tr
Prejem rokopisa – received: 2012-08-16; sprejem za objavo – accepted for publication: 2012-09-14
A Cu-1.5 % Cr-0.5 % Zr alloy was prepared with the powder-metallurgy (PM) method. Cu-Cr-Zr powders were pressed under
390 MPa of uniaxial compression and sintered at 1000 °C for 2 h. After the holding time, the samples were immediately taken
out of the furnace and pressed at 850 MPa. The sintered samples were solution-treated at 1000 °C for 15 min and water
quenched. Then they were deformed by 20 % at room temperature and aged at 470 °C for (2, 4, 6 and 8) h. A SEM investigation
revealed that the Cr and Zr particles having a limited solubility in Cu distributed homogeneously in the copper matrix. An XRD
analysis showed that each sample (sintered and aged at 470 °C for different times) has the same phases: copper and trace Cr2O3.
The relative density of the sinter-pressed sample increased from 92 % to 96 % due to cold deformation. The microhardness of
the samples ranged from 77 HV for the solution-treated and water-quenched sample to 116 HV for the aged sample, and the
electrical conductivity of the samples increased from 76 % IACS for the sinter-pressed sample to 87 % IACS for the sample
aged for 8 h.
Keywords: Cu-Cr-Zr, powder metallurgy, age hardenable, electrical conductivity, hardness
Zlitina Cu- 1,5 % Cr- 0,5 % Zr je bila izdelana po metodi pra{ne metalurgije (PM). Prahovi Cu-Cr-Zr so bili enoosno stisnjeni s
390 MPa in 2 h sintrani na temperaturi 1000 °C. Potem so bili vzorci vzeti iz pe~i in stisnjeni z 850 MPa. Sintrani vzorci so bili
raztopno `arjeni 15 min na temperaturi 1000 °C in ohlajeni v vodi. Nato so bili deformirani za 20 % pri sobni temperaturi in
starani (2, 4, 6 in 8) h na temperaturi 470 °C. Preiskava s SEM je odkrila, da imajo delci Cr in Zr omejeno topnost v Cu in da so
enakomerno razporejeni v osnovi iz bakra. Rentgenska analiza (XRD) je pokazala, da ima vsak vzorec (sintran in razli~no dolgo
staran na 470 °C) enake faze: baker in sledove Cr2O3. Relativna gostota sintranega in stisnjenega vzorca je s hladno deformacijo
narasla iz 92 % na 96 %. Mikrotrdota vzorcev je bila v obmo~ju od 77 HV za raztopno `arjen in v vodi ohlajen vzorec ter do
116 HV za staran vzorec. Elektri~na prevodnost sintranega in stisnjenega vzorca je bila 76 % IACS in je narasla na 87 % IACS
pri vzorcu, staranem 8 h.
Klju~ne besede: Cu-Cr-Zr, pra{na metalurgija, utrjevanje s staranjem, elektri~na prevodnost, trdota
1 INTRODUCTION
Age hardenable Cu-Cr-Zr alloys are widely used for
many applications such as trolley wires, electrodes for
resistance welding and lead-frame materials due to their
high strength, high electrical and thermal conducti-
vities1–3. Thermal aging is used to obtain high strength
and, simultaneously, high electrical conductivity of the
Cu–Cr–Zr alloys. The high electrical conductivity of an
alloy is due to a very low solubility of Cr and Zr in the
Cu matrix, whereas the excellent strength is attributed to
the combined effect of precipitation hardening, work
hardening, texture strengthening and alloy strengthen-
ing1,3–6. In order to control the microstructure and
improve the properties of a Cu–Cr–Zr alloy, it is of great
importance to optimize the aging process and identify
the composition of the precipitates. There has been no
unanimous agreement on the precipitation phase of the
alloy1,7. In general, the manufacturing processing of
Cu-Cr-Zr alloys includes casting, hot rolling, solution
treatment, water quenching, cold deformation and aging
steps1,2. In the present study, we characterized the
age-hardenable Cu-Cr-Zr alloys produced in the open
atmosphere with the powder-metallurgy technique. There
is little information about this method in the available
literature.
2 EXPERIMENTAL DETAILS
In this study, an Cu-1.5 % Cr-0.5 % Zr alloy was
prepared with the powder-metallurgy (PM) method. For
this purpose, the Cu-Cr-Zr powders obtained from Alpha
Aesar with the average grain sizes of (10, <10, 2–3) μm,
respectively, were mixed mechanically for 2 h, pressed
under 390 MPa of uniaxial compression and sintered at
1000 °C for 2 h in an open-atmospheric, electric-resist-
ance furnace. After the holding time, the samples were
immediately taken out of the furnace and pressed at 850
MPa. The sinter-pressed samples were solution-treated at
1000 °C for 15 min and water-quenched. Then they were
mechanically deformed by 20 % at room temperature
and aged at 470 °C for (2, 4, 6 and 8) h. A phase analysis
of the sinter-pressed, solution-treated, water-quenched
and aged samples was performed with the XRD-analysis
Materiali in tehnologije / Materials and technology 47 (2013) 1, 111–114 111
UDK 621.762:621.762.5:621.785.797 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(1)111(2013)
technique using the Cu K radiation with a wave length
of 15.418 nm over a 2 range of 10° "20° "90°. The
microstructures of the products were examined by means
of scanning electron microscopy, energy dispersive
spectroscopy (SEM-EDS). Relative densities of the
sintered samples were determined with the Archimedes’
method. The microhardness of the polished specimens
was measured with a load of 50 g.
3 RESULTS AND DISCUSSION
Figure 1 shows the SEM micrographs of a sinter-
pressed sample and a sintered, solution-treated, water-
quenched sample aged for 8 h. The microstructures in
Figure 1 generally include homogenously distributed
dark-grey particles in the light-grey matrix. The light-
grey areas show the copper matrix, while the dark-grey
particles mainly include chromium and small amounts of
copper and zirconium (Figure 2a, Mark 4 and Figure
2b, Marks 3, 4, 10). In addition, the lighter-grey particles
in the copper matrix are rich in zirconium which was
confirmed by EDS analyses. As in the chromium-rich
regions, the zirconium-rich regions include small
M. IPEK: CHARACTERIZATION OF AN Cu-Cr-Zr ALLOY SYNTHESIZED WITH THE POWDER-METALLURGY TECHNIQUE
112 Materiali in tehnologije / Materials and technology 47 (2013) 1, 111–114
Figure 1: SEM micrographs of the samples: a) sinter-pressed, b) sin-
tered, solution-treated, quenched and aged for 8 h
Slika 1: SEM-posnetka vzorcev: a) sintrano-stisnjeno, b) sintrano in
raztopno `arjeno, hitro ohlajeno, starano 8 h
Mark Mass fractions, w/%O Cr Cu Zr
1 2.6 1.3 95.4 0.7
2 29.6 8.6 8.3 53.5
3 27.1 35.3 37.4 0.2
4 39.5 49.2 11.0 0.3
5 16.3 4.2 20.6 58.9
6 19.5 0.7 11,0 68.8
7 0.0 0.2 99.6 0.2
8 0.6 0.2 99.0 0.2
9 10.9 9.8 78.7 0.6
10 37.4 52.4 10,0 0.2
Figure 2: SEM micrographs and EDS point analyses of the samples:
a) sinter-pressed, b) sintered, solution-treated and quenched
Slika 2: SEM-posnetka in EDS to~kaste analize vzorcev: a) sintrano-
stisnjeno, b) sintrano-raztopno `arjeno-hitro ohlajeno
Mark
Mass fractions, w/%
O Cr Cu Zr
1 10.1 0.4 56.1 33.4
2 17.0 4.4 31.0 46.6
3 18.8 0.1 11.9 69.2
4 28.5 65.2 5.8 0.5
5 13.4 0.3 45.7 40.6
6 100.0
7 100.0
amounts of copper and chromium (Figure 2a, Marks 1,
2, 5 and Figure 2b, Marks 2, 5, 6). In addition, oxygen
was detected in the chromium- and zirconium-rich
regions. SEM-EDS analyses revealed that the solution
heat treatment increases the solubility of Cr and Zr in the
copper matrix, but this solubility is still low. XRD
patterns of the sinter-pressed samples and the sintered,
solution-treated, water-quenched, 20 % cold-pressed and
aged samples are similar to each other and have
predominantly copper and trace Cr2O3 phases (Figure 3).
In the present study, EDS analyses also revealed that
chromium oxidized and this shows that the heat
treatments of this alloys must be done in a controlled
atmospheric medium. Relative density, hardness and
electrical conductivity of the samples are given in Table
1 as a function of process steps. While the relative den-
sities of the sinter-pressed and sintered, solution-treated
samples are approximately 92 %, the cold-deformed and
aged samples have relative densities of over 94 %. It is
obvious that cold deformation increases relative densi-
ties. The hardness values were determined by taking the
average of the five different measurements made on each
sample.
The hardness values for the sinter-pressed samples
and the solution-treated, water-quenched samples are
123 HV and 77 HV, respectively. It was found that the
hardness of a sample decreased after the processes of
solution treatment and water quenching. The plastic
deformation after the sintering may have caused the
deformation hardening and the hardness therefore
increased. After the aging the hardness value amounted
to 116 HV for the sample aged for 8 h. A large number
of dislocations and vacancies, introduced into the alloy
by cold deformation, provide the nucleation sites for the
formation of the Cr- and Zr-rich clusters and Guinier-
Preston zones during the aging treatment, resulting in the
acceleration of the precipitation process of the solute
atoms8. These hardness values are lower than those from
the other reports on the Cu-Cr-Zr alloys, which have
150–200 HV deformed by approximately 70 %.6,9
Electrical-conductivity values for casting and hot rolling
the Cu-Cr-Zr alloys given in the literature are 70–80 %
IACS10,11 and in the present study, an electrical con-
ductivity of 87 % IACS was obtained for the sample
aged for 8 h. A remarkable improvement in the mecha-
nical and electrical properties of the Cu–Cr–Zr alloy
during the aging treatment is due to the precipitation
which hardens the copper matrix with the high density of
the nano-scale particles and, simultaneously, depletes the
solute concentration in the copper matrix. In addition,
cold-deformation characteristics, such as dislocations
and vacancies, provide nucleation sites for the precipi-
tation8.
4 CONCLUSION
The following results can be drawn from the present
study: cold deformation enhances the relative density of
the sinter-pressed samples; cold deformation and aging
increase the hardness and electrical conductivity from 77
HV to 116 HV, and from 76 % IACS to 87 % IACS,
respectively.
5 REFERENCES
1 H. Li, S. Xie, X. Mi, P. Wu, J. Materials Science Technology, 23
(2007) 6, 795–800
2 H. Fuxiang, M. Jusheng, N. Honglong, Geng Zhiting, L. Chao, G.
Shumei, Y. Xuetao, W. Tao, L. Hong, L. Huafen, Scripta Materialia,
48 (2003), 97–102
3 M. Hatakeyama, T. Toyama, Y. Nagai, M. Hasegawa, M. Eldrup, B.
N. Singh, Materials Transactions, 49 (2008) 3, 518–521
4 C. A. Poblano-Salasa, J. D. O. Barceinas-Sanchez, J. Alloys and
Compounds, 485 (2009), 340–345
M. IPEK: CHARACTERIZATION OF AN Cu-Cr-Zr ALLOY SYNTHESIZED WITH THE POWDER-METALLURGY TECHNIQUE
Materiali in tehnologije / Materials and technology 47 (2013) 1, 111–114 113
Figure 3: XRD patterns of the sinter-pressed sample and the sintered,
solution-treated, quenched and aged samples
Slika 3: Rentgenski posnetek vzorcev sintrano-stisnjeno in sintrano-
raztopno `arjeno-hitro ohlajeno-starano
Table 1: Variation of the relative density, hardness and electrical conductivity of the test materials as a function of process
Tabela 1: Spreminjanje relativne gostote, trdote in elektri~ne prevodnosti preizku{anih materialov v odvisnosti od postopka
Sample Process Relative density, % Hardness, HV Electrical conduc-tivity, % IACS
Cu-Cr-Zr
Sintered and hot-pressed 91.9 123 76
Solution-treated and water-quenched 92.2 77 77
Cold pressed and
aged (470 oC)
Time, t/h
2 94.5 86 85
4 95.5 96 85
6 95.3 98 86
8 96.2 116 87
5 J. Su, P. Liu, H. Li, F. Ren, Q. Dong, Materials Letters, 61 (2007),
4963–4966
6 J. P. Tu, W. X. Qi, Y. Z. Yang, F. Liu, J. T. Zhang, G. Y. Gan, N. Y.
Wang, X. B. Zhang, M. S. Liu, Wear, 249 (2002), 1021–1027
7 J. Su, Q. Dong, P. Liu, H. Li, B. Kang, Materials Science and Engi-
neering A, 392 (2005), 422–426
8 C. Xia, Y. Jia, W. Zhang, K. Zhang, Q. Dong, G. Xu, M. Wang,
Materials and Design, 39 (2012), 404–409
9 Z. Wang, Y. Zhong, X. Rao, C. Wang, J. Wang, Z. Zhang, W. Ren, Z.
Ren, Transactions of Nonferrous Metals Society of China, 22 (2012),
1106–1111
10 Z. Wang, Y. Zhong, Z. Lei, W. Ren, Z. Ren, K. Deng, J. Alloys and
Compounds, 471 (2009), 172–175
11 http://www.albaksan.com/fr_CuCrZr.htm
M. IPEK: CHARACTERIZATION OF AN Cu-Cr-Zr ALLOY SYNTHESIZED WITH THE POWDER-METALLURGY TECHNIQUE
114 Materiali in tehnologije / Materials and technology 47 (2013) 1, 111–114
E. ALTUNCU et al.: CUTTING-TOOL RECYCLING PROCESS WITH THE ZINC-MELT METHOD
CUTTING-TOOL RECYCLING PROCESS WITH THE
ZINC-MELT METHOD FOR OBTAINING
THERMAL-SPRAY FEEDSTOCK POWDER (WC-Co)
POSTOPEK RECIKLIRANJA ORODIJ ZA REZANJE Z METODO
TALJENJA V Zn ZA PRIDOBIVANJE PRAHU WC-Co ZA
TERMI^NO NABRIZGAVANJE
Ekrem Altuncu1, Fatih Ustel2, Ahmet Turk3, Savas Ozturk4, Garip Erdogan2
1Sakarya University, Tech. Fac., Dept. Metallurgical and Materials Eng. Sakarya, Turkey
2Sakarya University, Eng. Fac., Dept. Metallurgical and Materials Eng. Sakarya, Turkey
3Celalbayar University, Eng Fac., Dept. Metallurgical and Materials Eng. Manisa, Turkey
4Ýzmir K. Çelebi University, Eng Fac., Dept. Metallurgical and Materials Eng. Ýzmir, Turkey
altuncu@kocaeli.edu.tr
Prejem rokopisa – received: 2012-08-16; sprejem za objavo – accepted for publication: 2012-09-19
Various recycling processes for WC-Co cermets from cutting tools, such as chemical modification, thermal modification, the
cold-stream method and the electrochemical method have been investigated and some of them are actually employed in industry.
However, these conventional methods have many problems to be solved and they are not always established technologies.
Therefore, a more economical and high-efficiency recycling procedure needs to be developed. In this study we investigated the
applicability of the zinc-melt method (ZMM) for recycling WC-Co as a powder from cutting-tool scraps. It was proven that
ZMM is an available technique for recovering the WC powder from the cutting tools. WC-Co powders are recovered and then
spray dried, sintered and obtained as a feedstock material for thermal-spray coating processes.
Keywords: cutting tool, WC-Co, recycling, zinc-melt method, spray dryer
Preiskovane so bile razli~ne metode za recikliranje cermetov WC-Co iz orodij za rezanje, kot so kemijska modifikacija,
termi~na modifikacija, metoda s hladnim tokom in elektrokemijska metoda, nekatere od njih pa se `e uporabljajo v industriji.
Vendar pa imajo te konvencionalne metode veliko problemov, ki jih je treba re{iti in nimajo vedno ustaljene tehnologije. Zato je
treba razviti bolj ekonomi~ne in visoko zmogljive postopke recikliranja. V tej {tudiji je bila preu~evana uporabnost metode s
taljenjem odpadkov rezilnih orodij v cinku (ZMM) za recikliranje WC-Co v obliki prahu. Dokazano je bilo, da je ZMM
uporabna tehnika za pridobivanje WC-prahu iz orodij za odrezavanje. WC-Co-prahovi so bili najprej pridobljeni, nato posu{eni
z razpr{evanjem ter sintrani v obliko, primerno za termi~no nabrizgavanje prevlek.
Klju~ne besede: orodje za rezanje, WC-Co, recikliranje, metoda s taljenjem cinka, su{enje z razpr{evanjem
1 INTRODUCTION
The $2-billion, worldwide tungsten-carbide industry
generates large quantities of scrap due to the parts
rejected at various stages of the production and the
worn-out cutting tools. The most basic recycling
approach would be to break down the scrap pieces into
powders and then fabricate more WC-based cutting
tools. This approach would cause a severe equipment
wear due to the abrasive nature of the cutting-tool mater-
ials and is, therefore, not feasible. As a result, the
recycling is done by chemical means, such as the zinc-
recovery process, electrolytic recovery, and extraction by
oxidation. The conventional recycling processes for
cutting tools have many problems to be solved, so they
are not efficient technologies. One of them takes a very
long processing time and the other requires high-scale
equipment. Another process leads to decomposition and
an undesirable phase occurrence. The recycling rate of
WC-Co is only about under 20 % and the rates are still
low1–5. Nowadays, the approaches to recovering WC and
Co materials are not only economically important but
also, due to the environmental factors, ecologically
significant. Globally, one third of the consumption of
WC for the cutting tools is being produced from their
scrap. In recent years, a new technique for recovering
WC and Co from the hard cutting-tool scraps using a
molten-zinc (Zn) bath has been studied. Due to its
efficiency and applicability, the zinc-melt process is
considered to have a higher potential for recycling the
cutting tools3. In this study, the zinc-melt-method
parameters, such as temperature and time were
optimized. A biscuit-structured WC-Co bulk was ground
to obtain fine powders and then mixed with a Co powder
by ball milling. These powders were produced with a
spray dryer, having a spherical form and utilizing the
thermal-spray process (Figure 1). The microstructure
and chemical properties of these powders were studied
with SEM and XRD.
2 EXPERIMENTAL WORK
The powder microstructures and surface morpho-
logies were examined with the scanning electron
microscopy (SEM). An elemental distribution analysis of
the recovered powder was conducted with the energy-
Materiali in tehnologije / Materials and technology 47 (2013) 1, 115–118 115
UDK 621.9.02:628:477.6 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(1)115(2013)
dispersive X-ray spectroscopy (EDS). The crystalline
phases were identified with the X-ray diffractometry
(XRD). The density of the sintered body was measured
in water using the Archimedes method. The flowability
was measured with a Hall flowmeter. The particle-size
distribution was measured with a laser particle sizer.
3 ZINC-MELT-METHOD (ZMM) RECYCLING
PROCESS
In a zinc-recovery process, cemented carbide scraps
are immersed in molten zinc in an electrical furnace at
1 atmosphere of inert gas at 650–800 °C. The zinc is
subsequently distilled at 700–950 °C.6,7 The optimum
conditions depend on the Co content and the zinc-to-
cobalt ratio. The properties of the reclaimed powders are
the same as the virgin powders. Scrap cemented carbides
can be sorted into medium (1.2–2 mm), coarse (4 mm),
and mixed grain sizes with optical microscopy and by
composition with x-ray spectroscopy before the zinc-
recovery process. When the WC-Co scraps are dipped in
a molten-zinc bath, the molten zinc starts to penetrate the
WC particles and the dissolving Co binder, i.e., the Co
film between the WC particles. The dissolution of the Co
film results in a separation of the WC particles and
independent particles float in the molten Zn-Co alloy.
Since zinc has a high vapor pressure, it is easily removed
in vacuum condition after the separation of WC particles.
As zinc evaporates under vacuum, the Co content
gradually increases and the Co metal has to precipitate
on the WC surface from the molten alloy. When zinc
evaporates completely, all Co precipitates on the WC
particle surface resulting in the WC-Co powders with the
same chemistry as found in the original powder before
E. ALTUNCU et al.: CUTTING-TOOL RECYCLING PROCESS WITH THE ZINC-MELT METHOD
116 Materiali in tehnologije / Materials and technology 47 (2013) 1, 115–118
Figure 1: Experimental route of recycling scrap cutting tools
Slika 1: Eksperimentalna pot recikliranja odpadkov rezilnih orodij
Figure 3: Recycling process of WC-Co
Slika 3: Postopek recikliranja WC-Co
Figure 2: Zinc-melt process8
Slika 2: Postopek s taljenjem v cinku8
the sintering8 (Figures 2 and 3). By increasing the
process temperature and time, the recycling efficiency
was improved. The zinc contamination was reduced
(Table 1).
Process specifications: Process temperature: 700–900 °C
under Ar+N2 atmosphere, Melting time: 1–10 h, Charge:
14–22 kg (Scrap/Zn:1/ 1.3), Efficiency: 82–97 %.
4 SPRAY DRY PROCESS
Spray drying which is the most versatile powder-
processing method consists of spraying a water-based
suspension (called slurry) of the materials to agglome-
rate into a stream of heated air. The rapid heat and mass
transfer which occurs during the drying combined with
the presence of various slurry compounds result in dried
granules having a large variety of shapes – from the
uniform solid spheres, which are regarded as ideal
granules for most ceramic systems, to elongated, pan-
cake, donut-shaped, needle-like or hollow granules9–11.
The spray-drying process transforms a solution with a
certain solid content into a powder of the solid in one
step. In this study we used volume fractions 25 %
WC-Co + water solution (slurry). The solid aggregates
are collected at the bottom of the chamber and separated
from the gas in cyclone collectors. The main controlled
operating parameters are the air temperature at the entry
(185–210 °C), at the exit (100–140 °C) and inside the
chamber (165–180 °C), the atomizing nozzle design and
the air- and slurry-flow rates. The flowability of thermal
spraying is good. The mean particle size is D50  48 μm.
According to the X-ray diffraction (XRD) investigations
E. ALTUNCU et al.: CUTTING-TOOL RECYCLING PROCESS WITH THE ZINC-MELT METHOD
Materiali in tehnologije / Materials and technology 47 (2013) 1, 115–118 117
Table 1: Biscuit structures after different thermal treatments
Tabela 1: Kola~i, dobljeni po razli~nih termi~nih obdelavah
Figure 4: Spray-dried powders and deposition
Slika 4: Prah, osu{en z razpr{evanjem in po depoziciji
of the spray-dried feedstock presented earlier, WC, WC2,
WO3, CoWO4 and metallic Co were present. The feed-
stock powder consisted of a-few-micrometers-sized
agglomerates where the WC particles of 1–5 μm were
finely distributed in the Co matrix.
Figure 4 shows the variety of shapes: from uniform
solid spheres to elongated, pancake, donut-shaped,
needle-like or hollow granules. It was found that the
higher the inlet temperature, the faster is the moisture
evaporation. A higher outlet temperature leads to a larger
size of the powder. And the outlet temperature also
controls the final moisture content of the powder. As the
viscosity is lowered, less energy or pressure is required
to form a particular spray pattern. Care must be taken
with high solid loadings to maintain proper atomization
ensuring a correct droplet formation. Then the
spray-dried WC-Co feedstock powder was sintered at
1100 °C for 8 h to attain a sufficient granule strength,
and then classified to a diameter of <30 μm. As the
sintering temperature increased, the peak intensity of
eta-carbide increased as well, while, simultaneously, the
peak intensity of the WC and W2C phases decreased.
Depending upon the particle size and the shape of the
tungsten-carbide-cobalt, spray-dried powder, the
flowability can be under control. The highest flowability
is obtained with the optimum spray-drying parameters.
The measured Hall-flow rate of the reprocessed powder
is approximately 160 g/min.
5 DEPOSITION PROCESS
Using the thermal-spray processing to deposit a coat-
ing of powders provides for a high-rate deposition
method allowing an effective pressure and the tempe-
rature required for sintering high-density powders.
High-velocity, oxy-fuel (HVOF), thermal-spray WC/Co
coatings have been used widely in many fields such as
metallurgy, energy sources and construction industry
where they can be subjected to severe abrasive wear due
to their excellent abrasive-wear resistance12–16. Spraying
was performed at Sulzer Metco Robot Controlled
Coating System during an HVOF process using a hydro-
gen-fuelled gun (DJ2600). The coating was free of any
macroscopic porosities or cracks exhibiting excellent
bonding to the substrate and a dense structure. The
optimum spray parameters were determined from the
preliminary experiments in order to minimize the coating
porosity and WC decomposition (Figure 4d).
6 CONCLUSION
Our recycling and modification of the powder
processes have been successful. The obtained granulated
powders can be used for the thermal-spray feedstock
materials. With respect to the zinc-melt process, the
bonding metal (cobalt) reacts with the high-purity zinc
during the cemented-carbide-recovery operation. The
zinc contamination is low. The zinc-melt process shows
a higher efficiency than the other process. The process
parameters were optimized according to the shape,
crystallinity and flowability properties. These powders
can be easily used in the HVOF process. However, prior
to starting further production of the recycled WC/Co
powders, it will have to be demonstrated that this is a
cost-effective production.
Acknowlodgements
The author is thankful to the members of the Sakarya
University Thermal Spray Laboratory.
7 REFERENCES
1 T. Kojima, T. Shimizu, R. Sasai, H. Itoh, Recycling process of
WC-Co cermets by hydrothermal treatment, J. of Materials Sci., 40
(2005), 5167–5172
2 J. C. Lee, E. Y. Kim, J. H. Kim, W. Kim, B. S. Kim, D. P. Banshi,
Recycling of WC-Co hardmetal sludge by a new hydrometallurgical
route, Int. J. of Refractory Metals and Hard Materials, 29 (2011)
3, 365
3 W. D. Venkateswaran, B. Schubert, M. Lux, B. D. Ostermann, W.
Kieffe, Scrap recycling by the melt bath technique, Int. J. of Ref.
Metals and Hard Materials, 14 (1996) 4, 263–270
4 C. Edtmaier, R. Schiesser, C. Megssl, W. D. Schubert, A. Bockb, A.
Schoen, B. Zegler, Selective removal of the cobalt binder in WC/Co
based hardmetal scraps by acetic acid leaching, Hydrometallurgy, 76
(2005), 63–71
5 R. Sasai, F. Inagak, H. Itoh, Resource Recovery from Cemented
Carbide by Subcritical Hydrothermal Treament, J. of the Soc. of.
Mat. Sci., Japan, 55 (2006) 3, 254–257
6 K. Hirose, I. Aoki, Recycling Cemented carbides without Pollut ion
– Sorting Charging Material for Zinc Process, Int. Conf. Process.
Mater. Prop., 1st 1993, 845–848 (English). Edited by Henein, Hani;
Oki, Takeo. Miner. Met. Mater. Soc: Warrendale, PA
7 A. K. Maiti, N. Mukhopadhyay, R. Raman, Effect of adding WC
powder to the feedstock of WC–Co–Cr based HVOF coating and its
impact on erosion and abrasion resistance, Surface and Coatings
Technology, 201 (2007), 7781–7788
8 Kohsei Co., Kitakyushu plant, http://www.kohsei.co.jp, 21. 05. 2012
9 W. J. Walker Jr., J. S. Reed, S. K. Verma, Influence of Slurry Para-
meters on the Characteristics of Spray-Dried Granules, J. of the
American Ceramic Society, 82 (1999) 7, 1711–1719
10 D. E. Walton, C. J. Mumford, Spray dried products—characterization
of particle morphology, Transactions of the Institution of Chemical
Engineers, 77 (1999), 21–38
11 Y. Wang, S. Jiang, M. Wang, S. Wang, T. D. Xiao, P. R. Strutt,
Abrasive wear characteristics of plasma sprayed nanostructured
alumina/titania coatings, Wear, 237 (2000), 176–185
12 M. C. Nyongesa, Preparation, Characterisation And Testing Of
Wc-Vc-Co Hp/Hvof Thermal Spray Coatings, School of Process and
Materials Engineering – Faculty of Engineering and the Built
Environment, WITS library, PHd thesis, 2006
13 D. A. Stewart, P. H. Shipway, D. G. McCartney, Abrasive wear beha-
viour of conventional and nanocomposite HVOF-sprayed WC–Co
coatings, Wear, 225–229 (1999), 789–798
14 R. Joost, J. Pirso, M. Viljus, S. Letunovit{, K. Juhani, Recycling of
WC-Co hardmetals by oxidation and carbothermal reduction in
combination with reactive sintering, Estonian J. of Eng., 18 (2012) 2,
127–139
15 M. G. Park, J. K. Han, Processing of Reclaimed Tungsten Carbide/
Cobalt Powders with Virgin Properties, Taehan Kumsok Hakhoechi,
30 (1992) 8, 996–1004 (Korean)
16 E. Altuncu, S. Ozturk, F. Ustel, Recycling of WC-Co from cutting
tools by zinc melt method, 16th International Metallurgy & Materials
Congress, TÜYAP Fair, Ýstanbul 2012
E. ALTUNCU et al.: CUTTING-TOOL RECYCLING PROCESS WITH THE ZINC-MELT METHOD
118 Materiali in tehnologije / Materials and technology 47 (2013) 1, 115–118
S. KARABAY et al.: PERFORMANCE TESTING OF AN OPTICAL GROUND WIRE COMPOSITE
PERFORMANCE TESTING OF AN OPTICAL GROUND
WIRE COMPOSITE
PREIZKU[ANJE ZMOGLJIVOSTI KOMPOZITNEGA
PODZEMNEGA OPTI^NEGA KABLA
Sedat Karabay, Ersin Asim Güven, Alpay Tamer Ertürk
Kocaeli University, Mechanical Engineering Department, Kocaeli, Turkey
sedatkarabay58@gmail.com
Prejem rokopisa – received: 2012-08-27; sprejem za objavo – accepted for publication: 2012-09-26
An optical ground wire (OPGW) composed of different materials was presented in detail before delivering the product to the
communication and electrical-energy markets. The performance level of its composite structure differs from the performance of
the original material. Therefore, to measure whether the OPGW had reached the required quality, it was exposed to several tests
simulating the real working conditions to detect the behavior of the composite structure. These included the
stress-strain/fibre-strain and tensile tests, aeolian vibration, galloping, creep, short circuit, temperature cycling and lightning
tests. Thus, the technical story of OPGW designed to serve the environment was explained in details and the test results were
interpreted. The required material improvements to the master alloys made due to the failures of the composite conductor
(OPGW) under heavy test conditions were also explained so that approval could be obtained.
Keywords: OPGW, lightning strike, creep, aeolian, composite structure, fiber failure
Opti~ni podzemni kabel (OPGW), sestavljen iz razli~nih materialov, je bil predstavljen do podrobnosti, preden je bil proizvod
poslan na trg komunikacij in energije. Zmogljivost kompozitne strukture se razlikuje od osnovnega materiala. Da bi preizkusili
zmogljivost kompozitnega materiala in izmerili, ali OPGW dose`e sprejemljivo kvaliteto, je bil kabel izpostavljen razli~nim
preizkusom, ki so simulirali realne razmere. To so natezna napetost – raztezek vlakna, eolianske vibracije, galopiranje, lezenje,
kratek stik, spreminjanje temperature in preizkus z bliskanjem. Podrobno je predstavljena celotna zgodba razvoja OPGW in
vpliva okolja, razlo`eni pa so tudi rezultati preizkusov. Razlo`ene so izbolj{ave osnovnih zlitin, potrebne zaradi napak v
kompozitnem prevodniku (OPGW), da bi se doseglo soglasje za uporabo.
Klju~ne besede: OPGW, udar strele, lezenje, eolianske vibracije, kompozitna struktura, poru{itev vlaken
1 INTRODUCTION
The OPGW cable today forms an integral part of any
power company’s transmission network and is utilized
for the mission-critical circuit control ensuring the
optimum operational efficiency and protection. The
OPGW cable is defined as a composite cable which
serves as a conventional overhead ground wire with the
added benefit of providing the optical-fiber communi-
cations. An optical communication carrier can be
completely separated from the power-transmission line
to form an additional revenue stream, whilst the cable
serves the traditional purpose of conducting fault
currents to the ground and protecting the power con-
ductors against lightning strikes. With proper design
considerations, the OPGW cable has proven its reliabi-
lity in protecting the optical fibers from electrical,
mechanical, and environmental stresses1. With an
increasing demand for more information transmission,
such as the widespread use of the internet in the recent
years, a much higher fiber-count OPGW is needed. The
previous study shows the structure and the main test
result of a stainless-steel tube with optical fibers in the
stainless-steel tubes used instead of the conventional
aluminum pipes in an extra-high multi-core OPGW2.
Thus, a technical story of OPGW designed for a long-
term reliability in the real working conditions was
explained and the test results interpreted. Remedies for
the failures of the composite conductor (OPGW) under
heavy test conditions required material improvements
that were also explained to reach approval in field
conditions.
The composite conductor used in the experiments is
composed of one stainless-steel tube with fibers, six
galvanized steel wires and 12 AA–6101 aluminum-alloy
wires at the outer layer (Figure 1)3.
2 CHARACTERISTICS OF OPGW IN A
STAINLESS-STEEL TUBE
2.1 Stress-Strain/Fiber-Strain and Tensile Test
The objective of this test is to monitor the optical
characteristics and verify the mechanical characteristics
of OPGW under the test up to the breaking load. An
OPGW sample was installed in a hydraulically activated,
horizontal test machine4–7. A displacement transducer
was fixed to the conductor to measure the cable
elongation over the 8 m gage length. The gage length for
attenuation measurements was taken for the length under
tension. The conductor elongation, the output signal
from the optical power meters and the conductor tension
as measured by a load cell were monitored using a
digital-data logging system. After completing the tests, if
Materiali in tehnologije / Materials and technology 47 (2013) 1, 119–124 119
UDK 66.017:620.168 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(1)119(2013)
the fiber elongation at 0.45 % of conductor elongation is
greater than 0.01 % and if at 72 % of RTS the temporary
increase in the fiber attenuation is greater than 1 dB/km,
as compared to the value measured before the test, and
there is a measurable permanent increase in the fiber
attenuation that is greater than 0.02 dB/km, the test shall
be considered unsuccessful. Some results of the com-
pleted test are presented in Figures 2 and 3. Figure 2
shows the data at the load (conductor tension) plotted
against the conductor strain. On the other hand, Figure 3
shows optical attenuation and the load/conductor tension
plotted against time.
2.2 Creep Test
The objective of the creep test is to measure the
room-temperature, long–term tensile-creep properties of
the conductor. The data from this test are used to assist in
the calculation of the sags and tensions. The test was
performed according to IEC 61395. The length of the
sample between the dead-end clamps was 15 m. The test
was carried out in a temperature-controlled laboratory at
20 °C ± 2 °C. In line with the Aluminum Association’s
method, the long-term tensile creep of the cable under a
constant tension is taken to be the permanent strain
occurring between 1 h and the specified test time.
The last reading during this test was taken at 1000 h.
A log-log plot of strain versus the elapsed time for
LVDT (the linear variable differential transformer) is
shown in Figure 4. On the completion of the test, the
best-fit straight line was fitted to the LVDT data and
extrapolated to 10 years (87000 h).
The equation of the line:
Strain = A*(Hours)B  y = 8.5073E – 0.5 X1.3965E–01
2.3 Temperature-Cycle Test
The objective of this test was to verify the good
performance of the fiber when the cable is subjected to
extreme thermal cycles. The test was performed in
accordance with EIA/TIA-455-3A. A reel with approxi-
mately 761 m of an OPGW cable was placed in a 5 m ×
6 m × 4 m environmental chamber. Three thermocouples
were placed in the environmental chamber to measure
the temperature. Two were placed on separate 25 cm
cable samples and located on either side of the cable
reel. The third was located under the first layer of the
cable reel. All twenty-four fibers were spliced to form
one continuous loop. The total test-fiber length was
approximately 18.3 km. The cable was subjected to two
thermal cycles. Each thermal cycle started with the
chamber temperature of 23 °C, which was then lowered
to –40 °C and held at this level for a minimum of 16 h.
The chamber temperature was then increased to 65 °C
and held at this level for a minimum of 16 h. To com-
S. KARABAY et al.: PERFORMANCE TESTING OF AN OPTICAL GROUND WIRE COMPOSITE
120 Materiali in tehnologije / Materials and technology 47 (2013) 1, 119–124
Figure 3: Optical attenuation plotted against time
Slika 3: Prikaz opti~nega du{enja v odvisnosti od ~asa
Figure 1: Design properties of OPGW
Slika 1: Sestav OPGW
Figure 4: Cable strain versus time
Slika 4: Raztezanje kabla v odvisnosti od ~asa
Figure 2: Load plotted against conductor strain
Slika 2: Obremenitev v odvisnosti od raztezka prevodnika
plete the cycle, the chamber temperature was returned to
23 °C. All the temperature transitions were conducted at
a rate of less than 20 °C/h. The chamber temperature was
based on one of the thermocouples on the 25 cm cable
samples, located on one side of the cable reel. The
cable-reel temperature and the optical data were recor-
ded every five minutes throughout the test.
The optical attenuation and the chamber temperature
versus time are shown in Figure 5. The variation in the
optical attenuation due to the temperature was no greater
than 0.006 dB/km. The maximum allowable change in
the attenuation, between the extreme temperature limits,
is 0.05 dB/km.
2.4 Aeolian-Vibration Test
The objective of the aeolian-vibration test is to assess
the fatigue performance of OPGW and the optical
characteristics of the fibers under typical aeolian
vibrations. The tests were performed according to IEC
60794-1-2, Method E19 and IEC 60794-4-1. Thus,
OPGW was pre-tensioned to 1795 N and an initial
optical measurement was taken. OPGW was then
tensioned to 17 903 N or 20 % of the RTS cable and the
exit angles of the cable from the suspension clamp were
measured. The initial target vibration frequency was 54.4
s–1, which is the frequency produced by a 4.5 m/s wind
(i.e., frequency = 830 ÷ diameter of OPGW in mm). The
actual vibration frequency was the system resonance that
was nearest to the target frequency and provided a better
system stability, while the target free-loop peak-to-peak
antinode amplitude was 5.08 mm or one third of the
OPGW diameter. This amplitude was maintained at this
level in the first free loop from the suspension assembly
towards the shaker. The amplitudes in the passive span
and the section between the shaker and the dead end in
the active span were maintained at the levels no greater
than one third of the cable diameter. OPGW was
subjected to 10-million vibration cycles. Optical
measurements were taken for 2 h after the completion of
the vibration cycles.
All twenty-four fibers were spliced to make the total
fiber length of 720 m (24 × 30 m). The test sample was
terminated beyond both dead ends so that the optical
fibers could not move relative to OPGW.
Dissection: After the completion of 10 million cycles,
the cable was dissected down to the stainless-steel tube
and visually examined. Active dead end: There were no
visible signs of breaks, cracks, failure or discoloration of
any of the dissected components of OPGW (Figure 6).
Passive dead end: There were no visible signs of breaks,
cracks, failure or discoloration of any of the dissected
components of OPGW. Suspension: There were no
visible signs of breaks, cracks, failure or discoloration of
any of the dissected components of OPGW.
2.5 Galloping Test
The objective of the galloping test is to assess the
fatigue performance of the fiber optical ground wire and
the optical characteristics of the fibers under typical
galloping conditions. The test was performed according
to IEC 60794-4-1. For that aim, an initial optical
measurement was taken one hour prior to the test. The
difference between the reference and test signals for the
initial measurement provided an initial base reading. The
change in this difference during the test indicated the
change in the attenuation of the test fiber. The cable was
subjected to 100 000 galloping cycles in the single-loop
mode. The free-loop peak-to-peak antinode amplitude
was maintained at the minimum of about 0.8 m or 1/25th
of the distance from the dead end to the suspension-
clamp length (i.e., 20 m). Optical measurements were
taken for two hours after the completion of the galloping
test. The galloping frequency at the start and during the
test was 1 s–1 without any variations. The free-loop
antinode amplitude in the active (driven) span was main-
tained at approximately 0.9 m. The free-loop antinode
amplitude in the passive span varied between 0.3 m to
0.4 m during the test. The tension in the cable fluctuated
between 529 N to 1432 N during the galloping. After the
completion of 100 000 cycles, the cable was dissected
S. KARABAY et al.: PERFORMANCE TESTING OF AN OPTICAL GROUND WIRE COMPOSITE
Materiali in tehnologije / Materials and technology 47 (2013) 1, 119–124 121
Figure 6: Aeolian-vibration test results
Slika 6: Rezultati preizkusa eolijskih vibracij
Figure 5: Records of the heat-cycle test of OPGW
Slika 5: Zapis iz preizkusa cikli~nega segrevanja OPGW
down to the stainless-steel tube and visually examined.
Active dead end: There were no visible signs of breaks,
cracks, failure or discoloration of any of the dissected
components of OPGW. Passive dead end: There were no
visible signs of breaks, cracks, failure or discoloration of
any of the dissected components of OPGW (Figure 7).
Suspension: There were no visible signs of breaks,
cracks, failure or discoloration of any of the dissected
components of OPGW.
2.6 Short-Circuit Test
The objective of the short-circuit test is to verify if
OPGW can withstand repeated short-circuit applications
without exceeding optical, physical or thermal require-
ments. The test was performed in accordance with the
TEIAS Specification and IEC 60794-1-2, Method H1.
The cable was first subjected to two low-level
calibration shots and then ten "official" shots. The
purpose of the calibration shots was to ensure that the
current level was correct. For the "official" shots, the
target values for the electrical parameters were: the
parameter target energy value of 109.5, the minimum
kA2s–1 fault current of 14.8 kA, the duration of 0.5 s, the
maximum possible asymmetric waveform to be symme-
trical after the 3rd cycle for each shot, the fault current
and the duration may vary slightly from the target values.
The objective was to achieve the minimum energy level
for each shot. To ensure that optical signals were stable,
the power meters were powered on and operating for at
least one hour before the first shot. The optical measure-
ment was normalized to zero before the first official shot.
The cables were visually inspected for birdcaging or
other damage during the test. The optical and tempe-
rature data were being acquired for one hour after the
tenth shot. The cable was maintained at the temperature
of 40 °C during this period.
As specified by IEC 60794-1-2, Method H1, the
acceptance criteria of the product are summarized below:
a) The temperature immediately after the current pulse
shall be less than 180 °C inside the optical unit. The
temperature inside the optical unit is measured by
thermocouple.
b) The attenuation increase during the tests shall be less
than 1.0 dB/km. There shall be no change in the
attenuation after the cable has cooled down to 40 °C.
c) There shall be no irreversible birdcaging. The cable
and hardware shall be dissected after the test and
visually examined for damage at each dead-end
assembly and at the midpoint of the span. Each
separable component of the cable shall be inspected.
There shall be no signs of birdcaging, excessive wear,
discoloration, deformation or other signs of a break-
down (Figures 8 and 9).
2.7 Lightning Tests
The essential function of OPGW in transmission
lines is to guard the aerial conductor from the lightning
strikes and its secondary job is to transmit the signals of
data and communications. The excessive lightning
energy generally flows through the outer layer of the
OPGW conductor. However, when this energy jumps
from the clouds to an OPGW conductor, a small region
on the outer layer is liable to overheating. Therefore, the
conductivity of the material used in an OPGW
conductor, including both electrical and heat transfer,
should be as high as possible. If not, regional melting
occurs as seen in Figures 10 a, b and finally the wires
are broken. These Figures 10 a, b refer to the first trial
of the lightning test. In this test, 2 × 10 m OPGW sam-
ples, connected in parallel to measure the attenuation of
the fibers and the effects of the overcurrent, failed due to
a lightning strike.
The test was realized under an amplitude of 200 A,
with a charge of 100 C and within the time of 500 ms.
The main acceptance restriction is to keep the resistance
increase below a 20 % change. This corresponds to three
wires breaking at the outer layer. However, at first the
trial 9–10 wires were broken.
S. KARABAY et al.: PERFORMANCE TESTING OF AN OPTICAL GROUND WIRE COMPOSITE
122 Materiali in tehnologije / Materials and technology 47 (2013) 1, 119–124
Figure 7: Galloping-test records of OPGW and attenuation of fibers
Slika 7: Zapis preizkusa galopiranja OPGW in opti~no slabljenje
vlaken
Figure 8: Applied short-circuit arc current and its time
Slika 8: Uporabljeni tok kratkega stika in njegovo trajanje
Figure 9: Attenuation range for OPGW when a lightning strike is
applied
Slika 9: Podro~je opti~nega slabljenja vlaken OPGW pri uporabi
udarca strele
Therefore, the AA–6101 aluminum alloy was modi-
fied with AlB2 (Figure 11) at the casting stage and then
the conductivity of the wires increased from 52.5 %
IACS to 57–58 % IACS.3 The second test was thus
performed with modified wires and stranded with a short
lay length to obviate the arc between the wires. The
second trial met all the requirements perfectly (Figures
12 a, b).
3 RESULTS
Before a new OPGW product, prepared with a com-
bination of different materials, is introduced to the
market, it should be exposed to several tests to determine
its mechanical and electrical behaviors under simulated
working conditions. Here, the required important tests
were applied to the OPGW composite structure. The
tested product passed most of them perfectly, but the
lightning test destroyed it completely. Therefore, the
designed and constructed composite structure should be
changed or the conductive material must be modified.
4 DISCUSSION
The initially designed and constructed OPGW
conductor successfully passed most of the tests defined
previously, except for the lightning test. As a remedy,
AlB with AlB2 phases of the master alloy was fed into
the molten AA–6101 alloy as 3 kg per ton. Then conduc-
tivity of the AA–6101 wires increased from 52 % IACS
to 57–58 % IACS. An increase in electrical conductivity
also causes an increase in heat conductivity. When the
above modification is applied other properties such as
tensile strength, elongation, 1 % elongation strength, etc.
remain constant.
S. KARABAY et al.: PERFORMANCE TESTING OF AN OPTICAL GROUND WIRE COMPOSITE
Materiali in tehnologije / Materials and technology 47 (2013) 1, 119–124 123
Figure 12: a) Lightning arc, b) view after a strike to the OPGW con-
ductor without broken wires on the outer layer
Slika 12: a) Oblok bliska, b) po udaru v OPGW prevodnik brez poru-
{enih `ic na zunanji strani
Figure 11: AlB2 master alloy used to increase the conductivity of the
AA–6101 aluminum alloy by inoculating it in the casting stage in a
foundry tandish3
Slika 11: Osnovna zlitina AlB2, uporabljena za pove~anje prevodnosti
aluminijeve zlitine AA-6101 z inokulacijo med ulivanjem v livarski
vmesni posodi3
Figure 10: a, b) Spot melting of AA-6101 aluminum-alloy wires due
to an application of lightning strike
Slika 10: a, b) To~kasto taljenje aluminijeve `ice AA-6101 zaradi
udarca strele
By modifying the wires and reducing the lay length
of the conductor, a new test sample was manufactured.
Then a lightning strike was applied again. Now the
results met the requirements and the standard used in the
test and the product passed perfectly all the required
tests.
5 CONCLUSION
The design, construction and modification of an
OPGW aerial conductor using a combination of different
materials is presented by explaining its behavior under
type tests before introducing it into the communication
and energy markets. This research also shows that a
lightning strike is the hardest test applied to the con-
ductor. OPGW can resist the lightning strike when we
strand the wires tightly, decrease the lay length and
increase electrical conductivity by inoculating it with
AlB2 in a tundish at 750 °C. The short-circuit test applied
to OPGW is not the only way of determining its perfor-
mance in the event of lightning.
6 REFERENCES
1 M. Yokoya, Y. Katsuragi, Y. Goda, Y. Nagata, Y. Asano, IEEE
Transaction on Power Delivery, 9 (1994), 1517–1523
2 F. Jakl, IEEE Transaction on Power Delivery, 15 (2000), 1524–1529
3 S. Karabay, Y. Tayþý, Materials and Manufacturing Process, 19
(2004), 1–13
4 T. Torvath, Journal of Electrostatics, 60 (2004), 265–275
5 S. J. Guavac, M. D. Nimrihter, L. R. Geric, Electrical Power System
Research, 78 (2007), 556–583
6 D. Ruiz, C. Torraba, Journal of Electrostatics, 67 (2009), 496–500
7 Y. Goda, Y. Shigenu, S. Watanabe, IEEE Transaction on Power
Delivery, 19 (2004), 1734–1739
S. KARABAY et al.: PERFORMANCE TESTING OF AN OPTICAL GROUND WIRE COMPOSITE
124 Materiali in tehnologije / Materials and technology 47 (2013) 1, 119–124
J. @UPAN et al.: INVESTIGATION OF THE COOLING PROCESS WITH NANOFLUIDS ...
INVESTIGATION OF THE COOLING PROCESS WITH
NANOFLUIDS ACCORDING TO ISO 9950 AND ASTM D
6482 STANDARDS
PREISKAVA POSTOPKA OHLAJANJA V NANOTEKO^INI PO
STANDARDIH ISO 9950 IN ASTM D 6482
Josip @upan, Darko Landek, Tomislav Filetin
University of Zagreb, Faculty of Mechanical Engineering and Naval Architecture,
Quenching Research Centre (QRC), I. Lucica 5, 10000 Zagreb, Croatia
josip.zupan@fsb.hr
Prejem rokopisa – received: 2012-09-01; sprejem za objavo – accepted for publication: 2012-09-19
Introducing nanofluids as liquid quenchants in order to improve the heat-transfer characteristics is a novel approach in heat
treatment. These new liquid quenchants are called nanoquenchants and are colloid suspensions of nanoparticles in the base fluid
(BF). For the purpose of this research standard liquid quenchants, such as water and polymer solution, were used as base fluids.
Nanoparticles were added to the base fluid in order to increase its thermal properties without a significant effect on the viscosity
of the fluid. The added nanoparticles cause an enhancement of the heat-transfer characteristics of the liquid quenchants. In this
research TiO2 nanoparticles, with the average size of 50 nm, were added to the base fluid. The cooling curves for every tested
quenchant were recorded using the IVF SmartQuench system and the quenching process parameters were determined and
compared. The tested quenchants were deionised water and polyalkylene-glycol (PAG) water solutions of (5, 10 and 20) % in
volume fractions of polymer concentrations. Quenching experiments were first conducted in pure BFs without the addition of
nanoparticles and without agitation according to the ISO 9950 standard. The experiments were then repeated with an addition of
0.2 g/l TiO2 nanoparticles. The second series of experiments was conducted in a quenching bath with agitation according to the
ASTM D 6482 standard. All of the recorded cooling curves were compared and the effects of the nanoparticle addition and
agitation were investigated. The addition of the nanoparticles and quenching with agitation caused an increase in the maximum
cooling rate and a shorter full-film stage.
Keywords: quenching, nanofluids, agitation, TiO2 nanoparticles
Uvajanje nanoteko~in za kaljenje za izbolj{anje lastnosti prenosa toplote je nov na~in toplotne obdelave. Te nove teko~ine za
kaljenje se imenujejo nanokalilne teko~ine, ki so koloidna suspenzija nanodelcev v osnovni teko~ini (BF). Za preiskavo je bila
kot standardna teko~ina za kaljenje uporabljena me{anica vode in raztopine polimera. Nanodelci so bili dodani osnovni teko~ini
za pove~anje termi~nih lastnosti brez velikega u~inka na viskoznost teko~ine. Dodani nanodelci so pove~ali toplotno prevodnost
teko~ine za kaljenje. V tej raziskavi so bili dodani osnovni teko~ini nanodelci TiO2 povpre~ne velikosti 50 nm. Krivulje
ohlajanja za vsako preizku{eno kalilno teko~ino so bile posnete s sistemom IVF SmartQuench in procesni parametri kaljenja so
bili dolo~eni in primerjani. Preizku{ene kalilne teko~ine so bile deionizirana voda, vodna raztopina polialkilen glikola (PAG) z
volumenskim dele`em (5, 10 in 20) % polimera. Preizkusi kaljenja so bili najprej izvr{eni v osnovni teko~ini (BF) brez dodatka
nanodelcev in brez preme{avanja, skladno s standardom ISO 9950. Eksperimenti so bili nato ponovljeni z dodatkom
nanodelecev 0,2 g/l TiO2. Druga serija preizkusov je bila izvr{ena v kalilni kopeli s preme{avanjem, kot dolo~a standard ASTM
D 6482. Vse posnete krivulje so bile primerjane in preiskovan je bil tudi u~inek dodatka nanodelcev in preme{avanja. Dodatek
nanodelcev in kaljenje s preme{avanjem sta povzro~ila pove~anje maksimalne hitrosti hlajenja in kraj{e stanje popolnega
prekritja s paro.
Klju~ne besede: kaljenje, nanoteko~ine, me{anje, nanodelci TiO2
1 INTRODUCTION
Nanofluids (NFs) are colloidal suspensions of a base
fluid (BF) and particles that are usually less than 100 nm
in diameter1. The particles are added with the aim of
improving the thermal properties of the pure base fluids,
which primarily refers to the significant increase in the
heat-transfer dynamics of the cooling process. It is
important that the rheological properties of NFs do not
change much2–4. Base fluids can be water, ethylene gly-
col, polyalkylene glycol (PAG) or quenching oil, while
metal nanopowders (Cu, Au, Ag), oxide-ceramic parti-
cles (Al2O3, SiO2, TiO2), carbon powders and nanotubes
are added to a base fluid. This is the reason why
nanofluids are interesting as liquid quenchants. Physical
properties of a quench-hardened work piece highly
depend on the cooling rate during the quenching process.
In order to achieve specific properties, different cooling
rates have to be applied5. So far, several studies were
conducted to demonstrate the efficiency of nanofluids as
liquid quenchants6–9. The experiments showed that
nanofluids exhibit a higher heat-transfer coefficient
(HTC) and a better thermal conductivity than base fluids,
but also a shorter full-film boiling phase. Further analy-
sis of the cooling curves showed an increase in the criti-
cal heat flux (CHF) caused by an addition of nanoparti-
cles to BF.
So far, all of the experiments were conducted accord-
ing to the ISO 9950 standard. There are no published
results regarding the quenching in nanofluids with
mechanical agitation. For this reason a series of experi-
ments according to the ASTM D 6482 standard were
Materiali in tehnologije / Materials and technology 47 (2013) 1, 125–127 125
UDK 621.78.084:620.3 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(1)125(2013)
conducted to compare the cooling curves of BFs and
NFs, but also to see the effect of agitation.
2 EXPERIMENTS
A standard test probe used in this set of experiments
is in compliance with the international standards and is a
part of the IVF SmartQuench system consisting of a
data-acquisition unit, a certified standard test probe, a
furnace and software. The data acquisition unit and
software have the capacity of gathering 100 samples per
second, which is considered to be high enough for
registering all the rapid changes that occur during the
quenching process. The experiments with still fluids
were conducted in a one-litre beaker, and a two-litre
quenching bath was constructed for the experiments with
agitation. The electric-motor rotational speed is set to
1000 r/min and agitation is induced with a three-blade
propeller. For the preparation of NFs, the titanium-oxide
(TiO2) particles produced by Degussa, Germany, were
used. The average size of the nanoparticles was 50 nm,
and they come in the form of a white powder. No disper-
sants, surfactants or activating agents were used. Pure
TiO2 nanoparticles were added to BFs and homogenized.
The method of homogenization consisted of adding 0.2
g/l of TiO2 nanoparticles to deionised (DI) water and
then sonified for 60 min. This was the nanoparticle con-
centration for all of the tested nanoquenchants. This was
followed by mixing the nanoparticles with a defined
volume of PAG. The fluid was then sonified for
additional 30 min and mechanically stirred from time to
time. The ultrasonic bath, type BRANSONIC 220, with
the frequency of 50 kHz and power of 120 W, was used
to homogenize the nanofluids. With respect to the
water-based polymer solution, polyalkylene glycol,
known as Ucon Quenchant E, was added to water. Its
tested concentrations were volume fractions (5, 10 and
20) % of PAG, while the concentration of nanoparticles
was 0.2 g/l. The bath temperature of all liquid
quenchants was 44 °C ± 1 °C.
3 RESULTS
A set of quenching experiments was conducted with
and without agitation in order to determine the cooling
characteristics of the liquid quenchants. First, the base
fluids were tested. To see the effect of a nanoparticle
addition, nanofluids were used as liquid quenchants in
both still and agitated conditions. All of the recorded
cooling curves are shown in Figure 1. Water and water-
based nanofluids show very small differences in the
cooling characteristics. The maximum cooling rate is
within 10 % but the most significant influence is
observed during the full-film phase. The transition
temperature from the full-film phase to the nucleate
boiling phase (Tvp) is much higher for NFs and for both
J. @UPAN et al.: INVESTIGATION OF THE COOLING PROCESS WITH NANOFLUIDS ...
126 Materiali in tehnologije / Materials and technology 47 (2013) 1, 125–127
Figure 1: Cooling-curve diagrams for: a) water, b) 5 %, c) 10 % and d) 20 % volume fractions of PAG with the curves for BF 1 and NF 2 in still
conditions and the cooling curves for BF 3 and NF 4 with agitation
Slika 1: Diagrami ohlajevalnih krivulj za: a) vodo, b) 5 %, c) 10 % in d) 20 % volumenskih dele`ev PAG s krivuljami za BF 1 in NF 2 v miro-
vanju in ohlajevalne krivulje za BF 3 in NF 4 s preme{avanjem
cases of quenching with agitation. All the specific values
of the cooling curves for each quenching case are given
in Table 1. The polymer solution with 5 % PAG shows
that, in still conditions, an addition of nanoparticles
shortens the full-film stage and causes a 17 % increase in
the maximum cooling rate (CRmax).
When quenching with agitation, the cooling curves
for BFs and NFs are almost identical. The full-film phase
is more than 6 s shorter than BF and the increase in
CRmax is 34 %. Note that the CRmax of BF is higher than
in the case of NF. As the polymer concentration is
increased, the difference between the cooling rates is
getting smaller. All of the cooling curves with agitation
show that an addition of nanoparticles does not cause
better cooling characteristics. This means that agitation
is the main cause of a shorter full-film phase and a
higher CRmax. During the research it has been noted that
the effect of the deposition of nanoparticles on the
surface probe was more evident after the quenching in
still conditions.
4 CONCLUSION
The research results led to the following conclusions:
• Quenching with agitation caused a shorter full-film
phase and a higher CRmax.
• Nanofluids provided an enhancement of the
heat-transfer characteristics in still conditions for all
the cases except for the high polymer concentration.
• There is almost no difference between the cooling
curves for BFs and NFs with agitation.
• Agitation, not an addition of nanoparticles, is the
main cause for a shorter full-film phase and higher
CRmax.
• The effect of an addition of nanoparticles becomes
lower as the polymer concentration increases.
Acknowledgments
The presented investigations were achieved within
the project No. 120-1201780-1779 – Modelling of
Material Properties and Process Parameters supported by
the Ministry of Science, Education and Sports of the
Republic of Croatia.
5 REFERENCES
1 S. K. Das, S. U. S. Choi, W. Yu, T. Pradeep, Nanofluids – science
and technology, John Wiley & Sons, Inc., Hoboken, New Jersey
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fluid quenching characteristics, International Heat Treatment and
Surface Engineering, 6 (2012) 2, 56–60
J. @UPAN et al.: INVESTIGATION OF THE COOLING PROCESS WITH NANOFLUIDS ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 125–127 127
Table 1: Resulted characteristics derived from the cooling curves
Tabela 1: Zna~ilnosti, ki izhajajo iz ohlajevalnih krivulj
Quenchant CRmax/(°C/s) T(CRmax)/°C t(CRmax)/s Tcp/°C Tvp/°C
Water 187.1 577.8 6.7 127.0 786.5
Water + 0.2 g TiO2 209.5 584.0 4.9 152.6 843.4
Water with agitation 196.1 556.5 3.9 87.4 850.1
Water + 0.2 g TiO2 with agitation 182.1 496.4 5.1 72.8 849.4
Water + 5 % PAG 138.8 535.7 12.5 134.5 707.8
Water + 0.2 g TiO2 + 5 % PAG 162.5 564.7 9.2 127.5 757.1
Water + 5 % PAG with agitation 185.8 566.2 5.6 151.5 786.1
Water + 0.2 g TiO2 + 5 % PAG with agitation 175.6 579.4 5.7 154.6 780.3
Water + 10 % PAG 127.5 585.9 11.3 181.8 727.6
Water + 0.2 g TiO2 + 10 %PAG 141.8 596.2 9.1 164.3 757.8
Water + 10 %PAG with agitation 154.2 597.7 6.2 163.8 772.3
Water + 0.2 g TiO2 + 10 %PAG with agitation 158.7 595.0 6.2 162.7 772.2
Water + 20 % PAG 107.1 604.2 11.9 241.0 717.0
Water + 0.2 g TiO2 + 20 % PAG 99.6 605.0 11.2 206.9 728.2
Water + 20 % PAG with agitation 115.0 622.4 5.3 224.4 767.0
Water + 0.2 g TiO2 + 20 % PAG with agitation 115.8 631.6 6.7 482.1 761.0

M. R. I. FARUQUE et al.: ANALYSIS OF THE EFFECTS OF METAMATERIALS ON THE RADIO-FREQUENCY ...
ANALYSIS OF THE EFFECTS OF METAMATERIALS ON
THE RADIO-FREQUENCY ELECTROMAGNETIC FIELDS
IN THE HUMAN HEAD AND HAND
ANALIZA U^INKOV METAMATERIALOV NA
RADIOFREKVEN^NO ELEKTROMAGNETNO POLJE V ^LOVE[KI
GLAVI IN ROKI
Mohammad Rashed Iqbal Faruque1,2, Mohammad Tariqul Islam1,
Nik Abdullah Nik Mohamed1, Baharudin Yatim1, Mohd. Alauddin Mohd. Ali1
1Institute of Space Science (Angkasa), Faculty of Engineering and Built Environment Building, Universiti Kebangsaan Malaysia, 43600 UKM,
Bangi, Selangor, Malaysia
2Dept. of Electrical, Electronic and Systems Engineering, Faculty of Engineering and Built Environment, Universiti Kebangsaan Malaysia,
43600 UKM, Bangi, Selangor, Malaysia
rashedgen@yahoo.com
Prejem rokopisa – received: 2012-09-02; sprejem za objavo – accepted for publication: 2012-09-19
The development of split-ring resonators (SRRs) consisting of two concentric square rings of a conductive material exhibiting a
resounding electric response at a microwave frequency is described in this paper. Each square ring has a gap and is oppositely
placed in the gap of the other ring. A size reduction is always possible in this kind of structure, but one should be careful about
the set frequency. The described square type of metamaterial design is completely novel in the context of the head and hand with
respect to the specific absorption-reduction rate (SAR). This design was used because more arrays can be placed within a given
area. Mobile phones are relatively small, requiring small metamaterial designs. The finite-difference time-domain method with
the lossy-Drude model is adopted in this analysis by using CST Microwave Studio®. The technique of a SAR reduction is
discussed, and the effects of the relating position, distance, and size of the square metamaterials on the SAR reduction were
investigated. Using the square-type metamaterials, we have achieved a 53.03 % reduction of the initial SAR value for the case of
1 g SAR and a 68.08 % reduction for the case of 10 g SAR.
Keywords: antenna, head model, SRRs, SAR
V ~lanku je prikazan razvoj resonatorja iz deljivih obro~ev (SRR), ki so sestavljeni iz dveh koncentri~nih obro~ev kvadratnega
prereza iz prevodnega materiala, ki ka`e pomemben elektri~ni odziv pri mikrovalovni frekvenci. Oba kvadratna obro~a imata
re`o in vsak obro~ je nasprotno name{~en v re`i drugega obro~a. Zmanj{anje velikosti teh struktur je vedno mogo~e, vendar pa
je treba biti pozoren na na~rtovano frekvenco. Trdimo, da je kvadratna oblika metamateriala popolnoma nov dizajn v okviru
glave in roke za zmanj{anje stopnje absorpcije (SAR). Eden od pomembnih razlogov, zakaj je bila ta oblika uporabljena, je
mo`nost ve~ postavitev v danem podro~ju. Mobilni telefoni so razmeroma majhni in zahtevajo majhne naprave iz
metamaterialov. Domena kon~nih ~asovnih razlik z uporabo modela lossy-Drude je bila uporabljena v tej {tudiji uporabljajo~
CST Microwave Studia®. Prikazana je tehnika zmanj{anja SAR ter preiskovanih u~inkov prilo`enih pozicij, razdalje in velikosti
kvadratnega metamateriala na zmanj{anje SAR. S kvadratno obliko metamateriala je bilo dose`eno 53,03-odstotno zmanj{anje
za~etne vrednosti SAR pri 1 g SAR in 68,08-odstotno zmanj{anje pri 10 g SAR.
Klju~ne besede: antena, model glave, SRR, SAR
1 INTRODUCTION
Recently, there has been an increasing public concern
about the health risk caused by electromagnetic (EM)
waves emitted from cellular phones. These phones send
their signals using very small surges of high-frequency
EM waves, or microwaves, favored over most over-
the-air telecommunication systems. The fundamental
safety limits for the radio-frequency (RF) revelation are
defined in terms of the immersed power per unit mass,
which is expressed by the specific absorption rate (SAR)
in watts per kilogram (W/kg). These protection guide-
lines are set in terms of the utmost mass-normalized
rates of the electromagnetic energy deposition (SARs)
for 1 g or 10 g of tissue. The two most generally used
SAR limits today are IEEE, 1.6 W/kg for 1 g of tissue,
and ICNIRP 1–8, 2 W/kg for 10 g of tissue, disregarding
the extremities such as ankles, hands, and feet, where
higher SARs, up to 4 W/kg for any 10 g of tissue, are
allowed in both of these standards. Therefore, SAR
becomes an important performance parameter for the
marketing of mobile phones and underlines the interest
in low-SAR mobile phones by both consumers and
mobile-phone manufactures.
SAR is the parameter employed to properly quantify
the response of the biological structure in terms of
incident and induced field of the energy absorbed and
maintained inside the human body. It is the time deri-
vative (rate) of the incremental energy (dW) absorbed by
or dissipated in an incremental mass (dm) contained in a
volume element (dV) of a given density (), as seen in
(1):4
SAR
t
W
m t
W
m
= ⎛⎝
⎜ ⎞
⎠
⎟ =
⎛
⎝
⎜
⎞
⎠
⎟
d
d
d
d
d
d
d
d
(1)
Materiali in tehnologije / Materials and technology 47 (2013) 1, 129–133 129
UDK 621.37 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(1)129(2013)
The SAR is expressed in units of watts per kilogram
(W/kg) or, equivalently, in milliwatts per gram (mW/g).
It can also be related to the induced electrical field using
(2):
SAR = E2/ (2)
where E is the electric field’s root mean square (V/m), 
is the biological tissue’s electrical conductivity (S/m)
and  is the biological tissue’s density (kg/m3).
The demand on the portable mobile devices is
increasing progressively with the development of novel
wireless communication techniques. In this respect,
compact size, light weight, low profile and low cost are
now quite important challenges to be accomplished by
the designers of any wireless mobile component. One of
the most important components of any wireless system is
its radiating element. In addition to the physical require-
ments, the emerging requirements of wireless systems
are a high directivity, a large gain, and an efficient and
broadband operability of the antennas. Many broadband
techniques have been investigated to overcome the
trade-off between the antenna size and the minimum
achievable quality factor, which is dictated by Chu
formulations9. These techniques are mainly increasing
the thickness of the substrate, using differently shaped
slots or radiating patches10–14, stacking different radiating
elements or loading the antenna laterally or verti-
cally12–17, utilizing magneto-dielectric substrates18 and
engineering the ground plane as in the case of EBG
metamaterials9.
Metamaterials are artificially structured materials
providing electromagnetic properties not encountered in
nature. A left-handed material was first implemented in a
two-dimensional periodic array of split-ring resonators
and long wire strips by Smith19. The logical approach
was to excite the split-ring resonators and wire strips in
order to force the structure to behave like magnetic and
electric dipoles, respectively12.
In references20,21 it was found that the SAR value is
affected by various parameters, such as the attachment
position of conductive materials, the sizes and
configurations of the antenna ground plane. In the
reference22 studies, a reflector used for a SAR reduction
was inserted between the radiator and the head. Other
studies applied ferrite sheets2, shielding materials23,24,
resistive sheets9, and artificial magnetic conductors13 to
reduce the EM field around the antenna.
Recently, metamaterials, including electromagnetic
band-gap (EBG) structures, have been proposed for
handset antennas with low SAR characteristics.
Metamaterials combining a negative permittivity and
permeability, i.e., negative index materials (NIMs), can
be obtained by using tiny electrical circuits called split-
ring resonators and continuous wires as the metamaterial
constituent units.
The specific absorption rate (SAR) in the head can be
reduced by placing the metamaterials between the
antenna and the head. In the case of studying the SAR
reduction of an antenna operating at the GSM 900 band,
the metamaterial parameter (i.e., the permittivity is  < 0,
hence, the refractive index is n < 0) and the effective
medium would be set negative. Hence, the effectiveness
of different positions, sizes, and metamaterials with
various parameters are also analysed for SAR reduction.
2 MATERIALS AND METHODS
CST MWS is a device used as a major simulation
instrument based on the finite-integral time-domain
technique (FITD). An unvarying meshing scheme was
chosen to make the major computation devoted to
inhomogeneous mark boundaries and aimed at the fastest
and faultless results. Two-cut schemes are needed for a
complete model to show the region with the closely com-
M. R. I. FARUQUE et al.: ANALYSIS OF THE EFFECTS OF METAMATERIALS ON THE RADIO-FREQUENCY ...
130 Materiali in tehnologije / Materials and technology 47 (2013) 1, 129–133
Figure 2: Head model with a mobile phone for a SAR calculation
Slika 2: Model glave z mobilnim telefonom za izra~un SAR
Figure 1: Head model with a handset antenna
Slika 1: Model glave z ro~no anteno
pacted meshing onward to inhomogeneous boundaries.
Here the lowest and highest mesh sizes were 0.3 mm and
1.0 mm, respectively. As a result, a total of 2 122 164
mesh cells were generated and the simulation time was
1208 s, including the mesh generation for each effort on
an Intel Core TM 2 Duo E 8400 3.0 GHz CPU with a 4
GB RAM system, for a complete model. It is noted here
that at first the materials are placed between the antenna
and the human head, and then all of these are replaced by
a metamaterial.
For this research, the SAM head model was con-
sidered. It consists of about 2 097 152 cubical cells at a
1 mm resolution. Figure 1 represents a portable tele-
phone model with a handset antenna considered in the
SAR calculation.
The head models used in this study were obtained
from an MRI-based head model through the whole-brain
Atlas website. Six types of tissues, i.e., bone, brain,
muscle, eye ball, fat, and skin were involved in this
model9,19. A space domain enclosing the human head and
the phone model is also shown in Figure 2. Tables 1 and
2 show their dielectric properties for 900 & 1800 MHz.
Numerical simulations of the SAR value were performed
by the FDTD method. The parameters for the FDTD
computation were as follows. The simulation domain
was 128 × 128 × 128 cells. The cell sizes were set as x
= y = z = 1 mm. The computational domain was
terminated with 8 cells of a perfectly matched layer
(PML). A PIFA antenna was modeled with a thin-wire
approximation.
Table 1: Dielectric tissue properties at 900 MHz
Tabela 1: Dielektri~ne lastnosti tkiva pri 900 MHz
Material Density/(kg m–3)
Conductivity
/(S m–1)
Relative
permittivity, r
Fat, bone 1130 0.12 4.83
Muscle, skin 1020 1.5 50.5
Brain 1050 1.11 41.7
Eye ball 1000 2.03 68.6
Table 2: Dielectric tissue properties at 1800 MHz
Tabela 2: Dielektri~ne lastnosti tkiva pri 1800 MHz
Material Density/(kg m–3)
Conductivity
/(S m–1)
Relative
permittivity, r
Fat, bone 1130 0.11 4.48
Muscle, skin 1020 1.35 47.80
Brain 1050 1.09 39.50
Eye ball 1000 1.99 65.3
3 SRRS STRUCTURE AND DESIGN
With this work we established that, using the FDTD
analysis, square metamaterials (SMMs) can reduce the
peak SAR for 1 g and SAR for 10 g in the head. In this
section, the SMMs are evaluated in a cellular phone with
the 900 MHz and 1800 MHz bands. Periodically
arranged square split-ring resonators (SSRRs) can work
as SMMs. The SSRR structure involves two conductive-
material, concentric square rings. Both square rings have
a gap, and each ring is placed in the gap of the other ring.
The schematic of the SSRR structure used in this study is
shown in Figure 3.
To construct the SMMs for a SAR reduction, SSRRs
were used as the resonator model, as shown in Figure 3.
The resonators operated in the 900 MHz bands. The
SSRRs contain two square rings, each with a gap on the
opposite side5. The SSRRs were introduced by Pendry et
al.11 (1999), and subsequently used by Smith et al.
(2000) to synthesize the first left-handed artificial me-
dium19.
Figure 4 shows the fabricated SMM arrays used in
this measurement. The metamaterials in this work were
designed with periodic SSRR arrangements to reduce the
M. R. I. FARUQUE et al.: ANALYSIS OF THE EFFECTS OF METAMATERIALS ON THE RADIO-FREQUENCY ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 129–133 131
Figure 3: Square Metamaterial arrays used in this calculation
Slika 3: Kvadratna razporeditev metamateriala, uporabljena pri
izra~unu
Figure 4: Fabricated square metamaterial arrays used in this research
Slika 4: Izdelana kvadratna razporeditev metamateriala, uporabljena v
tej raziskavi
SAR value. By properly designing the SSRR structure
parameters, a negative effective medium parameter can
be achieved for the 900 Hz and 1800 Hz bands.
4 NUMERICAL RESULTS AND DISCUSSION
The designed SSRRs were placed between the
antenna and the human head, thus reducing the SAR
value. To study the SAR reduction of the antenna, oper-
ating at the GSM 900 band, different positions, sizes,
and metamaterials for the SAR-reduction effectiveness
are also analyzed by using the FDTD method in
conjunction with a detailed human-head model. The
antenna was put equidistant to the head axis. The
distances from 5 mm to 20 mm were checked and ulti-
mately the distance of 20 mm was selected. Moreover,
the maximum SAR reduction was achieved when the
proposed structure was placed to the near side of the
cellular phone. In the present paper, the power output of
the mobile was considered to be 600 mW at the opera-
tional frequency of 900 MHz. In the case when the main
network station is far from the mobile phone, the actual
power is 1 W to 2 W. The highest values were 2.002
W/kg for SAR 1 g and 1.293 W/kg for SAR 10 g,
calculated when the cellular phone model was at a
distance of 20 mm from the human-head model and
without any metamaterial attachment. Obviously, this
SAR reduction is more desirable than the metamaterial
free attachment. The results quoted in12,7 are 2.28 W/kg
and 2.17 W/kg for SAR 1 g. This is due to the use of
different antennae placed in different locations relative to
the head model. These SAR values agree with the
compared result reported in5, i.e., 2.43 W/kg for SAR
1 g. This was achieved using different radiating powers
and different antenna.
To study the effect of a SAR reduction with the use
of SMMs, the radiated power from the PIFA antenna
with μ = 1 and  = –3 was fixed at 600 mW. Calculating
SAR at 900 MHz, we obtained the value of 1.673 W/kg
for SAR 10 g without a metamaterial; however, in the
case of a metamaterial, the reduced SAR value for 10 g
was 0.737 W/kg. Thus, a reduction of 55.95 % was
achieved, whereas the sketch furnish design in5 acquired
a 32.03 % reduction. The reasons for this difference are
different densities, antennae, sizes of the metamaterial
and types of conductivity. The result implies that only
suppressing the maximum current on the front side of the
conducting box contributes significantly to the reduction
of the spatial-peak SAR, as the decreased quantity of the
power absorbed in the head is considerably larger than
that dissipated in the metamaterials. It is clear from the
simulation results that metamaterials can reduce the peak
SAR successfully where the antenna reduction is merely
pretended. So, both the inner capacitance and inductance
of the metamaterials are reinforced. At this stage, the
mediums will exhibit stop band walks with a single
negative medium parameter.
5 EXPERIMENTAL EVALUATIONS
The SAR measurement was performed using the
COMOSAR measurement system. The system uses a
robot to position the SAR probe inside the head phan-
tom. The head phantom is filled with a liquid that has
dielectric properties selected on the basis of the IEEE
standard 1528, which are r = 41.5 and  = 0.97 S/m for
900 MHz and r = 40 and  = 1.4 S/m for 1800 MHz.
The measured and simulated SAR values without the
inclusion of TMMs are shown in Table 3.
Table 3: SAR simulation and measurement results without the
inclusion of TMMs (distance between the head and the phone model,
d = 20 mm)
Tabela 3: Simulacije SAR in rezultati meritev brez upo{tevanja
TMMs (razdalja med glavo in modelom telefona, d = 20 mm)
Value of SAR
SAR 1 g SAR 10 g
Simulated 2.002 1.673
Measured 1.936 1.609
Table 3 shows that the simulated SAR value is
greater than 3.29 % for SAR 1 g and 3.82 % for SAR 10
g due to the fact that the distance between the head and
phone model was not correctly set at the measurement
stage. In addition, the distance between the source and
the internal surface of the phantom affects the SAR. For
a 5-mm distance, a positioning uncertainty of ±0.5 mm
would produce a SAR uncertainty of ±20 %. An accurate
device positioning is therefore essential for accurate
SAR measurements.
The simulated and measured SARs are obtained
using the tilted position of the SMMs with the antenna
revealing the simulated and measured SAR values of
1.1673 W/kg and 1.0623 W/kg for SAR 1 g, respec-
tively. The simulated and measured SARs differed by
6.27% for SAR 1 g. Regarding the difference in the
absolute values of the peak SAR, the phone-model
casing for the simulation was different from the casing
used for the measurement. In addition, Scotch tape was
used to attach the SMMs and antenna at the measure-
ment stages. The simulated and measured results also
differ because the parameters for the measurement
system change with water evaporation and temperature.
Furthermore, the measurement system contains several
parameters (source, network emulator, probe, electronic
evaluation procedures) that affect the SAR calculation
but are not included in the simulation.
6 CONCLUSION
The process of the EM absorption between an
antenna and the human head, while using new SMMs,
has been thoroughly and intensively experimented and
presented in this paper. After examining the substance of
the developed SMMs for the phone model, the SAR
values were found to be about 0.639 W/kg for SAR 10 g
M. R. I. FARUQUE et al.: ANALYSIS OF THE EFFECTS OF METAMATERIALS ON THE RADIO-FREQUENCY ...
132 Materiali in tehnologije / Materials and technology 47 (2013) 1, 129–133
and 1.0623 W/kg for SAR 1 g. These results will be the
core points when providing information in the field of
designing communication equipment, safe from interrup-
tions.
7 REFERENCES
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M. R. I. FARUQUE et al.: ANALYSIS OF THE EFFECTS OF METAMATERIALS ON THE RADIO-FREQUENCY ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 129–133 133

K. MICHALEK et al.: POSSIBILITIES FOR DESULPHURIZATION OF AN ALLOY STEEL ...
POSSIBILITIES FOR DESULPHURIZATION OF AN
ALLOY STEEL IN A VOD DEVICE WHILE USING
CHEMICAL HEATING
MO@NOST RAZ@VEPLJANJA LEGIRANEGA JEKLA V
VOD-NAPRAVI Z IZKORI[^ANJEM KEMIJSKEGA OGREVANJA
Karel Michalek1, Libor ^amek1, Karel Gryc1, Tomá{ Huczala2,
Vladimír Troszok2
1V[B-Technical University of Ostrava, FMME, Department of Metallurgy and Foundry, 17. listopadu 15/2172, 708 33 Ostrava, Czech
Republic
2Tøinecké `elezárny, a. s., Prùmyslová 1000, 739 70 Tøinec-Staré Mìsto, Czech Republic
karel.michalek@vsb.cz
Prejem rokopisa – received: 2012-09-03; sprejem za objavo – accepted for publication: 2012-09-18
The paper describes the knowledge about and results of the experimental heats performed in an electric steel plant. The aim was
to verify the possibilities of a controlled desulphurization of alloy steel in a VOD (Vacuum Oxygen Decarburization) device,
particularly when using chemical heating through the OVD process (Oxygen Vacuum Deoxidation/Degassing), as well as the
standard slag formers. Experimental procedures were used in the production of alloy tool steel. Both, the system of making the
reducing slag and optimizing the composition of individual oxides in order to achieve the desired basicity, are complicated by
the products of the chemical heating. Therefore, it was necessary to develop a new production technology that would eliminate
such oxide products of chemical heating and, thus, the maximum degree of desulphurization could be achieved. The main
principle of the technology is to create appropriate thermodynamic and kinetic conditions required for desulphurization. In
particular, this concerns a low oxygen activity in the steel, the composition of refining slag, and an intense bath stirring with
argon.
Keywords: steel desulphurization, VOD device, OVD process, slag, steel cleanliness
^lanek opisuje znanje in rezultate raziskave eksperimentalnih talin, izdelanih v elektrojeklarni. Namen je bil preveriti mo`nost
kontroliranega raz`vepljanja legiranega jekla v VOD (Vacuum Oxygen Decarburization)-napravi, posebno med uporabo
kemijskega ogrevanja med OVD (Oxygen Vacuum Deoxidation/Degassing)-postopkom, kot tudi standardnih gradnikov `lindre.
Eksperimentalni postopki so bili uporabljeni pri proizvodnji legiranega orodnega jekla. Oboje, tvorba reduktivne `lindre in
optimalna sestava posameznih vklju~kov, z namenom doseganja `elene bazi~nosti, ovirajo produkti kemijskega ogrevanja. Zato
je bilo potrebno razviti novo tehnologijo proizvodnje, ki bi prepre~ila oksidne produkte kemijskega ogrevanja, kar bi omogo~ilo
doseganje maksimalne stopnje raz`vepljanja. Osnovni princip nove tehnologije je zagotoviti ustrezne termodinami~ne in
kineti~ne pogoje, potrebne za raz`vepljanje. To pomeni nizko aktivnost kisika v jeklu, sestavo rafinacijske `lindre in intenzivno
me{anje kopeli z argonom.
Klju~ne besede: raz`vepljanje jekla, VOD-naprava, OVD-postopek, `lindra, ~istost jekla
1 INTRODUCTION
The technology of the production of alloy tool steels
require a special preparation of the input materials,
followed by a strict compliance with technological and
metallurgical processes when producing steel in a 10-ton
electric arc furnace (EAF), as well as the ladle’s optimal
preparation. The final steel processing in a VOD device
requires a continuous and systematic management of
individual production stages, in particular the chemical
heating, creating a high vacuum, production of reduction
and refining slag1,2, additional steel alloying, and conti-
nuous, optimal purification with argon.
The required steel desulphurization in the VOD
device can only be achieved by mastering the technology
and metallurgy of the processes, especially by opti-
mizing the slag regime and complying with the basic
thermodynamic and kinetic conditions of a desulphuri-
zation reaction.
2 PRINCIPAL FACTORS AFFECTING THE
REQUIRED DEGREE OF STEEL
DESULPHURIZATION
In the paper formerly presented3, the technology of
the low-sulphur-content steel production (up to the mass
fraction w = 0.002 %) in the conditions of EAF using
CaO and Refraflux additives was described in detail.
With regard to the product range of tool steels,
desulphurization technology in the VOD device with an
oxidation stage at the beginning of the process, used for
the chemical heating of steel, has been assessed and
gradually developed.
The desired degree of desulphurization in a vacuum
station when applying chemical heating is largely limited
by the output sulphur content from EAF. In the product-
ion of alloy steel grades in VOD, the chemical-heating
method creates entirely different starting conditions for
the formation of the reducing slag compared to the
standard technological processes in EAF. Nevertheless,
Materiali in tehnologije / Materials and technology 47 (2013) 1, 135–140 135
UDK 669-982:669.15 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(1)135(2013)
the reduction conditions ensuring a successful desulphu-
rization of steel can be created using special procedures
in this relatively harsh environment.
Standard patterns of the desulphurisation of steel and
their review are listed in our earlier works3,4. The
principles that must be met in order to obtain low
sulphur content in steel can be summarized into the
following:
• A high level of activity of free-oxygen anions in the
slag, i.e., a high basicity of the slag with a high
proportion of alkaline oxides and a low proportion of
acidic oxides.
• A low activity of oxygen ao in steel, i.e., a low con-
tent of dissolved oxygen and a low value of the
activity coefficient fo.
A negative factor affecting the degree of steel desul-
phurization is the presence of the "easily reducible"
oxides in the refining slag – besides FeO there are also
MnO, P2O5 and Cr2O3. The sum percentage content of
the aforementioned elements for a well-working refining
slag is usually recommended to be up to the mass
fraction w = 3 %.
From the kinetic point of view, an increased tempe-
rature positively affects steel desulphurization. An
increase in the temperature helps us decrease the
viscosity of the slag and metal, increase the sulphur
diffusion coefficient, and allows a reduction in the
surface tension, which results in the chemical reaction
more quickly reaching the state of equilibrium. However,
the effect of appropriate kinetics of the ongoing pro-
cesses is closely connected with meeting the basic
thermodynamic conditions.
Creating appropriate conditions for desulphurization
in the steel processing in VOD upon the use of chemical
heating requires a specific technology. A negative factor
after the chemical heating in VOD is a gradual decrease
in the temperature. In a formed temperature interval, and
in a gradual cooling of the steel, it is necessary to carry
out all the necessary technology operations to create
suitable thermodynamic and kinetic conditions for the
desired degree of desulphurization.
3 CURRENT TECHNOLOGICAL AND
METALLURGICAL PROCEDURES OF
FORMING A REDUCING SLAG
The original technology of the alloy-steel production
with a low sulphur content that is below 0.005 % was
based on the standard charge of recyclable waste. Thus,
it was ensured that the critical sulphur content would not
be exceeded during the processing in EAF, and that the
desired sulphur content of below w = 0.005 % would be
reached in the VOD device using the chemical heating
(based on Al + O2).
For the actual production process the calcium slag
containing calcium fluorite (CaF2) was used in the
electric steel plant of TØINECKÉ @ELEZÁRNY, a. s.
(T@, a. s.). The positive effect of CaF2 supporting the
thermodynamics and also the kinetics of the required
processes is widely known and has already been
discussed in our earlier work3. The disadvantages of
using fluorite are the reduction of the lifetime of the
lining of furnaces and ladles, and its adverse impact on
the working environment (formation of fluorites).
Currently, in the production of alloy steels in EAF the
fluorite is being replaced with a high-quality, industrially
produced synthetic slag. The chemical composition of
the tested synthetic slag is shown in the ternary diagram5
in Figure 1. The slag-formation reduction technology
with a high refining effect is based on the application of
a mixture of quicklime and synthetic slag Refraflux,
where both components are transported using alternating
dosing to the surface of the bath in EAF.
The original slag additive – fluorite – stays in use to
create the reducing slag in the secondary-metallurgy
processing VOD device. The main reason is the need for
the use of chemical heating with aluminium, where the
product of this chemical reaction – Al2O3 – negatively
affects the system of chemical optimization of the
reducing-slag composition, as well as the secondary
products of oxidation of FeO, MnO and Cr2O3. The
standard requirement for the range of chemical heating
in the production of alloy-steel grades is 130–180 °C,
which leads to a significant increase in the content of
Al2O3 in the slag after oxidation.
4 PRODUCTION CONDITIONS AND
PROCEDURES FOR EXPERIMENTAL HEATS
Production processes for desulphurization have been
tested on VOD using the technology of chemical heating
K. MICHALEK et al.: POSSIBILITIES FOR DESULPHURIZATION OF AN ALLOY STEEL ...
136 Materiali in tehnologije / Materials and technology 47 (2013) 1, 135–140
Figure 1: Area of the chemical composition of the tested synthetic
slag REFRAFLUX 4842 S in the ternary diagram Al2O3-CaO-SiO25
Slika 1: Podro~je kemijske sestave preizku{ene sinteti~ne `lindre
REFRAFLUX 4842 S v ternarnem diagramu Al2O3-CaO-SiO25
(OVD) in the production of the alloy tool-steel grade
19569 (X63CrMoV5.1). Table 1 shows the internal
release chemical composition of the steel grade 19569.
Table 1: Internal release chemical composition of 19569 steel in mass
fractions, w/%
Tabela 1: Interna kemijska sestava jekla 19569 v masnih dele`ih, w/%
Chemical composition 19569 (X63CrMoV5.1)w/%
C 0.58–0.68
Si 0.7–1.1
Mn 0.25–0.55
P max. 0.015
S max. 0.010
Cr 4.5–5.5
Mo 0.8–1.2
V 0.2–0.4
W max. 0.6
The output from EAF is the sulphur content of w =
0.005 % achieved as a standard, which can be, during the
subsequent processing with the VOD technology using
chemical heating, maintained at the achieved level or
reduced below the given limit.
The technological process involved the following
operations:
• Dislodging the reducing slag from the ladle after
tapping EAF. This part of the technology is, from the
perspective of preventing the reverse transition of
sulphur from the slag into the metal and the
possibility of creating a new reducing slag, very
favourable. However, it also negatively affects the
heat loss through the uncovered steel surface, which
increases the need for the strength of the chemical
heating.
• A creation of a new cover slag on the steel surface.
This was carried out by applying the basic charge of
50 kg lime and by adding again the same charge of
lime in addition to an approximately 20 kg of fluorite
upon the installation of the casting ladle (CL) into the
VOD device.
• An addition of aluminium and oxygen-blowing in
one or two stages combined with chemical heating, in
this way the temperature was increased by 150–200
°C above the liquidus temperature of the alloy steel
grade produced.
• A gradual formation of a non-foaming reducing slag
caused by the lime additive.
• Steel evacuation to perform a vacuum deoxidation
and to create conditions for increasing the cleanliness
of the steel including its precise additional alloying.
• Temperature and chemical homogenization with
argon at atmospheric pressure, whose duration was
set by the temperature margin of the gradual, con-
trolled cooling, within which the treatment of the
chemical composition of the reducing slag was
carried out, or small corrections to the chemical com-
position of the alloy tool steel were made.
• After reaching the required steel temperature above
the liquidus, the processing in VOD was completed
and the casting into the moulds was done.
In the process of gradual cooling from the tempe-
ratures after the chemical heating, it was difficult to
approach the optimum values of the chemical com-
position of the reducing slag, which should contain mass
fractions from 45 % to 55 % CaO, 18 % to 25 % Al2O3,
" 10 % SiO2. At the same time and in accordance with
the theory of steel desulphurization, it was also very
important to maintain a low oxygen activity in the metal.
This requirement has been, in terms of the VOD device,
met particularly by maintaining the maximum possible
content of aluminium in the liquid metal. Suitable kinetic
conditions were continuously provided with the optimum
control – blowing argon through the bottom porous block
with the volumetric flow rates of up to 20 l min–1.
5 ACHIEVED RESULTS
The course of each experimental heat was, in the
secondary metallurgy VOD with chemical heating, to
some extent, unique. A total of 10 heats were produced.
The aim was to verify the options of the controlled steel
desulphurization in the VOD device using chemical
heating (OVD) as well as the standard slag-forming
additives. The comparison of the basic chemical analyses
of the slag and steel of the selected 19569 heat grade
tested is shown in Table 2.
The achieved results describing the composition of
the final reducing slag after the steel processing show
some differences, especially the differences for Al2O3
and SiO2 oxides. These variations result from individual
processing procedures in the VOD device in relation to
the method of how to control chemical heating. In most
of the heats the CaO/Al2O3 ratio ranged from 1.0 to 2.63,
and the basicity (CaO/SiO2) was in the range from 2.15
to 10.
K. MICHALEK et al.: POSSIBILITIES FOR DESULPHURIZATION OF AN ALLOY STEEL ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 135–140 137
Table 2: Selected chemical analyses of the final reducing slag and steel grade 19569 in mass fractions, w/%
Tabela 2: Izbrane kemijske analize kon~ne redukcijske `lindre jekla 19569 v masnih dele`ih, w/%
Heat
No. CaO Al2O3 MgO SiO2 FeO Cr2O3 MnO Ssteel
Temperature before
chemical heating
Temperature after
chemical heating
4 35.49 35.82 2.74 13.22 0.35 0.19 0.17 0.0030 1559 °C 1682 °C
9 40.62 26.38 3.46 18.88 1.03 0.66 0.74 0.0030 1529 °C 1685 °C
6 41.00 31.70 5.20 13.90 0.19 0.32 0.45 0.0030 1544 °C 1723 °C
7 48.37 18.40 4.40 4.80 0.28 0.09 0.01 0.0020 1575 °C 1734 °C
In terms of a successful desulphurization it was very
important to maintain the parameters of the input
chemical analyses and to optimize the steel temperatures
before and after the chemical heating in the VOD device
so that the composition of the arising reducing slag was
not subsequently tasked, in particular by a high content
of aluminium oxide.
During the subsequent processing following the
chemical heating it was necessary to continue the adding
of lime and, based on the visual inspection results, to
substitute the additive for its liquefaction. Based on the
information obtained from the chemical analysis, it was
recommended to only alloy the materials that do not
allow batching the VOD device during the next phase of
vacuuming.
The changes in the chemical compositions of the
steel and the slag during the processing in the secon-
dary-metallurgy VOD device are shown in Figures 2 to
5.
In terms of reducing the slag formation, any deviation
from the recommended technology causes deflection
from its optimum composition, which prevents the alloy
steel from achieving the required degree of desulphuri-
zation.
The processing of the No. 4 heat can be shown as an
example (Figure 2). A low content of CaO in the
reducing slag did not ensure a further desulphurization of
the steel, so that the resulting sulphur content was
identical to the initial content (30 μg/g).
A similar course of processing can be noted in the
cast No. 9 (Figure 3), which resulted in a relatively rapid
drop in the temperature (due to the insufficient heat
capacity of the ladle) following the completion of the
vacuuming phase, leading to an unsatisfactory content of
CaO in the reducing slag due to the worsened assimi-
lation of lime. The final sulphur content of the heat was
identical to the initial content (30 μg/g).
From the course of processing the heat No. 6 (Figure
4), where the input sulphur content was 40 μg/g – it is
clear that after the completion of the vacuuming phase
the sulphur content was increased to 70 μg/g. The result
of the repeated measurement of the oxygen activity in
the steel showed a higher value (7.3 μg/g), which was
probably associated with a low total aluminium content
in the steel (w = 0.001 %). The subsequent additional
deoxidation by aluminium in the phase of the controlled
additional cooling decreased this activity (to 1.7 μg/g),
but, on the other hand, it shifted the optimum
composition of the reducing slag (too high a content of
Al2O3). Through specified treatments we managed to
keep the resulting sulphur content at the value of
30 μg/g.
The process of operating the experimental heat No. 7,
whose course is shown in Figure 5, has been found to be
K. MICHALEK et al.: POSSIBILITIES FOR DESULPHURIZATION OF AN ALLOY STEEL ...
138 Materiali in tehnologije / Materials and technology 47 (2013) 1, 135–140
Figure 3: Development of the chemical composition of the steel and
the slag during the processing in the secondary-metallurgy VOD
device with chemical heating in the heat No. 9
Slika 3: Razvoj kemijske sestave jekla in `lindre med procesom
sekundarne metalurgije v VOD-napravi s kemijskim segrevanjem
taline {t. 9
Figure 2: Development of the chemical composition of the steel and
the slag during the secondary-metallurgy processing in the VOD
device with chemical heating in the heat No. 4
Slika 2: Razvoj kemijske sestave jekla in `lindre med procesom se-
kundarne metalurgije v VOD-napravi s kemijskim segrevanjem taline
{t. 4
the optimal course of the steel processing in the
secondary-metallurgy VOD device with chemical
heating. The initial sulphur content in the steel of 50
μg/g was, after meeting the recommended procedures for
the steel-processing technology, reduced to the resulting
content of 20 μg/g. What contributed to this result were
not only a suitable composition of the reducing slag
(CaO > 50 %, B > 5, CaO/Al2O3 > 2, FeO < 0.3 %), and
a low oxygen activity in the steel (1.4 μg/g), but also the
stirring of the steel.
6 IMPACT OF TECHNOLOGY ON
CLEANLINESS OF STEEL
The carried-out experimental heats were assessed in
terms of the achieved cleanliness of the steel. An
electron microanalyzer Hitachi S-3500N was used for
the evaluation, on the basis of which we determined the
framework compositions of non-metallic inclusions,
their numbers and geometrical proportions. The achieved
results for the heats No. 6 and No. 7 are shown as the
ternary diagrams in Figure 6. As for these heats, the
lime additive was, in the final phase of the controlled
additional cooling, added with no fluorite additive. This
led to a cooling of the reducing slag, increasing its
viscosity, and a worsening of the kinetic conditions for
K. MICHALEK et al.: POSSIBILITIES FOR DESULPHURIZATION OF AN ALLOY STEEL ...
Materiali in tehnologije / Materials and technology 47 (2013) 1, 135–140 139
Figure 6: Ternary diagrams as a result of the chemical microanalysis
of the inclusions from the heats No. 6 and No. 7
Slika 6: Ternarni diagrami kot rezultat kemijske mikroanalize
vklju~kov iz talin {t. 6 in {t. 7
Figure 4: Development of the chemical composition of the steel and
the slag during the processing in the secondary-metallurgy VOD
device with chemical heating in the heat No. 6
Slika 4: Razvoj kemijske sestave jekla in `lindre med procesom
sekundarne metalurgije v VOD-napravi s kemijskim segrevanjem
taline {t. 6
Figure 5: Development of the chemical composition of the steel and
the slag processing in the secondary-metallurgy VOD-device with
chemical heating in the heat No. 7
Slika 5: Razvoj kemijske sestave jekla in `lindre med procesom
sekundarne metalurgije v VOD-napravi s kemijskim segrevanjem
taline {t. 7
the absorption of non-metallic inclusions, which also led
to a higher incidence of the inclusions based on
CaO-Al2O3. The micro-purity results for the other expe-
rimental heats were standard.
7 CONCLUSION
The production technology of the controlled steel
desulphurization with a requirement to keep the target
sulphur content below 0.0050 % was optimized in the
VOD device using chemical heating of the steel.
Technologies have been applied in the production of the
alloy tool-steel grade No. 19569 (X63CrMoV5.1). The
main issue in the design of the technology was that the
systems for the production of the reducing slag and
optimized compositions of individual oxides have been
disrupted by the chemical-heating products.
The essential problems are the amount of Al2O3,
which is formed after the chemical heating during the
oxidation of added aluminium by gaseous oxygen and,
simultaneously, the formed quantity of the easily
reducible oxides MnO, FeO and Cr2O3. With regard to
this, it is necessary to add high quantities of CaO to
achieve an appropriate content of free CaO and an
increase in the CaO/Al2O3 ratio. However, this quantity
results in a noticeable drop in the temperature during the
heat completion. The reduction of oxides is completed
after the vacuuming.
In accordance with the theory of steel desulphu-
rization, the efficiency of maintaining a low oxygen
activity in the metal where the FeO content in the slag is
below 1 % was confirmed. This requirement has been
ensured in the VOD device after the chemical heating
primarily by maintaining the maximum possible
aluminium content in the liquid metal. It is also
necessary to maintain the appropriate slag basicity (over
5) and the CaO/Al2O3 ratio, which should be in the range
of 1.8 to 2.2.
Results of the microanalysis of the inclusions showed
that, to achieve the desired cleanliness of the tested steel
grade, it is required, after the completed vacuuming, to
run the final phase of processing the reducing slag only
with the optimum steel temperature, weight and
chemical composition of the slag. It was found that an
increase in the CaO additive changes the chemical
compositions of the inclusions, in particular causing an
increase in the CaO content. The impact of the change in
the chemical compositions of the inclusions on the
beneficial properties of the steel will be monitored with a
final customer.
Acknowledgements
This work was supported by the Ministry of Industry
and Trade of the Czech Republic in the framework of the
projects No. FR-TI3/374 and FR-TI3/373.
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140 Materiali in tehnologije / Materials and technology 47 (2013) 1, 135–140
PROMOCIJA [TUDIJA IN RAZISKAV MATERIALOV
Slovensko dru{tvo za materiale, Lepi pot 11, 1000
Ljubljana, je 6. 12. 2012 `e deseto leto zapored pri-
pravilo sklop predavanj v okviru Promocije {tudija in
raziskav materialov. Prireditev je bila v predavalnici P5
Naravoslovnotehni{ke fakultete, Oddelek za materiale in
metalurgijo, A{ker~eva 12 v Ljubljani.
Na prireditvi je 7 predavateljev predstavilo dose`ke s
podro~ja kovinskih in polimernih materialov ter trdih
prevlek, prisotnih pa je bilo 122 dijakov in njihovih
spremljevalcev. Ker je bil eden od dijakov gluh, ga je
spremljal prevajalec za gluhoneme.
Najprej je predstavnica NTF OMM, dr. Maja Von-
~ina, predstavila {tudij materialov in metalurgije na
Univerzi v Ljubljani, Naravoslovnotehni{ki fakulteti,
Oddelek za materiale in metalurgijo, ki je organiziran po
bolonjskih na~elih in obsega Univerzitetni {tudijski pro-
gram 1. stopnje – In`enirstvo materialov ter Visoko{olski
{tudijski program 1. stopnje – Metalur{ke tehnologije.
Nova oblika organiziranja {tudija omogo~a tudi vklju~e-
vanje {tudentov v raziskovalno delo `e takoj na za~etku
{tudija.
V naslednjih predavanjih so bili predstavljeni rezul-
tati raziskav in smeri razvoja na podro~ju sodobnih
kovinskih materialov, trdih prevlek in polimernih mate-
rialov.
Prisotni dijaki so menili, da je taka oblika predavanj
koristna, saj se tako lahko seznanijo z mo`nostmi za {tu-
dij materialov in bolje spoznajo raziskovalno dejavnost
ter delo na podro~ju materialov.
Dokaz, da obstaja med mladimi zanimanje za mate-
riale, je leto{nja prireditev, ki je bila `e deseta zapo-
vrstjo. Dijaki in njihovi spremljevalci so izrazili `eljo, da
bi Slovensko dru{tvo za materiale tako prireditev {e
naprej pripravilo vsako leto.
Promocija {tudija in raziskav materialov je name-
njena dijakom, da pridobijo nekaj koristnih informacij s
podro~ja {tudija in raziskav materialov. Za mlaj{e
predavatelje pa je prireditev prilo`nost, da si pridobijo
potrebne izku{nje za nastopanje v javnosti.
doc. dr. Matja` Torkar
tajnik SDM
Slovensko dru{tvo za materiale
Lepi pot 11
1000 Ljubljana
Materiali in tehnologije / Materials and technology 47 (2013) 1, 141 141
Dijaki – udele`enci Promocije {tudija in raziskav materialov v preda-
valnici P5. Predavanja je prevajal tudi tolma~ za gluhoneme.
Predstavnica NTF OMM, dr. Maja Von~ina predstavlja mo`nosti {tu-
dija materialov
UDK 669:378(497.4) ISSN 1580-2949
Dogodki/Events MTAEC9, 47(1)141(2013)