VSEBINA – CONTENTS
Predgovor/Foreword
M. Torkar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95
IZVIRNI ZNANSTVENI ^LANKI – ORIGINAL SCIENTIFIC ARTICLES
Linear two-scale model for determining the mechanical properties of a textile composite material
Linearni dvostopenjski model za dolo~itev mehanskih lastnosti tekstilnega kompozita
T. Kroupa, P. Janda, R. Zem~ík . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 97
Influence of the process parameters and the mechanical properties of aluminum alloys on the burr height and the surface
roughness in dry drilling
Vpliv parametrov procesa in mehanskih lastnosti aluminijevih zlitin na vi{ino igle in hrapavost povr{ine pri suhem vrtanju
U. Köklü. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103
The performance of various artificial neurons interconnections in the modelling and experimental manufacturing of the
composites
Predstavitev razli~nih umetnih nevronskih povezav pri modeliranju in eksperimentalni izdelavi kompozitov
M. O. Shabani, A. Mazahery . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109
Experimental and theoretical investigation of drying technology and heat transfer on the contact cylindrical dryer
Eksperimentalna in teoreti~na raziskava tehnologije su{enja in prevajanja toplote na kontaktnem valjastem su{ilniku
S. Prvulovi}, D. Tolma~, M. Lambi}, D. Dimitrijevi}, J. Tolma~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 115
An RE/RM approach to the design and manufacture of removable partial dentures with a biocompatibility analysis of the F75
Co-Cr SLM alloy
RE/RM-pribli`ek, na~rtovanje in izdelava snemljivih delov zobovja z analizo biokompatibilnosti zlitine F75 Co-Cr SLM
D. P. Jevremovi}, T. M. Pu{kar, I. Budak, Dj. Vukeli}, V. Koji}, D. Eggbeer, R. J. Williams. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 123
Influence of transient response of platinum electrode on neural signals during stimulation of isolated swinish left vagus nerve
Vpliv prehodnega zna~aja platinaste elektrode na `iv~ni signal med stimulacijo izoliranega `ivca vagusa svinje
P. Pe~lin, F. Vode, A. Mehle, I. Gre{ovnik, J. Rozman . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 131
Effect of the antimony thin-film deposition sequence on copper-silicon interdiffusion
Vpliv zaporedja nanosa tankih plasti antimona na interdifuzijo baker-silicij
M. Nasser, B. Mokhtar, B. Mahfoud, R. Mounir, Z. Fouzia, B. Chaouki . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 139
Lime-metakaolin hydration products: a microscopy analysis
Produkti hidracije apno-metakaolin: mikroskopska analiza
A. L. Gameiro, A. Santos Silva, M. R. Veiga, A. L. Velosa . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 145
The impact of die angle on tool loading in the process of cold extruding steel
Vpliv kota matrice na obremenitev orodja pri hladni ekstruziji jekla
S. Randjelovi}, M. Mani}, M. Trajanovi}, M. Milutinovi}, D. Movrin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 149
Final-structure prediction of continuously cast billets
Napoved kon~ne mikrostrukture kontinuirno ulitih gredic
J. [tìtina, L. Klime{, T. Mauder, F. Kavi~ka . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 155
Outgassing of hydrogen from a stainless steel vacuum chamber
Razplinjevanje vodika iz nerjavnega jekla
S. Avdiaj, B. Erjavec . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 161
STROKOVNI ^LANKI – PROFESSIONAL ARTICLES
The quality of super-clean steels produced at @ÏAS, inc.
Kakovost super~istih jekel, izdelanih v podjetju @ÏAS, inc.
M. Balcar, L. Martínek, P. Fila, J. Novák, J. Ba`an, L. Socha, D. A. Skobir Balanti~, M. Godec. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 169
ISSN 1580-2949
UDK 669+666+678+53 MTAEC9, 46(2)93–190(2012)
MATER.
TEHNOL.
LETNIK
VOLUME 46
[TEV.
NO. 2
STR.
P. 93–190
LJUBLJANA
SLOVENIJA
MAR.–APR.
2012
Simulations of the shrinkage porosity of Al-Si-Cu automotive components
Modeliranje kr~ilne poroznosti Al-Si-Cu avtomobilskih ulitkov
L. Lavtar, M. Petri~, J. Medved, B. Taljat, P. Mrvar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 177
Wear-resistant intermetallic arc spray coatings
Obrabna obstojnost intermetalnih prevlek, napr{enih v elektri~nem obloku
E. Altuncu, S. Iriç, F. Ustel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 181
Effect of sintering parameters on the density, microstructure and mechanical properties of the niobium-modified
heat-resistant stainless steel GX40CrNiSi25-20 produced by MIM technology
Vplivi parametrov sintranja na gostoto, mikrostukturo in mehanske lastnosti z niobijem legiranga nerjavnega ognjevzdr`nega jekla
GX40CrNiSi25-20, izdelanega z MIM-tehnologijo
S. Butkovi}, M. Oru~, E. [ari}, M. Mehmedovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 185
PREDGOVOR/FOREWORD
Vsakovrstni materiali nas obkro`ajo na vsakem
koraku v na{em `ivljenju. Ali se kdaj vpra{amo, kako so
ti materiali nastali, kdo jih je razvil in izbolj{al, kako jih
najbolj{e uporabiti in kak{na je prihodnost materialov in
tehnologij? Odgovor na ta vpra{anja vsekakor ni
preprost, iskanje odgovorov pa omogo~a nadaljnji razvoj
in napredek dru`be, kot tudi obstoj revije Materiali in
tehnologije. Kot pove `e njeno ime, je revija namenjena
materialom in tehnologijam. Njen za~etek sega v leto
1967, ko je za~ela izhajati kot @elezarski zbornik, kjer so
svoja dognanja objavljali predvsem strokovnjaki sloven-
ske industrije jekla. Leta 1992 se je revija preimenovala
v Kovine zlitine tehnologije, kjer je svoja dela objavljalo
tudi vedno ve~ slovenskih raziskovalcev. Revija se je {e
enkrat preimenovala leta 2000, ko je dobila dana{njo
obliko in ime.
Revija Materiali in tehnologije je prehodila `e kar
~astitljivo pot razvoja, od svojih za~etkov kot @elezarski
zbornik, pa do dana{njih dni. Opa`amo, da postaja
cenjena v razli~nih strokovnih krogih, saj omogo~a
vpogled v dogajanje na podro~ju kovinskih, polimernih
in anorganskih materialov ter tehnologij s teh podro~ij.
Resno delo urednikov v preteklosti je omogo~ilo tudi
vklju~itev revije v sistem SCI, kar ji je {e pove~alo
dodano vrednost. Po sedanjih mednarodnih merilih se
kvaliteta revije izra`a s faktorjem SCI, ki je pomemben
del tudi pri oceni kvalitete publikacij raziskovalcev, ki v
njej objavljajo svoje ~lanke. Pove~uje se {tevilo sloven-
skih avtorjev, pove~uje pa se tudi interes avtorjev iz
tujine. To je dokaz, da gre razvoj revije v pravo smer.
Revija je pomembna tudi z nacionalnega stali{~a.
Ve~ina znanstvenih in strokovnih ~lankov je v angle{kem
jeziku, v slovenski jezik pa so prevedeni: naslov, povze-
tek, klju~ne besede, naslovi tabel in podnaslovi slik. To
prispeva k {irjenju slovenske strokovne terminologije,
kar je pomembno za bogatenje in nadaljnji razvoj
slovenskega jezika.
V decembru 2011 so bili imenovani nov glavni in
odgovorni urednik ter ~lani uredni{kega odbora. To
postavlja pred imenovane nove izzive, ki bodo omogo~ili
nadaljnji razvoj revije, njeno {irjenje po znanstveni in
strokovni javnosti ter ve~anje ugleda revije na podro~ju
materialov in tehnologij.
Prizadevali si bomo ~im bolj skraj{ati rok od sprejetja
~lanka do njegove objave. Pri tem imajo pomembno
vlogo tudi avtorji, saj kvalitetna vsebina in ustrezna
tehni~na priprava ~lanka olaj{ata delo vsem, ki
sodelujemo pri pripravi ~lankov za objavo.
Zahvaljujem se dosedanjemu glavnemu in odgovor-
nemu uredniku prof. dr. Francu Vodopivcu za ves trud in
prizadevanje, ki ga je vlo`il v revijo.
Glavni in odgovorni urednik
doc. dr. Matja` Torkar
A wide variety of materials surround us in everyday
life. However, do we always ask ourselves where these
materials came from, how these materials were prepared,
who developed them, who improved them, what is their
best application and what is the future of materials and
technologies? The answers to these questions are not easy,
but the search for the answers enables the further
development of society, as well as the existence of our
journal – Materials and Technology. As is clear from the
journal’s name, it is devoted to materials and technology.
The earliest issue dates back to 1967, when the journal
was called @elezarski zbornik: a periodical where
metallurgical engineers from Slovenian steelworks could
publish their results. In 1992 the journal changed its name
to Kovine zlitine tehnologije (Metals Alloys Technologies)
and a larger number of Slovenian researchers started to
publish articles. The last change of name was in 2000,
when the journal Materiali in tehnologije (Materials and
Technology) took its present name and form.
Materials and Technology has come a long way from
its beginnings as @elezarski zbornik. It is clear that the
journal is much appreciated in a variety of professional
circles, where it provides an insight into activities in the
fields of metallic materials, polymer materials and
ceramic materials, as well as the technology of these
fields. Intense efforts made by previous editors made it
possible for the journal to enter the SCI system, which
greatly increases the added value of the journal. After
some recent changes the quality of the journal is now
reflected in its SCI factor, which is an important element
in the evaluation of the quality of researchers. The
increasing number of Slovenian authors and the ever-
larger number of authors from abroad are evidence that
the journal is developing in the proper direction.
The journal is also important from the national point
of view. Most of the articles are published in English, but
with translations of the abstract, keywords, table titles,
figure captions into Slovene. This helps with the spread-
ing of Slovenian terminology, which is important for the
further development and enrichment of Slovene.
A new Editor-in-Chief and Editorial Board were
elected in December 2011. Their mission will be to ensure
the successful development of the journal Materials and
Technology, its broader recognition among material
scientists and an increase of the journal’s reputation in the
field of materials and technologies.
We will strive to reduce the time from the acceptance
to the publication of articles. However, an important role
is also played by the authors, as the quality of the content
and the technical correctness of the paper’s preparation
facilitates the work of all the staff included in the process
of publishing the articles.
Finally, I would like to express my thanks to the
former Editor-in-Chief, prof. Franc Vodopivec, for all his
hard work and the effort that he put into the journal.
Editor-in-Chief
A/Prof.Dr. Matja` Torkar

T. KROUPA et al.: LINEAR TWO-SCALE MODEL FOR DETERMINING THE MECHANICAL PROPERTIES ...
LINEAR TWO-SCALE MODEL FOR DETERMINING THE
MECHANICAL PROPERTIES OF A TEXTILE
COMPOSITE MATERIAL
LINEARNI DVOSTOPENJSKI MODEL ZA DOLO^ITEV
MEHANSKIH LASTNOSTI TEKSTILNEGA KOMPOZITA
Tomá{ Kroupa1, Petr Janda2, Robert Zem~ík3
1University of West Bohemia in Pilsen, Department of Mechanics, Univerzitní 22, 306 14, Plzeò, Czech Republic
2University of West Bohemia in Pilsen, Department of Machine design, Univerzitní 22, 306 14 Plzeò, Czech Republic
3University of West Bohemia in Pilsen, Department of Mechanics, Univerzitní 22, 306 14, Plzeò, Czech Republic
kroupa@kme.zcu.cz
Prejem rokopisa – received: 2011-02-01; sprejem za objavo – accepted for publication: 2011-10-10
The engineering mechanical constants for a description of mechanical macro-scale models of carbon and aramid textile
composite materials are calculated using finite-element analyses. Two sub-scale models of representative volumes are used. The
micro-scale model represents a periodically repeated volume consisting of fibers and a matrix within each interweaved yarn.
The meso-scale model represents a unit cell of four interweaved yarns, which is repeated within the whole composite with the
properties obtained from a micro-scale model and matrix. The finite-element models are built with the commercial packages
Siemens NX 7.5 and MSC.Marc 2008r1 using subroutines.
Keywords: composite, textile, linear, carbon, aramid, epoxy, tensile, finite-element analysis
Z uporabo metode kon~nih elementov so izra~unane in`enirske konstante za opis mehanskega makrodimenzionalnega
ogljik-aramidnega modela kompozita. Uporabljena sta dva poddimenzionalna modela za manj{e ustrezne prostornine.
Mikrodimenzionalni model je periodi~no ponavljanje prostornine, ki se ponavlja za ves kompozit za matico in vpleteno vlakno.
Mezodimenzionalni model je spletna celica iz {tirih vpletenih vlaken in se ponavlja v vsem kompozitu z lastnostmi
mikromodela in matice. Modeli kon~nih elementov so vgrajeni v paketa Siemens NX 7.5 in MSC.Marc 2008r1 z uporabo
podrutin.
Klju~ne besede: kompozit, tekstil, linearen, ogljik, aramid, epoksi, natezne lastnosti, kon~ni elementi
1 INTRODUCTION
A knowledge of the precise values of the mechanical
properties of materials is crucial for the capability of
models to predict the behavior of analyzed structures.
This is also the case for the modeling of composite
materials. Several material properties of the composites
can be calculated directly from experimental results
(Young’s moduli from tensile tests, etc.). The presented
paper is aimed at a determination of the elasticity
constants of textile composites using sub-scale models to
determine the properties that cannot be measured and
calculated directly from the experiment (shear modulus,
etc.). The models were used for the prediction of the
elasticity constants of two materials with a simple plain
weave. This type of material was chosen because of the
possibility to measure directly the Young’s moduli in the
principal material directions using tensile tests.
Nevertheless, the shear modulus was fitted on the linear
part of the measured curves using a gradient-opti-
mization algorithm and the Poisson’s ratios of the whole
textile composites were identified using a digital image
correlation1. The material data of the constituents were
given by the manufacturer and the dimensions of the
periodically repeated volume (unit-cell element – UCE)
of the textiles were measured using a digital camera.
2 EXPERIMENT
The effective material parameters were determined
using simple tensile tests performed on a Zwick/
Roell Z050 test machine on thin strips with the
dimensions given in Table 1.
Table 1: Dimensions of the strips
Tabela 1: Dimenzije traka
Carbon Aramid
Length mm 100.00 100.00
Width mm 10.00 10.00
Thickness mm 0.30 0.35
Three types of specimens were used for each
material. One of two principal directions of the textile
fabric form the angles 0°, 45° and 90° with the direction
of the loading force. Once the force-displacement
diagrams (Figures 1 and 2) were measured, the Young’s
moduli in directions 1 and 2 and the shear modulus were
fitted on the linear parts of the curves using a com-
bination of a plane-stress finite-element (FE) model and
the gradient-optimization algorithm implemented in
OptiSLang software (the methodology is described in2,3).
A digital image correlation1 was used for the calculation
Materiali in tehnologije / Materials and technology 46 (2012) 2, 97–101 97
UDK 66.017:519.61/.64:620.17 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)97(2012)
of the Poisson’s ratios. The elasticity parameters of the
textiles are shown in Table 2.
Table 2: Elasticity parameters of textile composites
Tabela 2: Parametri elasti~nosti za tekstilne kompozite
Carbon/Epoxy Aramid/Epoxy
E1 GPa 31.05 15.85
E2 GPa 29.73 15.66
G12 GPa 1.83 1.24
12 – 0.19 0.31
3 UNIT-CELL ELEMENTS
Axes orientations in UCE (Figure 4) are shown in
Figure 3. Dimensionless geometry of UCE of yarns is
shown in Table 3. The ideal distribution of fibers in the
yarns, the perfect saturation of the yarns by the matrix
and the volume fiber fractions Vf = 0.6 and Vf = 0.7 are
considered in the calculations.
Table 3: Dimensions of the unit cell of the yarn
Tabela 3: Dimenzije spletne celice
l1 [-] 1
l2 [-] 4
l3 [-] 4  3
T. KROUPA et al.: LINEAR TWO-SCALE MODEL FOR DETERMINING THE MECHANICAL PROPERTIES ...
98 Materiali in tehnologije / Materials and technology 46 (2012) 2, 97–101
Figure 4: UCE geometry of yarn with Vf = 0.6 (fibers – black, matrix
– gray)
Slika 4: UCE-geometrija vlakna z Vf = 0.6 (vlakna – ~rno, matica –
sivo)
Figure 2: Force-displacement diagrams (gray) for Aramid/Epoxy,
with fitted parts (black)
Slika 2: Odvisnost sila – premik (sivo) za aramid/epoksi s pribli`ki
(~rno)
Figure 1: Force-displacement diagram (gray) for Carbon/Epoxy, with
fitted parts (black)
Slika 1: Odvisnost sila – premik (sivo) za ogljik/epoksi s pribli`ki
(~rno)
Figure 3: Material axes: yarn (left), textile (right)
Slika 3: Osi materiala: vlakno (levo), tekstil (desno)
Figure 5: Photograph of Carbon/Epoxy specimen
Slika 5: Posnetek vzorca ogljik/epoksi
The dimensions of the UCE of both materials, which
were measured using detailed photographs provided by a
Canon EOS 400D digital camera (Figures 5 and 6) are
shown in Table 4.
Table 4: Dimensions of textile unit cells
Tabela 4: Dimenzije spletne celice tekstila
Carbon/Epoxy Aramid/Epoxy
l1 mm 5.00 3.00
l2 mm 5.00 3.00
l3 mm 0.30 0.35
4 BOUNDARY CONDITIONS
In the FE model of the UCE (Figure 7) it is
necessary to invoke pure tensile conditions or pure shear
conditions to determine the elasticity parameters.
Furthermore, the UCE has to be fixed in space to
eliminate rigid body modes. Finally, the periodic
boundary conditions have to be satisfied. The effect of
the periodic boundary condition on the UCE with two
periodically tied faces is sketched in Figure 8.
Each corresponding pair of nodes on opposite faces
of the FE model must fulfill the conditions 4–6
u u di i u
i
B A− = , v v d
i i
v
i
B A− = (P)
w w di i w
i
B A− = for i=1...N
where i is the index of the corresponding constrained
faces; N is the total number of the periodically con-
strained faces; u, v and w are the displacements in the 1,
2 and 3 direction; and d d du
i
v
i
w
i, , are displacements of
the appropriate retained nodes in which the loading is
applied (Figure 9). The UCE of the yarns is periodically
tied in all three directions and the UCE of the textiles is
tied in direction 1 and 2 (Figure 9).
For the determination of the Young’s modulus E1 and
the Poisson’s ratio 12 of the textile composite the model
is loaded with 1  0 and the other loadings are equal to
zero. Normal strains in direction 1 and 2 are calculated
as
1
1
1
1
=
d
l
, 2
2
2
2
=
d
l
(e12)
and the Young’s modulus and Poisson’s ratio are
E v1
1
1
12
2
1
= =




, (E112)
T. KROUPA et al.: LINEAR TWO-SCALE MODEL FOR DETERMINING THE MECHANICAL PROPERTIES ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 97–101 99
Figure 9: Boundary conditions and links used for periodic boundary
conditions for UCE of Aramid/Epoxy textile
Slika 9: Mejni pogoji in povezave, uporabljene za periodi~ne mejne
pogoje za UCE aramid/epoksi tekstil
Figure 6: Photograph of Aramid/Epoxy specimen
Slika 6: Posnetek vzorca aramid/epoksi
Figure 7: UCE geometry of Carbon/Epoxy textile composite (yarns –
black, matrix – gray)
Slika 7: UCE-geometrija kompozita ogljik/epoksi (trakovi – ~rno,
matica – sivo)
Figure 8: Scheme of the effect of the periodic boundary conditions
Slika 8: Shema u~inka periodi~nosti mejnih pogojev
For the determination of the shear modulus G12 the
model is loaded by 12  0. The other loadings are equal
to zero. The shear strain in plane 12 is calculated as
12
2
1
1
1
2
2
= +
d
l
d
l
(g12)
and the shear modulus is
G12
12
12
=


(G12)
The same scheme is used for the determination of the
elasticity constants in the other directions or planes.
5 INPUT PARAMETERS
Carbon (Toray T600) and Aramid (Twaron K1055)
fibers are transversely isotropic materials. Their elasti-
city parameters are given in Table 5.
Table 5: Elasticity parameters for fibers
Tabela 5: Parametri elasti~nosti za vlakna
Carbon Aramid
E1 GPa 230.00 104.00
E2 GPa 7.05 5.40
E3 GPa 7.05 5.40
12 – 0.30 0.40
23 – 0.30 0.40
31 – 0.02 0.02
G12 GPa 50.00 12.00
G23 GPa 50.00 12.00
G31 GPa 50.00 12.00
The matrix is manufactured from epoxy resin
(MGS® L 285) and hardener (MGS® 285). It is con-
sidered to be a linear isotropic material (Table 6).
Table 6: Elasticity parameters for the matrix
Tabela 6: Parametri elasti~nosti za matico
Epoxy
E GPa 3.00
 – 0.30
6 RESULTS
The effect of the periodic boundary conditions is
shown in Figures 10 and 11. Opposite faces of the UCE
are deformed in the same shape. The elastic properties of
the yarns are shown in Table 7. The results are shown
for both fiber volume fractions (Vf). Similarly, the results
for the textiles are shown for both Vf (Table 8).
Table 7: Calculated elasticity parameters of the yarns
Tabela 7: Izra~unani parametri elasti~nosti za spleta
Carbon/Epoxy Aramid/Epoxy
Vf – 0.60 0.70 0.60 0.70
E1 GPa 138.87 161.54 63.46 73.55
E2 GPa 7.05 8.33 4.36 4.60
E3 GPa 7.05 8.33 4.36 4.60
12 – 0.30 0.30 0.36 0.37
23 – 0.36 0.34 0.40 0.40
31 – 0.02 0.02 0.02 0.02
G12 GPa 4.26 5.90 3.42 4.34
G23 GPa 3.88 5.45 3.15 4.01
G31 GPa 4.26 5.90 3.42 4.34
Table 8: Calculated elasticity parameters for textiles
Tabela 8: Izra~unani parametri elasti~nosti za tekstil
Carbon/Epoxy Aramid/Epoxy
Vf – 0.60 0.70 0.60 0.70
E1 GPa 27.75 31.89 14.08 15.72
E2 GPa 27.75 31.89 14.08 15.72
G12 GPa 2.60 3.30 2.21 2.60
12 – 0.33 0.33 0.28 0.29
T. KROUPA et al.: LINEAR TWO-SCALE MODEL FOR DETERMINING THE MECHANICAL PROPERTIES ...
100 Materiali in tehnologije / Materials and technology 46 (2012) 2, 97–101
Figure 11: Deformed FE model of the UCE of the textile under shear
loading in plane 12 (shown values of shear stress 12)
Slika 11: Deformiran FE-model UCE za tekstil pri stri`ni obremenitvi
v ravnini 12 (prikazane vrednosti stri`ne napetosti 12)
Figure 10: Deformed FE model of the UCE of yarn under shear
loading in plane 23 (shown values of shear stress 23)
Slika 10: Deformiran FE-model za UCE-spleta pri stri`ni obremenitvi
v ravnini 23 (prikazane vrednosti stri`ne napetosti 23)
7 CONCLUSION
Multi-scale models for the prediction of the elasticity
constants of textile composites with a simple plain weave
were developed and presented. Good agreement of the
Young’s moduli was achieved between the calculated
and experimental values. However, the shear moduli are
slightly over-predicted. The calculated Poisson’s ratios
were calculated with acceptable accuracy only for the
Aramid/Epoxy textile. Future research will be aimed at
the non-linear, plastic and damage behavior of the
matrix, the damage behavior of the fibers and an
investigation of the imperfections and the unit-cell
element dimensions of the textiles.
Acknowledgement
The work has been supported by the projects GA
P101/11/0288 and European project NTIS – New
Technologies for Information Society No. CZ.1.05/
1.1.00/02.0090.
8 REFERENCES
1 R. Kottner, R. Zem~ík, V. La{, Mechanical characteristics of rubber
segment – shear test. In: Experimental Stress Analysis 2007
2 T. Kroupa, R. Zem~ík, J. Klepá~ek, Temperature dependence of
parameters of non-linear stress-strain relations for carbon epoxy
composites, Mater. Tehnol., 43 (2009) 2, 69–72
3 T. Kroupa, R. Zem~ík, V. La{, Improved non-linear stress-strain
relation for carbon-epoxy composites and identification of material
parameters, JCM, 45 (2011) 9, 1045–1057
4 P. P. Camanho, C. G. Dávila, S. T. Pinho, J. J. C. Remmers,
Mechanical response of Composites, ECCOMAS, Springer 2008
5 B. Hassani, E. Hinton, Homogenization and structural topology
optimization, Springer 1999
6 R. Zem~ík, K. Kunc, T. Kroupa, R. Kottner, P. Janda, Micromodel
for failure analysis of textile composites, In: YSESM, Trieste, Italy,
2010, 48–51
T. KROUPA et al.: LINEAR TWO-SCALE MODEL FOR DETERMINING THE MECHANICAL PROPERTIES ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 97–101 101

U. KÖKLÜ: INFLUENCE OF THE PROCESS PARAMETERS AND THE MECHANICAL PROPERTIES ...
INFLUENCE OF THE PROCESS PARAMETERS AND THE
MECHANICAL PROPERTIES OF ALUMINUM ALLOYS
ON THE BURR HEIGHT AND THE SURFACE
ROUGHNESS IN DRY DRILLING
VPLIV PARAMETROV PROCESA IN MEHANSKIH LASTNOSTI
ALUMINIJEVIH ZLITIN NA VI[INO IGLE IN HRAPAVOST
POVR[INE PRI SUHEM VRTANJU
Ugur Köklü
Department of Manufacturing Engineering, Faculty of Technology, University of Dumlupýnar, 43500, Simav-Kütahya, Turkey
ugurkoklu@gmail.com
Prejem rokopisa – received: 2011-05-10; sprejem za objavo – accepted for publication: 2011-10-28
In this paper, the effect of the mechanical properties of aluminum alloys, cutting speed, feed rate and the drill diameter on burr
height and surface roughness of drilling holes were investigated, using the Taguchi method. Al-2024, Al-7075 and Al-7050 were
selected as the workpiece materials for experiments. The analysis of variance and signal-to-noise ratio were employed to
analyze the effect of the drilling parameters. The results of the statistical analysis indicated that feed rate and cutting speed
minimize significantly both the height of the exit burrs and the surface roughness. Moreover, the mechanical properties of the
workpieces are different influential factors on both responses.
Keywords: drilling, burr height, surface roughness, Al-2024, Al-7075, Al-7050
Raziskan je vpliv mehanskih lastnosti aluminijevih zlitin, hitrosti rezanja in podajanja ter premera svedra na vi{ino igle in
hrapavost povr{ine vrtane povr{ine z uporabo Taguchi metode. Za preizkuse so bile izbrane zlitine Al-2024, Al-7075 in
Al-7050. Analize variance in intenzitete signal-ozadje so bile uporabljene za dolo~itev vpliva parametrov vrtanja. Rezultati
statisti~ne analize so pokazali, da hitrosti podajanja in rezanja zmanj{ujeta vi{ino izhodne igle in hrapavost povr{ine. Tudi
mehanske lastnosti obdelovanca imajo razli~en vpliv na oba odgovora.
Klju~ne besede: vrtanje, vi{ina igle, hrapavost povr{ine, Al-2024, Al-7075, Al-7050
1 INTRODUCTION
Drilling is one of the most important material
removal process that has been widely used in the
aerospace, aircraft and automotive industries. Although
modern metal-cutting methods, including electron-beam
machining, ultrasonic machining, electrolytic machining
and abrasive jet machining, have improved in the manu-
facturing industry, conventional drilling still remains one
of the most common machining processes1,2. Aluminum
is used in many industrial areas to make different
products and it is significant for the world economy.
Structural components made from aluminum and
aluminum alloys are vital in the aerospace industry and
very important in other areas of transportation and
building in which durability, strength and light weight
are expected3.
The drilling process produces burrs on both the
entrance and the exit surfaces of the workpiece. The exit
burr is part of the material extending off the exit surface
of the workpiece. Most burr-related problems in drilling
are caused by the exit burr because it is much larger than
the entrance burr4. The presence of these exit burrs
requires additional manufacturing steps for disassembly
and deburring. These additional steps are typically not
easy to automate and are generally performed manually5.
Burr formation affects workpiece accuracy and quality in
several ways: deterioration of the surface quality,
dimensional distortion on the part edge, challenges to
assembly and handling caused by burrs in sensitive
locations on the work and damage done to the workpiece
subsurface from the deformation associated with burr
formation6.
In several studies burr formation and surface
roughness were investigated. Nouari et al.7 examined the
effect of the machining parameters on the hole-surface
roughness and diameter deviations for different coated
drills. The results show that, small constant feed rate,
low cutting speeds are appropriate for the dry machining
of AA-2024. Kilickap6 presented an application of
Taguchi and response surface methodologies for
minimizing the burr height and the surface roughness in
drilling Al-7075. The optimization results showed that
the combination of low cutting speed, low feed rate and
high point angle is necessary to minimize both burr
height and surface roughness. Kurt et al.3 investigated the
role of different coatings, point angle and cutting
parameters on the hole quality in the drilling of Al-2024
alloy and concluded that the cutting parameters and the
coatings have different effects on hole quality. They have
Materiali in tehnologije / Materials and technology 46 (2012) 2, 103–108 103
UDK 669.715:621.95:620.179.11 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)103(2012)
obtained effective results using a low cutting speed and
feed rate. Ko and Lee8 used several materials that were
drilled by several cutting conditions, velocity and feed
rate. They indicated that burr formations were highly
dependent on the material properties, the drill geometry
and the cutting condition. Lauderbaugh9 used simulation
tools and analysis of variance to identify the influence of
process parameter on the height of exit burrs and
concluded that feed rate, chisel-edge-to-drill diameter
ratio, drill diameter, yield strength and point angle are
significant for the height of exit burrs. Ko et al.10 carried
out an experimental investigation of the role of various
shapes of drills and materials (SM45C, SS400,
A6061-T6 and A2024-T4) on the burr in drilling. Their
experimental results showed that the burr height from
ductile materials is larger than from brittle materials.
Kurt et al.11 investigated the influence of the cutting
parameters and the mechanical properties of a workpiece
on the burr formations in a dry drilling process. His
experimental results showed that the machining
parameters and the mechanical properties of a workpiece
effect the burr formation. In addition to these they have
classified the burrs into three types. Kalidas et al.12
compared the performance of the three types of coatings
on the hole quality under dry- and wet-drilling condi-
tions of aluminum alloys. The result of the experiments
indicates that, the use of coatings did not seem to affect
the surface roughness of the hole produced.
In this study a statistical analysis of the experimental
data of the cutting parameters and the mechanical pro-
perties of aluminum alloys on the burr height and surface
roughness of the produced hole in the dry drilling of
Al-2024, Al-7075 and Al-7050 have been investigated
and analyzed with the Taguchi method.
2 MATERIALS, CUTTING CONDITIONS AND
PLAN OF EXPERIMENTS
Burr height and the surface roughness of the drilled
hole surface were determined by cutting condition. The
drilling experiments were conducted in dry cutting
conditions on a Johnford VMC model three axes CNC
milling machine with a Fanuc controller. In this study,
Al-2024, Al-7075 and Al-7050 were chosen as the work
materials with the specimen dimensions 200 mm × 140
mm × 30 mm. The mechanical properties of the three
aluminum alloys are presented in Table 1.
Uncoated, conventional, high-speed-steel twist drills
with diameters of 8 mm, 10 mm and 12 mm were used
for drilling experiment. A new drill bit was used for each
drilling experiment. The burr heights (H) and surface
roughness (Ra, the arithmetic average) of each drilled
hole were measured by means of an optical microscope
and a Mahr Perthometer surface roughness tester using a
meter cut off of 5.6 mm. The burr height and the surface
roughness of the machined hole measurement points of
the workpiece are shown in Figure 1. Each specimen
was measured from four different points (0°, 90°, 180°
and 270°) for both, burr height and surface roughness.
The drilling experiments were planned using Tagu-
chi’s orthogonal array. Three experimental parameters
were the cutting speed, feed rate and drill diameter were
selected for the present investigation. Three levels of
each control factor were taken into account. Taguchi’s
orthogonal array of L27 was chosen for the experimental
plan. The considered experimental factors and their
levels are listed in Table 2.
Table 1: Mechanical properties of Al-2024, Al-7075 and Al-7050
materials
Tabela 1: Mehanske lastnosti zlitin Al-2024, Al-7075 in Al-7050
Materials
Tensile
Strength
(MPa)
Yield
Strength
(MPa)
Elongation
%
Hardness
(HRB)
Al-2024 469 324 17 75
Al-7075 570 505 11 87
Al-7050 521 467 10 84
Table 2: Control factors and their levels used for drilling experiments
Tabela 2: Kontrolni dejavniki in njihov nivo, uporabljen pri preizku-
sih vrtanja
Symbol Factors Unit
Level
1 2 3
A Cutting speed m/min 20 30 40
B Feed rate mm/r 0.05 0.1 0.15
C Drill diameter mm 8 10 12
3 EXPERIMENTAL RESULTS AND ANALYSIS
The Taguchi method is very popular for solving
optimization problems in the field of manufacturing
engineering13. In this method, the term "signal" (S)
represents the desired value and the "noise" (N)
represents the undesired value. The objective of using
the S/N ratio is a measure of the performance to develop
products and processes that are insensitive to noise
factors. The S/N ratio indicates the degree of predictable
performance of a product or process in presence of noise
factors. The process parameter settings with the highest
S/N ratio always yield the optimum quality with mini-
mum variance. The difference between the functional
U. KÖKLÜ: INFLUENCE OF THE PROCESS PARAMETERS AND THE MECHANICAL PROPERTIES ...
104 Materiali in tehnologije / Materials and technology 46 (2012) 2, 103–108
Figure 1: Burr height and surface-roughness measurement points of
the produced hole
Slika 1: To~ke meritev vi{ine igle in hrapavosti povr{ine pri izvrtinah
value and the objective value is emphasized and iden-
tified as the loss function. The loss function is derived as
Eq. (1)
L y
L m y m
k y m k MSD( )
"( )( )
!
( ) ( )=
−
= − =
2
2
2
(1)
where L(y) is the loss function, y is the value of the
quality characteristic, m is the target value of y, k is the
commensurately constant, which depends on financial
criticality of y, and MSD is the mean square deviation.
Eq. (1) can be expressed by the signal-to-noise ratio (ç)
and can be rewritten as:
	 = −10 10lg ( )MSD (2)
The value of the loss function is further transformed
into a signal-to-noise (S/N) ratio. In the present investi-
U. KÖKLÜ: INFLUENCE OF THE PROCESS PARAMETERS AND THE MECHANICAL PROPERTIES ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 103–108 105
Table 3: Experimental layout using L27 orthogonal array and experimental values
Tabela 3: Na~rt preizkusov z uporabo ortogonalne razporeditve L27 in vrednosti preizkusov
Trial no. Factor Level Al-2024 Al-7075 Al-7050
A B C H/mm Ra/μm H/mm Ra/μm H/mm Ra/μm
1 1 1 1 3.29 6.200 1.43 2.967 1.14 2.042
2 1 1 2 3.32 6.188 1.55 3.061 1.45 2.275
3 1 1 3 3.54 6.484 1.68 3.073 1.60 2.352
4 1 2 1 4.27 6.395 1.64 3.112 1.22 2.331
5 1 2 2 4.40 6.580 1.88 3.335 1.55 2.422
6 1 2 3 4.51 6.934 1.92 3.429 1.65 2.421
7 1 3 1 5.28 6.883 2.76 3.265 2.21 2.487
8 1 3 2 4.84 6.930 2.59 3.312 2.25 2.529
9 1 3 3 5.06 7.234 2.62 3.451 2.45 2.587
10 2 1 1 3.04 6.847 1.73 3.088 1.28 2.242
11 2 1 2 3.22 7.090 2.12 3.208 1.89 2.421
12 2 1 3 3.30 7.158 2.30 3.322 1.31 2.582
13 2 2 1 4.04 6.906 1.87 3.457 1.98 2.627
14 2 2 2 3.81 7.120 1.98 3.742 2.42 2.728
15 2 2 3 5.27 7.476 2.88 3.751 2.38 2.838
16 2 3 1 6.28 7.118 2.44 3.634 2.11 2.854
17 2 3 2 6.44 7.515 2.65 3.361 2.20 2.975
18 2 3 3 6.61 7.700 2.93 3.814 2.56 3.081
19 3 1 1 3.51 7.322 1.93 3.676 1.72 2.679
20 3 1 2 5.75 7.441 1.93 3.332 1.45 2.751
21 3 1 3 5.01 7.397 2.13 3.751 1.97 2.883
22 3 2 1 6.21 7.809 2.92 3.209 2.21 2.667
23 3 2 2 5.57 7.930 2.78 3.831 2.55 2.948
24 3 2 3 6.31 7.974 2.42 3.928 2.22 3.027
25 3 3 1 6.58 7.970 2.97 3.287 2.63 3.031
26 3 3 2 6.42 8.114 3.75 4.107 2.67 2.942
27 3 3 3 4.89 7.939 3.78 4.321 3.53 3.288
Table 4: Response table for burr height and surface roughness
Tabela 4: ANOVA-rezultati za vi{ino igle
Factors
Mean S/N ratios (dB) for H Mean S/N ratios (dB) for Ra
Level 1 Level 2 Level 3 Level 1 Level 2 Level 3
Al–2024
Cutting speed –12.502 –12.99 –14.80 –16.441 –17.16 –17.80
Feed rate –11.341 –13.72 –15.23 –16.762 –17.16 –17.47
Drill diameter –13.143 –13.47 –13.68 –16.943 –17.13 –17.33
Al–7075
Cutting speed –5.8132 –7.181 –8.482 –10.151 –10.83 –11.36
Feed rate –5.3261 –6.874 –9.276 –10.282 –10.94 –11.12
Drill diameter –6.5163 –7.173 –7.788 –10.353 –10.79 –11.20
Al–7050
Cutting speed –4.4502 –5.851 –7.073 –7.5241 –8.606 –9.269
Feed rate –3.5881 –5.886 –7.900 –7.8052 –8.490 –9.103
Drill diameter –4.9243 –6.002 –6.448 –8.0763 –8.478 –8.844
1, 2 and 3: Optimum level and Rank
gation, the objective is to minimize the burr height and
the surface roughness; therefore, "smaller is better" as a
quality characteristic is selected, which is a logarithmic
function given as:
S/N
r
Ri
i
r
( ( )	
 = −
=
∑10 110 2
1
lg i=1,2,...,r (3)
where Ri is the value of the burr height or the surface
roughness for the ith trial in r number of measure-
ments14.
The experimental values obtained from the experi-
ments related to burr height and surface roughness are
illustrated in Table 3. The S/N ratios for the burr height
(H) and the surface roughness (Ra) were calculated using
the output parameter values given in Table 3. The S/N
ratio for each parameter level was calculated by
averaging the S/N ratios obtained when the parameter
maintained at that level. Table 4 shows the S/N ratio
obtained for different parameter levels.
The response graphs for the S/N ratios of the burr
height and the surface roughness are shown in Figures 2
and 3, respectively. It is observed from the S/N response
graph that the optimum parameter level combinations for
the minimum values of Al-2024, Al-7075 and Al-7050
are A1B1C1, for both burr height and surface roughness.
As shown in Table 4 and Figure 2, the feed rate is the
dominant parameter on the burr height followed by the
cutting speed. The drill diameter has a lower effect on
burr height. Although a lower burr height is always
preferred, burr formation in drilling is not desirable. In
U. KÖKLÜ: INFLUENCE OF THE PROCESS PARAMETERS AND THE MECHANICAL PROPERTIES ...
106 Materiali in tehnologije / Materials and technology 46 (2012) 2, 103–108
Table 5: The result of ANOVA for burr height
Tabela 5: ANOVA-rezultati za vi{ino igle
Factors Dof SS V F P/%
Al-2024
Cutting speed 2 8.0732 4.0366 7.60* 21.31
Feed rate 2 18.9565 9.4782 17.84* 50.04
Drill diameter 2 0.2276 0.1138 0.21 0.60
Error 20 10.6262 0.5313 28.05
Total 26 37.8834 100
Al-7075
Cutting speed 2 2.3905 1.1953 13.76* 23.98
Feed rate 2 5.3525 2.6762 30.80* 53.68
Drill diameter 2 0.4903 0.2451 2.82 4.92
Error 20 1.7376 0.0869 17.42
Total 26 9.9709 100
Al-7050
Cutting speed 2 1.6389 0.8194 10.37* 20.27
Feed rate 2 4.3023 2.1511 27.23* 53.19
Drill diameter 2 0.5671 0.2835 3.59* 7.01
Error 20 1.5800 0.0790 19.53
Total 26 8.0883 100
* Significant at 95 % confidence level
Table 6: The result of ANOVA for surface roughness
Tabela 6: ANOVA-rezultati za hrapavost povr{ine
Factors Dof SS V F P/%
Al-2024
Cutting speed 2 5.6317 2.8159 131.94* 69.83
Feed rate 2 1.5560 0.7780 36.45* 19.30
Drill diameter 2 0.4501 0.2250 10.54* 5.58
Error 20 0.4268 0.0213 5.29
Total 26 8.0646 100
Al-7075
Cutting speed 2 1.09547 0.54773 13.10* 35.90
Feed rate 2 0.56992 0.28496 6.81* 18.68
Drill diameter 2 0.54954 0.27477 6.57* 18.00
Error 20 0.83654 0.04183 27.42
Total 26 3.05146 100
Al-7050
Cutting speed 2 1.28385 0.64192 87.02* 54.05
Feed rate 2 0.69896 0.34948 47.38* 29.43
Drill diameter 2 0.24479 0.12240 16.59* 10.31
Error 20 0.14753 0.00738 6.21
Total 26 2.37513 100
*Significant at 95 % confidence level
Figure 2: Effect of cutting parameters on the burr height: a) Al-2024,
b) Al-7075 and c) Al-7050
Slika 2: Vpliv parametrov rezanja na vi{ino igle: a) Al-2024, b) Al-
7075 in c) Al-7050
the present investigation, when cutting speed 20 m/min,
feed rate 0.05 mm/r and drill diameter 8 mm are used,
the burr height is minimized. The height of the exit burr
increases as the feed rate, the cutting speed and the drill
diameter increase. As shown in Table 4 and Figure 3,
cutting speed is the dominant parameter on surface
roughness, followed by the feed rate. The drill diameter
has a lower effect on surface roughness. In the present
investigation, when applied by cutting speed 20 m/min,
feed rate 0.05 mm/r and the drill diameter 8 mm, the
surface roughness is minimized. The roughness of the
drilled surface increases as the feed rate, the cutting
speed and the drill diameter increase.
The results of the analysis of variance (ANOVA) for
the burr height are presented in Table 5. From the
analysis, for all three aluminum alloys the feed rate is a
highly significant factor and plays a major role in
affecting the burr height. It can be observed from Table
5 that cutting speed also affects the burr height. The
effect of the drill diameter does not make any impact on
the responses, except for Al-7050. Percent (%) is
described as the significance rate of the process para-
meters on the burr height. It can be observed from the
ANOVA Table that the cutting speed, feed rate and drill
diameter are effect on the burr height 21.31 %, 50.04 %
and 0.60 %; 23.98 %, 53.68 % and 4.92 %; 20.27 %,
53.19 % and 7.01 % in drilling of Al-2024, Al-7075 and
Al-7050, respectively.
The results of the ANOVA for the surface roughness
are presented in Table 6. From the analysis, for the all
three aluminum alloys the cutting speed is a highly
significant factor and plays a major role in affecting the
surface roughness. It can be observed from Table 6 that
the feed rate and the drill diameter also affect the surface
roughness. It can be observed from the ANOVA Table
that cutting speed, the feed rate and the drill diameter are
effect on the surface roughness 69.83 %, 19.30 % and
5.58 %; 35.90 %, 18.68 % and 18.00 %; 54.05 %, 29.43
% and 10.31 % drilling of the Al-2024, Al-7075 and
Al-7050, respectively.
A series of experiment were conducted on three types
of aluminum. The properties of the workpiece material
have a significant influence on the burr height15. The burr
formation process is heavily dependent on the yield
strength, ultimate strength9 and ductility4. Also consi-
dering the ductility of materials represented as elon-
gation8 in Table 1 for the alloys Al-2024, Al-7075 and
Al-7050. The higher value of elongation represents
better ductility of the material. Al-2024 shows more
ductility than the Al-7075 and Al-7050 alloys. The
elongation percentage of workpieces used in the
experiments affects the formation of the burr height and
the surface roughness. The amount of burr around the
hole, which is drilled in Al-2024 alloy material is greater
for Al-7075 and Al-7050, because Al-2024 is more
ductile than Al-7075 and Al-7050. Also the difference of
burr height in Al-7075 and Al-7050 is not large, Al-7050
produces the smaller burr. As a result, much more burr
occurs in ductile materials. This tendency was also
mentioned by various other researchers4,8–10. In summary,
burrs are formed as a result of plastic deformation and
fracture. The final burr geometry determined by the
amount of plastic deformation is determined by the
ductility of the material represented as elongation8.
Al-7050 alloy machined surface, shows a lower value
of the surface roughness compared to Al-2024 and
Al-7075 alloys. Higher surface roughness values of
Al-2024 alloy can be explained by the highly ductile
nature of the alloy, which increases the tendency to form
a built-up edge (BUE). A relatively higher workpiece
ductility increases the BUE formation tendency16. The
presence of the BUE in the drilling process causes an
increase in the tool wear and a rougher surface finish.
U. KÖKLÜ: INFLUENCE OF THE PROCESS PARAMETERS AND THE MECHANICAL PROPERTIES ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 103–108 107
Figure 3: Effect of cutting parameters on the surface roughness: a)
Al-2024, b) Al-7075 and c) Al-7050
Slika 3: Vpliv parametrov rezanja na hrapavost povr{ine: a) Al-2024,
b) Al-7075 in c) Al-7050
4 CONCLUSIONS
In order to minimize the burr height and the surface
roughness of Al-2024, Al-7075 and Al-7050, the effects
of various cutting parameters have been investigated in
drilling using the Taguchi method and the analysis of
variance. Based on the S/N ratios and the ANOVA results
it is concluded that feed rate was the most influential
controllable factor among input parameters which affect
the burr height. The cutting speed was the second factor
at burr formation. The drill diameter has the lowest effect
on burr height. In view of the surface roughness, cutting
speed is a dominant parameter and it is followed by feed
rate and drill diameter, respectively. Moreover, the best
parametric combination of the three control factors
minimizing both the burr height and the surface
roughness were as follows: 20 m/min cutting speed, 0.05
mm/r feed rate and 8 mm drill diameter.
The mechanical properties of the workpieces are an
influential factor on burr height and surface roughness
formed at the hole. Due to the ductility of the material,
the amount of burr around the hole in Al-2024 alloy
material is much more than in Al-7075 and Al-7050, and
it can be explained with elongation. In addition, the
surface roughness obtained by Al-2024 is worse than by
Al-7075 and Al-7050 alloys. The higher surface rough-
ness values of the Al-2024 alloy can be explained by the
highly ductile nature of the alloy.
It is highly important to avoid burr formation, to
minimize it or to take it under control as an additional
manufacturing step is needed in order to be able to
eliminate the burrs formed by drilling. Minimization of
burr formation is an important problem of manufactur-
ing. The analysis of variance and Taguchi techniques
were applied in order to determine the effects of the
drilling parameters. Through the utilizing optimal
conditions obtained with S/N ratio, the burr around the
hole is minimized which contibutes the reduction of the
overall manufacturing cost by reducing the number of
processing requirement.
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U. KÖKLÜ: INFLUENCE OF THE PROCESS PARAMETERS AND THE MECHANICAL PROPERTIES ...
108 Materiali in tehnologije / Materials and technology 46 (2012) 2, 103–108
M. O. SHABANI, A. MAZAHERY: THE PERFORMANCE OF VARIOUS ARTIFICIAL NEURONS INTERCONNECTIONS ...
THE PERFORMANCE OF VARIOUS ARTIFICIAL
NEURONS INTERCONNECTIONS IN THE MODELLING
AND EXPERIMENTAL MANUFACTURING OF THE
COMPOSITES
PREDSTAVITEV RAZLI^NIH UMETNIH NEVRONSKIH POVEZAV
PRI MODELIRANJU IN EKSPERIMENTALNI IZDELAVI
KOMPOZITOV
Mohsen Ostad Shabani, Ali Mazahery
Karaj Branch, Islamic Azad University, Karaj, Iran
vahid_ostadshabany@yahoo.com
Prejem rokopisa – received: 2011-05-31; sprejem za objavo – accepted for publication: 2011-07-06
This study reports the performance of different artificial neural network (ANN) training algorithms in the prediction of
mechanical properties. First, an experimental investigation was carried out on the mechanical behavior of an A356 composite
reinforced with B4C particulates and then an ANN modeling was implemented in order to predict the mechanical properties,
including the yield stress, UTS, hardness and elongation percentage. After the preparation of the training set, the neural network
was trained using different training algorithms, hidden layers and the number of neurons in hidden layers. The test set was used
to check the system accuracy for each training algorithm at the end of the learning. The results show that the
Levenberg-Marquardt learning algorithm gave the best prediction for the yield stress, UTS, hardness and elongation percentage
of the A356 composite reinforced with B4C particulates.
Keywords: composite, hardness, mechanical properties, ANN
V tem delu smo najprej opredelili mehanske lastnosti, vklju~no z mejo plasti~nosti, natezno trdnostjo, trdoto in raztezkom
kompozita A356, oja~enega z delci B4C, in nato uporabili kombinacijo umetne nevronske mre`e in metode kon~nih elementov.
Po pripravi treningpostavitve je bila nevronska mre`a preizku{ena z uporabo razli~nih algoritmov, skritih plasti in {tevila
nevronov v skritih plasteh. Treningpostavitev je bila uporabljena za preverjanje natan~nosti za vsak algoritem na koncu u~enja.
Rezultati ka`ejo, da da Levenberg-Marquardtov u~ni logaritem najbolj{o napoved meje plasti~nosti, natezne trdnosti, trdote in
raztezka za kompozit A356, ki je oja~en z delci B4C.
Klju~ne besede: kompozit, trdota, mehanske lastnosti, ANN
1 INTRODUCTION
Large quantities of castings are made each year from
the aluminium alloy A356 (also known as Al-7Si-
0.3Mg). This alloy is one of the most popular alloys used
in industry due to its high fluidity and good "casta-
bility"1–5.
The addition of hard particles to a ductile metal
matrix produces a material whose mechanical properties
are intermediate between the matrix alloy and the
ceramic reinforcement. The casting cooling rate, the
reinforcement volume fraction, size, shape, and spatial
distribution are the most important parameters, playing a
role in the enhancement of the composite’s mechanical
properties. A stronger adhesion at the particle/matrix
interface improves the load transfer, increasing the yield
strength and stiffness, and delays the onset of
particle/matrix de-cohesion6.
An ANN is a logical structure with multi-processing
elements, which are connected through interconnection
weights. The knowledge is represented by the inter-
connection weights, which are adjusted during the learn-
ing phase. This technique is especially valuable in
processes where a complete understanding of the
physical mechanisms is very difficult, or even impossible
to acquire, as in the case of material properties where no
satisfactory analytical model exists7–14.
The aim of this study was to investigate the
prediction performance of various training algorithms
using a neural network computer program for the
mechanical properties of the A356 composite reinforced
with B4C particulates. The results showed that the
Levenberg-Marquardt learning algorithms gave the best
result for this study.
2 EXPERIMENTAL
In this study, A356 was used as the matrix material
and different volume fractions of B4C particles (1 % to
15 % B4C) with particle sizes ranging from 1 μm to 5 μm
were used as the reinforcements.
The melt-particle slurry was produced by a mecha-
nical stirrer. Approximately 5 kg of A356 alloy was
charged into the graphite crucible and heated up to a
temperature above the alloy’s melting point (750 °C).
The graphite stirrer, fixed on the mandrel of the drilling
Materiali in tehnologije / Materials and technology 46 (2012) 2, 109–113 109
UDK 620.17:519.61/.64:669.018.95 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)109(2012)
machine, was introduced into the melt and positioned
just below the surface of the melt. It was stirred at
approximately 600 r/min and 750 °C. Then the step
casting was poured into the CO2-sand mould.
Microscopic examinations of the composites and
matrix alloy were carried out using an optical micro-
scope. The porosity measurements of the composites
were obtained using Archimedes’s method. Hardness
and tensile tests were used to assess the mechanical
behavior of the composites and the matrix alloy.
2.1 Prediction of cooling rate and temperature gra-
dient with EEM
The numerical model is applied to simulate the
solidification of binary alloys; the mathematical
formulation of this solidification problem is given15:
C
T x y z t
t
K T x y z t q
∂
∂
( , , , )
( , , , )= ∇ +2 (1)
where /(kg/m3) is the density, K/(W/(m K)) is the
thermal conductivity, C/(J/(kg K)) is the specific heat,
q/(W/m3) is the rate of energy generation, T/K is the
temperature, and t/s is the time.
The release of latent heat between the liquidus and
solidus temperatures is expressed by:
q L
f
t
= 


s
(2)
where L/(J/kg) is the latent heat and fs is the local solid
fraction.
The fraction of solid in the mushy zone is estimated
by the Scheil equation, which assumes perfect mixing in
the liquid and no solid diffusion. With the liquidus and
solidus having constant slopes, fs is then expressed as:
f
T T
T T
k
s
f
f liq
= −
−
−
⎛
⎝
⎜
⎞
⎠
⎟
−
1
1 10/ ( )
(3)
where Tf/K is the melting temperature, TLiq/K is the
liquidus temperature and k0 is the partition coefficient.
Then15:


f
t k T T
T T
T T
k k
s
f liq
f
f liq
=
− −
−
−
⎛
⎝
⎜
⎞
⎠
⎟
− −
1
10
2 0 0
( )( )
( )/ ( 1 )


T
t
(4)
The latent heat released during the solidification of
the remaining liquid of eutectic composition was taken
into account by a device that considers a temperature
accumulation factor.
C
T x y z t
t
K T x y z t q’
( , , , )
( , , , )
∂
∂
= ∇ +2 (5)
where C’ can be considered as a pseudo-specific heat
given by:
C C L
f
T
’ _= M
s

(6)
C f C f CM s s s= − +( )1 1 (7)
where the subscripts L, S and M refer to liquid, solid
and mushy, respectively. The other properties, such as
the thermal conductivity and the density in the mushy
zone, are described in a similar way to the specific heat:
K f K f KM s s s= − +( )1 1 (8)
  M s s s= − +( )1 1f f (9)
The finite-element method (FEM) was used for
discretization. Based on the above transient-temperature
model, the FEM method is used to calculate the transient
temperature, cooling rate and temperature gradient (G).
2.2 Neural network training algorithms
There are various training algorithms used in neural
network applications. However, it is difficult to predict
which of these will be the fastest one for any problem.
Generally, it depends on some factors: the structure of
the networks, in other words, the number of hidden
layers, weights and biases in the network, aimed error
during the learning, and application area, for instance,
pattern recognition or classification or the function
approximation problem. However, the data structure and
the uniformity of the training set are also important
factors that affect the system accuracy and performance.
Some of the famous training algorithms are as
follows7–14,16–26:
Resilient back propagation (Rprop): is a network
training function that updates weight and bias values
according to the Rprop algorithm.
Random order incremental training with learning
functions: trains a network with weight and bias learning
rules using incremental updates after each presentation
of an input. Inputs are presented in a random order.
Gradient descent back propagation: is a network
training function that updates weight and bias values
according to the gradient descent.
BFGS quasi-Newton back propagation: is a network
training function that updates weight and bias values
according to the BFGS quasi-Newton method.
Bayesian regularization: is a network training
function that updates the weight and bias values accord-
ing to LM optimization. It minimizes a combination of
squared errors and weights, and then determines the
correct combination so as to produce a network that
generalizes well. The process is called Bayesian regulari-
zation.
In the analysis of the performance of various training
algorithms, the same prepared learning and test set were
used in the training processes of each learning algorithm.
The performance analyses were made from the
viewpoint of training duration, error minimization and
prediction achievement. The neural network predictions
were directly compared with the experimentally obtained
data to evaluate the learning performance. The mean
square error (MSE), which is a statistical and scientific
M. O. SHABANI, A. MAZAHERY: THE PERFORMANCE OF VARIOUS ARTIFICIAL NEURONS INTERCONNECTIONS ...
110 Materiali in tehnologije / Materials and technology 46 (2012) 2, 109–113
error-computation method, was used to analyze the
error25.
3 RESULTS AND DISCUSSION
Microscopic examinations were carried out on the
metal-matrix composite. Figure 1 shows that the B4C
particles were distributed between the dendrite branches
and were frequently clustered together, leaving the
dendrite branches as particle-free regions in the material.
Figure 2 shows the variation of porosity with B4C
content. It indicates that an increasing amount of
porosity is observed with increasing the volume fraction
of the composites. The porosity level increased, since the
contact surface area was increased27–31.
Figure 3 displays the results of the hardness tests.
The hardness of the MMCs increases with the volume
fraction of particulates in the alloy matrix. The higher
hardness of the composites could be attributed to the fact
that the B4C particles act as obstacles to the motion of
dislocations32–36. Figure 4 shows the typical stress-strain
curves obtained from uniaxial tension tests. The consi-
derable increase in strain-hardening observed during the
plastic deformation of composites is rationalized by the
resistance of the hard reinforcing particles to the slip
behavior of the Al matrix. The elongation to fracture of
the composite materials was found to be very low, and no
necking phenomenon was observed before fracture. On
the other hand, the elongation to fracture of the
un-reinforced Al alloy was about 15 %.
The input and output data set of the model is
illustrated schematically in Figure 5. In Figure 6, the
obtained MSE values for training data were given for
each training algorithm. The obtained error values for
M. O. SHABANI, A. MAZAHERY: THE PERFORMANCE OF VARIOUS ARTIFICIAL NEURONS INTERCONNECTIONS ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 109–113 111
Figure 4: Stress-strain curves for volume fractions Al/ 3 % B4C (B),
Al/ 7 % B4C (C), Al/ 10 % B4C (D), Al/ 12 % B4C (M) and Al/ 15 %
B4C (N)
Slika 4: Odvisnosti napetost – deformacija za volumenske dele`e Al/
3 % B4C (B), Al/ 7 % B4C (C), Al/ 10 % B4C (D), Al/ 12 % B4C (M)
in Al/ 15 % B4C (N)
Figure 2: Variations of porosity as a function of the volume fraction
of B4C
Slika 2: Spremembe poroznosti v odvisnosti od volumenskega dele`a
B4C
Figure 3: Variations of the hardness value of the samples as a function
of the volume fraction of B4C
Slika 3: Spremembe trdote vzorcev v odvisnosti od volumenskega
dele`a B4C
Figure 1: Typical optical micrographs: a) the composite with the
volume fraction of B4C 4 %, b) the composite with 13 % B4C
Slika 1: Tipi~en opti~ni posnetek: a) kompozit z volumenskim dele-
`em B4C 2 %, b) kompozit s 15 % B4C
different numbers of neurons in the hidden layers and the
number of hidden layers were analyzed and presented
graphically. This figure also gives information about the
accuracy of five famous training algorithms depending
on the number of neurons in the hidden layers and the
number of hidden layers. It is evident from this figure
that the smallest error value was obtained by using the
Levenberg-Marquardt training algorithm with two
hidden layers and eight neurons (MSE = 6.4). BFGS
quasi-Newton back propagation with three hidden layers
and nine neurons in the hidden layers follows the
M. O. SHABANI, A. MAZAHERY: THE PERFORMANCE OF VARIOUS ARTIFICIAL NEURONS INTERCONNECTIONS ...
112 Materiali in tehnologije / Materials and technology 46 (2012) 2, 109–113
Figure 6: Evaluation of the training performance of the networks for different training algorithms according to the MSE values with: a) one
hidden layer, b) two hidden layers, c) three hidden layers and d) four hidden layers
Slika 6: Ocena parametrov treninga mre`e za razli~ne treningalgoritme po MSE-vrednostih za: a) eno skrito plast, b) dve skriti plasti, c) tri skrite
plasti in d) {tiri skrite plasti
Figure 5: Schematic representation of the neural network architecture
Slika 5: Shemati~en prikaz nevronske arhitekture
Figure 7: Comparison between the experimental and predicted values:
a) elongation percentage, b) UTS
Slika 7: Primerjava med eksperimentalnimi in predvidenimi vrednost-
mi za: a) raztezek in b) natezno trdnost
Levenberg-Marquardt algorithm (MSE = 8.1), and thirdly
the gradient descent back propagation including four
hidden layers and six neurons in the hidden layers has
clearly much more error than the previous two cases
(MSE = 14.4). The most error was obtained from the
Resilient back-propagation training algorithm and the
Random order incremental training with learning
functions. The Levenberg-Marquardt training algorithm
was found to be the fastest training algorithm; however,
it requires more memory with the same error conver-
gence bound compared to the training methods25. MSE is
a good criterion to have information about learning
performance. The iterations were continued until it is
decided that the minimum MSE error is obtained.
Figure 7 shows the efficacy of the optimization
scheme by comparing the ANN results with the
experimental values. There is a convincing agreement
between the experimental values and the predicted
values for UTS and the elongation percentage of the
A356 composite reinforced with B4C particulates for the
Levenberg-Marquardt training algorithm.
4 CONCLUSION
1) The mechanical properties modeling was developed
to predict the hardness, yield stress, ultimate tensile
strength and elongation percentage.
2) The effect of various training algorithms on the
prediction of the mechanical properties of the
fabricated A356 composite reinforced with B4C
particulates was investigated. The prediction of the
ANN model was found to be in good agreement with
the experimental data.
3) According to the results, the Levenberg-Marquardt
learning algorithm gave the best prediction for
hardness, yield stress, ultimate tensile strength and
elongation percentage for the A356 composite. It is
believed that an ANN with two hidden layers and
eight neurons (MSE = 6.4) gave an accurate
prediction of the mechanical properties of the
fabricated A356 composite reinforced with B4C
particulates.
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Materiali in tehnologije / Materials and technology 46 (2012) 2, 109–113 113
M. O. SHABANI, A. MAZAHERY: THE PERFORMANCE OF VARIOUS ARTIFICIAL NEURONS INTERCONNECTIONS ...

S. PRVULOVI] et al.: EXPERIMENTAL AND THEORETICAL INVESTIGATION OF DRYING TECHNOLOGY ...
EXPERIMENTAL AND THEORETICAL INVESTIGATION
OF DRYING TECHNOLOGY AND HEAT TRANSFER ON
THE CONTACT CYLINDRICAL DRYER
EKSPERIMENTALNA IN TEORETI^NA RAZISKAVA
TEHNOLOGIJE SU[ENJA IN PREVAJANJA TOPLOTE NA
KONTAKTNEM VALJASTEM SU[ILNIKU
Slavica Prvulovi}, Dragi{a Tolma~, Miroslav Lambi}, Dragana Dimitrijevi},
Jasna Tolma~
University of Novi Sad, Technical faculty "Mihajlo Pupin", Djure Djakovi}a bb, 23000 Zrenjanin, Serbia
prvulovicslavica@yahoo.com
Prejem rokopisa – received: 2011-06-05; sprejem za objavo – accepted for publication: 2011-11-23
An experimental and theoretical research on the technology and application of contact drying on the rotating cylinder dryer is
presented. Results of measurements of drying parameters of drying starch solution in real working conditions of drying are
analysed. Based on tests numerical values of the coefficient of heat transfer, heat transfer model, energy performance, curves of
drying kinetics and drying kinetics equations are given, characteristic for drying technology of drying and other relevant
parameters of the process.
Keywords: drying technology, heat transfer, contact cylindrical dryers
Predstavljena je eksperimentalna in teoreti~na raziskava tehnologije in uporabe kontaktnega su{enja na su{ilniku z vrte~im se
valjem. Analizirani so rezultati meritev parametrov su{enja raztopine {kroba v realnih obratovalnih razmerah. Na podlagi
numeri~nih vrednosti koeficientov prenosa toplote so prikazani model prenosa toplote, poraba energije, krivulje kinetike su{enja
in kineti~ne ena~be su{enja, karakteristi~ne za tehnologijo su{enja in drugi za proces pomembni parametri.
Klju~ne besede: tehnologija su{enja, prenos toplote, kontaktni valjasti su{ilnik
1 INTRODUCTION
Drying on the rollers is a method recognized world-
wide as a continuous and very economical technology.
Contact roll kiln is applied in several branches of
industry, especially in chemical and food industry. It is
also applied in drying powdered products in water
dispersion with 35–40 % dry matter, colloidal solution
and suspension, viscous liquids and pastes.1 Different
dispersions of certain powdered products unequally react
by drying, which depends on the properties of the
processed powdered material. Thus, it is very difficult to
propose a unique technique of cylinder dryer.2 In
comparison to other systems, roller drying accommo-
dation requires less space, service and maintenance are
very simple and performed with a small labor force.
These advantages, as well as the economical transfer of
heat provide a low price of dried final product.
The drying time is in most cases of only a few
seconds and it is very important by drying of products
sensitive to heat, such as vitamins, which makes better
high temperature in short period of time than lower
temperatures in long period of time. For substances such
as starch solutions in water, contact drying on heated
rollers is a good solution of the problem of drying.
Adhesion of dried solution is intensive (in the thin
layer) and the drying is done in one incomplete roll
revolution. The layer thickness of dried material is
regulated by the size of gap between the main and
applying rollers, which usually ranges from 0.1–1.0 mm.
Thanks to the principle of drying based on direct contact
of heated roll and wet material, intensive exchange of
heat and mass occurs.
The kiln shown in relation to other technological
solutions, has better technical and economic indicators of
work. In addition to work efficiency, the essential
precondition for any wider application of roll drying is
the relatively low specific energy consumption. This
consumption by the drying roller is significantly less
than by using of other drying mechanisms and spray
drying, or in any other way. Generally it is in the range
of 1.2–1.6 (kg steam/kg evaporated water). The
revolution rate is in the range 5–25 r/min and it depends
on the type of dried material. For roll heating used
aerated water pressure 3–8 bar, hot water or organic
liquids with high boiling temperature are used.3,4
The drying efficiency is estimated with the quantity
of evaporated moisture, which is in range 10–60 kg h–1
m–2 depending on the size of the drying cilinder. The
relatively simple construction and low specific energy
consumption make these drying roll devices very
attractive for applications in industry. For these contact
roll dryers there are very few technical data on experi-
mental plants which would enable exact calculations
Materiali in tehnologije / Materials and technology 46 (2012) 2, 115–121 115
UDK 536.2 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)115(2012)
necessary for designing of dryers. Drying on cylinder
dryers is an improved drying technique. It provides high
quality of dried material and high efficiency and
economy of plants, i.e. the reduction of the power and
investment costs.5 Descriptions of cylinder dryers and
other drying systems are found in references.6,7
2 EXPERIMENTAL FACILITY AND RESULTS
OF MEASUREMENTS
The tests were performed on the industrial dryer with
cylinder diameter d = 1 220 mm and length L = 3 048
mm, heated inside by steam vapor shown in Figure 1.
By cylinder heating and by constant working pressure of
p = 4 bar, the stationary conditions necessary in
experimental measurements are obtained for a great
number of rotations.
Used measuring instruments:
1) Water vapor pressure
pp = 4 bar
2) Water vapor temperature
Tp = 140 °C
3) Number of cylinder rotations
n = 7.5 min–1
4) Thickness of cylinder envelope
1 = 35 mm
5) Thickness of the dried material moisture
2 = 0.25 mm
6) Cylinder surface A = 11.5 m2
7) Water content of dried material
– start of drying w1 = 65 %
– end of drying w2 = 5 %
8) Water vapor consumption mp = 268 kg h–1
9) The dried material is 35 % mass solution of potato
starch in water.
For experimental measurements, the next measuring
instruments and accessories were used:
1) For measuring of water pressure: membrane
manometer of 0–10) bar range and precision of 1.6
%;
2) Water temperature: bimetal thermometer with range
0–200 °C and precision of ±2.5 %;
3) Cylinder surface temperature and temperatures in
direct vicinity of the cylinder: digital thermometer,
type KD-23; with thermo couple NiCr-Ni as sensor,
with range –69–199.9 °C and ±0.1 °C precision.
4) Ambient temperature: glass mercury thermometers
with range 0–50 °C.
5) Air speed in direct vicinity of the cylinder:
anemometer with incandescent wire, type TA 400,
Airflow Developments Canada – LTD, with range
0–2) m/s and precision of ±0.02 m s–1.
6) Moisture of drying material: digital meter for
humidity, type Metler-LP16 and precision of ±0.1 %.
S. PRVULOVI] et al.: EXPERIMENTAL AND THEORETICAL INVESTIGATION OF DRYING TECHNOLOGY ...
116 Materiali in tehnologije / Materials and technology 46 (2012) 2, 115–121
Figure 1: Scheme of experimental contact rool dryer; 1- cylinder; 2-
bringing cylinders; 3- scattering cylinder; 4- knife; 5- pipeline for the
wet material transporting; 6- worm conveyer; 7- steam pipeline; 8-
scheme of measuring points
Slika 1. Shema eksperimentalnega kontaktnega valjastega su{ilnika;
1- valj; 2- dovodna valja; 3- razpr{ilni valj; 4- no`; 5- cevovod za
transport su{enega materiala; 6- pol`asti prenosnik; 7- parni cevovod;
8- shema merilnih to~k
Table 1: Average values of the results of temperature measuring of
drying material Tm/°C and content of moisture, w/%
Tabela 1: Povpre~ne vrednosti izmerjenih temperatur su{enega mate-
riala Tm/°C in vsebnost vlage v masnih dele`ih, w/%
Measuring
place,
according to
the Figure 1
Average values of
the temperature
dried material,
Tm/°C
Moisture of
the dried
material,
w/%
Time drying
t/s
3 – 65 0
4 80 54 1
5 81 41 2
6 82 29.5 3
7 84 18.5 4
8 89 10 5
1 96 5 6
Table 2: Results of temperature measuring in drying material layer on
the cylinder surface, T/°C
Tabela 2: Povpre~ne temperature v plasti su{enega materiala na
povr{ini valja, T/°C
Number of
measure-
ment points
Distance from cylinder; x/m
0 0.005 0.01 0.02 0.03 0.04
1. 96 65 43.5 39.8 37.2 35.0
2. 98 59 40 38 37 35.5
3. – – – – – –
4. 80 53 42.0 40.0 31.0 30.0
5. 81 48 35.0 31.0 30.0 29.0
6. 82 45 36.0 28.5 38.0 27.0
7. 84 43 30.0 27.3 26.0 25.0
Mean value
T/°C 85.0 50.4 37.8 34.2 31.6 30,2
7) Water power consumption: measurement of conden-
sate mass out of the dryer with the balance "Scalar
100" with the range of 0–100 kg.
The results of temperature measuring with dried
materials layer on cylinder surface, and dried material
moisture are given in Table 1. The measuring was per-
formed in the plane of cylinder cross-section according
to experimental points marked in Figure 1.
The results of temperature measuring with dried
materials layer on cylinder surface are given in Table 2.
3 DETERMINATION OF HEAT TRANSFER
COEFFICIENT
The overall heat flux from vapor into surrounding air
can be calculated as:
qu = U (Tp – T) (1)
For great cylinder diameters and with relation to
envelope thickness, it is possible with great accuracy use
the term for the coefficient of heat transfer as for the flat
wall the Eq. (2). The difference of heat transfer
coefficient for flat wall and for the cylinder diameter d =
1 220 mm and cylinder wall thickness 1 = 35 mm is of
1.66 % with regard heat transfer coefficient for cylinder
body. Because of simpler form, for calculation of the
total coefficient of heat transfer Eq. (2) will be applied8.
When in the cylinder surface is covered by a layer of
drying material, the overall heat transfer coefficient is
defined according to equation (2):
U
l
h k k h
=
+ + +
1
1
1
1
1
2 
2m com
(2)
The influential parameter of the mechanism of heat
transfer is the combined coefficient of heat transfer (hm),
Table 3 and value of Nussle’s number is defined with the
equation.
Nu
h d
k
BRe= =c
a
c (3)
On the basis of grouping influential parameters that
influence the most coefficient of heat transfer, the results
of experimental and theoretical researches are being
correlated using the equation of Nussle’s type9–11.
h
k
d
B
dG
c
a
c
=
⎛
⎝
⎜
⎞
⎠
⎟


(4)
The constants (B) and (c), are defined by the method
of the least squares.
4 RESULTS AND DISCUSSION
Applying correlation theory on experimental results
of measuring empirical equations of drying kinetics were
dedeuced.12,13
In the initial drying period the dependence of
moisture versus drying time is almost linear with initial
of drying at t = 0–1) s (Figure 2). Thus, in the initial
period the drying rate is constant. In the second period of
drying in temporal interval t = 1–6 s, the drying depen-
dence changes to a second rank polio. At the end of
drying, the content of moisture is of w2 = 5 %.
In Figure 3, the drying rates curves are presented.
Initially, the drying rate is constant, and in the second
period the drying rate decreases. When the content of
moisture is reduced to that of balanced moisture w = 5 %,
the evaporation rate is dw/dt = 0.06 (kg water/ kg dried
solid).
The presented thermal drying curve in Figure 4
corresponds to a dependence of polio of second rank.
The initial drying temperature is of 80 °C and in the end
of drying it is of 96 °C.
Using the method of smallest squares in processing
the experiment data the next empiric equations were
derived:
S. PRVULOVI] et al.: EXPERIMENTAL AND THEORETICAL INVESTIGATION OF DRYING TECHNOLOGY ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 115–121 117
Figure 2: Dependence content of moisture and drying time
Slika 2: Odvisnost vsebnosti vlage od ~asa su{enja
Figure 3: Dependence quantity of dried solid and drying time
Slika 3: Odvisnost koli~ine trdnega materiala od ~asa su{enja
– dependence content of moisture material and drying
time (Figure 2):
w = 66.166 – 14.196 t + 0.636 t2 (5)
– dependence drying rate and drying time (Figure 3):
dw/dt = 0.744 – 0.168 t + 0.008 t2 (6)
– drying temperature and drying time (Figure 4):
Tm = 82.40 – 2.721 t + 0.821 t
2 (7)
These empiric equations for drying are obtained from
experimental data, they define the character of the drying
process and are in accord with previous researches.14,15
In the initial period of drying, the surface of dried
material with high content of moisture is covered by a
thin layer of water, it behaves as free moisture and the
evaporation is accelerated also with taking up physically
tied moisture (Figure 3). In the second period, the drying
rate is lower, also for tied moisture.
Applying the correlation theory for experimental
results we obtained empirical equations for change of
temperature (T) in function of distance (x), for distance
of cylinder surface in every measuring point. The
temperature in the plane of the central cross-section of
cylinder with consideration of standard deviations is
given in Figure 5.
The empiric equation for the dependence of mean
temperature and distance x (m), from the cylinder
surface (Figure 5), is:
T = 75.50 – 3 611.88x + 64 595.87x2 (8)
Based on the results of experimental and theoretical
investigation16,17 the following correlative equations were
deduced:
Nu = 0.569 Re
0.691 (9)
hc = 0.569 Re
0691 (k/d) (10)
qm = – 3.29 (dT/dx)x = 0 (11)
In the layer of drying material, the total heat flux
consists of part of flux equal to the product of heat
conductivity of humid material and temperature gradient
and of the flux part equal to the product of material flux
of humidity, specifically the humidity enthalpy, i.e. the
flux originating from evaporation of humidity. The
intensity of this flux is a relevant factor in total heat flux
by drying the material on the surface.
On the basis of local temperature, the heat flux is
variable along the rim of the rotating cylinder. In the
second drying period, and especially at the end of
drying, the temperature gradient has a rising tendency.
During the drying process on cylinder dryers, humidity
remnants near the end of drying are evaporated at rising
temperature on material surface and cause higher
variables of temperature gradient.
S. PRVULOVI] et al.: EXPERIMENTAL AND THEORETICAL INVESTIGATION OF DRYING TECHNOLOGY ...
118 Materiali in tehnologije / Materials and technology 46 (2012) 2, 115–121
Table 3: Combined heat transfer coefficient (hcom), heat transfer
coefficient with convection (hc), heat transfer coefficient with
radiation (hr) and heat transfer coefficient by evaporation of humidity
(hw)
Tabela 3: Kombinirani koeficient prenosa toplote (hcom), koeficient
prenosa toplote s konvekcijo (hc), koeficient prenosa toplote s seva-
njem (hr) in koeficient prenosa toplote z izparevenjem vlage (hw)
Number
of
measur-
ing
place
Convective
heat transfer
coefficient
hc /
(W m–2 K–1)
Radiation heat
transfer
coefficient
hr /
(W m–2 K–1)
Evaporation
heat transfer
coefficient
hw /
(W m–2K–1)
Combined
coefficient of
heat transfer
hcom /
(W m–2K–1) =
hc + hr + hw
4 15.0 7.0 475 497
5 15.8 7.2 335 358
6 17.0 7.1 189 214
7 17.8 7.2 128 153
8 14.7 7.3 87 109
1 16.9 7.1 41 65
Mean
value 15.8 7.2 210 233
Figure 5: Dependence of temperature in the plane of central cross and
distance from cylinder surface
Slika 5: Odvisnost temperature v ravnini srednjega prereza od razdalje
do povr{ine valja
Figure 4: Dependence temperature of drying layer and time
Slika 4: Odvisnost temperature v su{eni plasti od ~asa su{enja
In Table 3, are given the results of deduction of heat
transfer coefficient with convection (hc), heat transfer
coefficient with radiation (hr), heat transfer coefficient
with evaporation of humidity (hw) and combined heat
transfer coefficient (hcom). The heat transfer coefficient
with convection from drying material layer in air is
variable along the cylinder rim.
The mean value of heat transfer coefficient is 15,8
W m–2 K–1. The maximal value of heat transfer coefficient
found in the lower zone of cylinder is 17.8 W m–2 K–1. To
greater values of Reynolds’s number, correspond the
higher temperature gradient, greater values of heat
transfer with convection and higher Nussle’s number
(Figure 6).
The investigated drying process is in reality a natural
air streaming around a rotating cilinder with low
streaming speed measured at eight measuring points in
close aproximity of the cilinde. The air streaming com-
bines natural flux and flux due to the cylinder rotation
with rotation rate set for an optimal drying. Since the
Raynolds’s number depends on air streaming speed, it
changes in a given interval, presented on (Figure 6).
According to Figure 6, the Reynolds’s number is
lower than Rek = 5 · 105 and indicates to a laminar con-
vection in direct vicinity of the cylinder surface.
The thermal resistance consisting of combined
coefficients of heat transfer (1/hcom) has an important
effect on the overall heat transfer coefficient (U), as
shown in Table 4.
The mean value for thermal resistance of heat trans-
fer of R = 8.26 · 10–3 m2 K W–1 agrees to the mean value
of overall heat transfer coefficient U = 118 W m–2 K–1,
(Table 4). According to data from,18,19 heat transfer
coefficient is in the range of 105–345 W m–2 K–1.
Taking into account the local values of combined
heat transfer coefficients (hm), (Table 4), the variation
along the cylinder size indicates to changes of technical
resistances of heat transfer from the cylinder to air
(1/hcom) and also originate changes of overall heat
transfer coefficient (U) along the cylinder circumference.
The dominant effect on changes of overall heat
transfer coefficient (U) is due to changes of coefficients
of heat transfer with evaporation of humidity (Table 4).
This effect is represented as thermal resistance of heat
transfer (1/hcom). The research results for these dryers
include various values of Reynolds’s number, which
cover air convection speeds from 0.1 m/s to 1 m/s i.e. Re
= 10000–34500 by standard cylinder size of d = 1 220
mm.
The overall energy of vapour as thermal flux is:
q
m r
Ap
p
=
⋅
(12)
The energy balance is presented in order to check the
acquired results. For the value of temperature gradient
(dT/dx)x=0 = –3 611 Eq. (8), heat flux is qm = 11 880 W m–2,
Eq.(11). The overall energy of vapour as thermal flux is
qp = 13 825 W m–2, Eq. (12).
The difference of both values is 1 945 W m–2, and it
is heat loss. Thermal degree of energy use is 	T = 0.859.
During contact drying there is high degree of heat use
due to direct contact of drying material layer and the
cylinder heated surface.
The evaluation of uncertainty is an ongoing process
consuming time and resources. It consists of:
(a) Uncertainty value is from the surface of cylinder
temperature and temperatures in direct vicinity of the
cylinder and uncertainty of the digital thermometer,
type KD-23; and thermo couple NiCr-Ni.
(b) Uncertainty value is from air speed in direct vicinity
of the cylinder.
Uncertainty of the anemometer with incandescent
wire, type: TA 400, 0–2 m s–1.
Applying the correlation theory to measurement
results we have obtained the empirical Eq. (8), (9), (10),
(11) with a high coefficient of correlation, equation (8):
R = 0.985, eq. (9), (10): R = 0.963, eq. (11): R = 0.978,
therefore, the total uncertainty is of 5–7.8 %. The
uncertainty analysis of the whole work shows that
Materiali in tehnologije / Materials and technology 46 (2012) 2, 115–121 119
S. PRVULOVI] et al.: EXPERIMENTAL AND THEORETICAL INVESTIGATION OF DRYING TECHNOLOGY ...
Figure 6: Dependence of change of Nussle’s and Raynolds’s number
with cylinder (d = 1 220 mm, v = 0.35 m/s, Tm = 85 °C)
Slika 6: Odvisnost spremembe Nusslevega in Raynoldsovega {tevila
pri valju (d = 1 220 mm, v = 0.35 m/s, Tm = 85 °C)
Table 4: Overall heat transfer coefficient (U)
Tabela 4. Splo{ni koeficient prenosa toplote (U)
Measur-
ing
point
Thermal resistance of heat transfer
103 (m2 K W–1)
R =(1/h1 + 1/k1 + 2/k2m + 1/hcom)
Overall heat
transfer
coefficient
U/(W m–2 K–1)
4 0.1 0.76 3.1 2.0 167
5 0.1 0.76 3.1 2.7 150
6 0.1 0.76 3.1 4.6 117
7 0.1 0.76 3.1 6.5 96
8 0.1 0.76 3.1 9.1 76
1 0.1 0.76 3.1 15.3 52
Mean
value 0.1 0.76 3.1 4.3 118
temperature and air speed measurements have a
relatively little influence on the accuracy of results. Thus
it is concluded that the obtained results can be used in
practice.
5 CONCLUSIONS
On the basis of experimental results and their
analysis, the following conclusions are proposed:
• Local values of temperature, heat flux and heat
transfer coefficient are different along the cylinder
rim;
• maximal values of heat flux originate in the upper
cylinder zone (i.e. in initial drying period);
• The values of heat transfer complex coefficient from
the surface of drying material on surrounding air
produce changes of thermal resistances and heat
transfer and cause variations of total heat transfer
coefficient along the cylinder rim. The greatest is the
effect of heat transfer coefficient with humidity
evaporation.
• On the basis of the research results, the mean value of
overall heat transfer coefficient U = 118 W m–2 K–1
was obtained.
• On the basis of experimental and theoretical results
the thermo dynamical analysis of the problem was
performed and temperature gradients, heat flux and
heat transfer coefficients were calculated. In this way,
a new approach is given to the drying theory in the
last fifteen to twenty years;
The obtained results can be used:
• For defining the essential dependences and
parameters of heat transfer with rotating cylinders
heated inside by vapor;
• For the design and development of new drying
cylinders or selection of optimal parameters of heat
transfer.
• The research results can be used because of
experimental data taken at real plant as base. For this
reason, the results can be useful for: researchers,
designers and manufacturers of such and similar
drying systems, as well for educative purposes.
• The determined relevant parameters of heat transfer
have had as objective a more complete energy
description of rotating cylinders for cylinder dryers
and drying and to complement the existing knowl-
edge and explanations of some, so far incompletely
explained phenomena in simpler devices.
6 REFERENCES
1 L. Renshu, W. Weihong, H. Jun, A Study on contact drying with
flexible screen, Journal of Forestry Research, 11 (2000) 1, 51–53
2 M. W. Meshram, V. V. Patil, S. S. Waje, B. N. Thorat, Simultaneous
gelatinization and drying of maize starch in a single-screw extruder,
Drying Technology, 27 (2009) 1, 113–122
3 S. Prvulovic, D. Tolmac, M. Lambic, Z. Blagojevic, Researching
results of contact drying, Energetic Technologies, 5 (2008), 3–6
4 D. Tolmac, M. Lambic, Heat transfer through rotating rol of contact
dryer, Int. Comm. in Heat and Mass Transfer, 24 (1997), 569–573
5 S. Prvulovic, D. Tolmac, M. Lambic, Convection drying in the food
Industry, Agricultural Engineering International the CIGR E journal,
9 (2007), 1–12
6 T. Aihara, W. S. Fu, Y. Suzuki, Numerical analysis of heat and mass
transfer from horizontal cylinders in downward flow of air-water
mist, Journal of Heat Transfer, 112 (1990), 472–478
7 N. L. Yong, W. J. Minkowycz, Heat Transfer characteristics of the
annulus of two coaxial cylinders with one cylinder rotating, Heat and
Mass Transfer, 32 (1989), 711–721
8 D. Tolmac, S. Prvulovic, M. Lambic, The Mathematical model of the
heat transfer for the contact dryer, FME Transactions, 35 (2007),
5–22
9 D. Maillet, A. Degiovanni, R. Pasquetti, Inverse heat conduction
applied to the measurement of heat transfer coefficient on a cylinder,
Journal of Heat Transfer, 113 (1991), 549–557
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boiling heat transfer from a horizontal cylinder to sub cooled liquid,
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11 D. Tolmac, Determination of heat transfer coefficient of contact
dryer’s rotating cylinder, Ph. D. dissertation, University of Novi Sad,
Zrenjanin, Serbia, 1995
12 D. Hez, Comparison of processing economics of different starch
dryers. Journal of Starch/Strake, 36 (1984), 369–373
13 D. Tolmac, M. Lambic, The mathematical model of the temperature
field of the rotating cylinder for the contact dryer, Int. Comm. in
Heat and Mass Transfer, 26 (1999), 579–586
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in a two-pass rotating rectangular duct, Heat and Mas transfer, 40
(2004) 6–7, 467–475
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convective heat transfer in radial rotating rectangular ducts, Journal
of Heat Transfer, 113 (1991), 604–611
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convection micro polar boundary layer about a horizontal permeable
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112 (1990), 504–506
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transfer in a horizontal rotating cyilinder of the contact dryer, Facta
Universitatis, 5 (2007), 47–61
18 H. H. Cho, S. Y. Lee, D. H. Rhee, Effects of cross ribs on heat and
mass transfer in a two-pass rotating duct, Heat and Mass Transfer, 40
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List of symbols
d/m – roll diameter
n/(r/min) – number of roll rotations
p/bar – pressure
T/°C – temperature
q/(W m–2) – heat flux
x/m – distance
Nu – Nuselts number
Re – Reynolds number
A/m2 – the cylinder surface
G/(kg s–1 m–2) – mass speed stream warm air
μ/(kg s–1 m–1) – dynamic viscosity warm air
(dT/dx)/K m–1 (or °C m–1) – temperature gradient
w/% – moisture
S. PRVULOVI] et al.: EXPERIMENTAL AND THEORETICAL INVESTIGATION OF DRYING TECHNOLOGY ...
120 Materiali in tehnologije / Materials and technology 46 (2012) 2, 115–121
Tp/°C – water vapor temperature for cylinder heating
r/(kJ kg –1) – heat evaporation steam water
U/(W m –2 K–1) – overall heat transfer coefficient from
condensing vapor in cylinder interior on surrounding
air
ka/(W m–1 K–1) – termical conductivity of air
h1/(W m–2 K–1) – coefficient of heat transfer from con-
densing vapor on cylinder wall
1/m – thickness of cylinder envelope
2/m – mean thickness of drying material layer
k1/(W m–1 K–1) – thermo conductivity of cylinder enve-
lope
k2m/(W m–1 K–1) – mean thermo conductivity of material
at drying
hcom/(W m–2 K–1) – combined coefficient of heat transfer
R/(m2 K W–1) – thermical resistance of heat transfer
S. PRVULOVI] et al.: EXPERIMENTAL AND THEORETICAL INVESTIGATION OF DRYING TECHNOLOGY ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 115–121 121

D. P. JEVREMOVI] et al.: AN RE/RM APPROACH TO THE DESIGN AND MANUFACTURE ...
AN RE/RM APPROACH TO THE DESIGN AND
MANUFACTURE OF REMOVABLE PARTIAL DENTURES
WITH A BIOCOMPATIBILITY ANALYSIS OF THE F75
Co-Cr SLM ALLOY
RE/RM-PRIBLI@EK, NA^RTOVANJE IN IZDELAVA SNEMLJIVIH
DELOV ZOBOVJA Z ANALIZO BIOKOMPATIBILNOSTI ZLITINE
F75 Co-Cr SLM
Danimir P. Jevremovi}1, Tatjana M. Pu{kar2, Igor Budak3, Djordje Vukeli}3,
Vesna Koji}4, Dominic Eggbeer5, Robert J. Williams5
1Clinic for Prosthodontics, School of Dentistry, Pan~evo, University Business Academy, Novi Sad, Serbia
2Clinic for Prosthodontics, Medical Faculty – Department of dentistry, University of Novi Sad, Serbia
3Faculty of Technical Sciences, University of Novi Sad, Serbia
4Oncology Institute of Vojvodina, Sremska Kamenica, Serbia
5Centre for Dental Technology and the National Centre for Product Design and Development Research,
Cardiff Metropolitan University, Cardiff, United Kingdom
budaki@uns.ac.rs
Prejem rokopisa – received: 2011-07-14; sprejem za objavo – accepted for publication: 2011-08-24
The implementation of computer-aided technologies and systems has paved the way towards a significant advancement of the
conventional modelling and manufacture of dental replacements. In this research the focus is on approach that combines reverse
engineering as a modelling technique and rapid manufacture, i.e., selective laser melting, as the manufacturing technology, with
special emphases on material selection in the fabrication of removable partial dentures. The paper presents the results of a
biocompatibility analysis of the F75 Co-Cr dental alloy using the MTT eluate test.
Keywords: reverse engineering, rapid manufacturing, selective laser melting, removable partial dentures, biocompatibility,
Co-Cr alloys, MTT eluate test
Uporaba ra~unalni{kih tehnologij in sistemov je odprla pot do pomembnega napredka konvencionalnega modeliranja in izdelave
snemljivih zobnih nadomestkov. Te`i{~e te raziskave je pribli`ek, ki kombinira obratno in`enirstvo kot tehniko modeliranja in
hitre izdelave z laserskim taljenjem kot izdelavno tehnologijo s posebnim poudarkom na izbiri materiala za snemljive dele
zobovja. V ~lanku so predstavljeni rezultati analize biokompatibilnosti zobne zlitine F75 Co-Cr z MTT-preizkusom eluiranja.
Klju~ne besede: obratno in`enirstvo, hitra izdelava, selektivno lasersko taljenje, snemljivi deli zobovja, biokompatibilnost,
Co-Cr zlitina, MTT-preizkus eluiranja
1 INTRODUCTION
Dental prosthetics (also known as prosthodontics) has
always maintained close relationships with engineering
disciplines, relying mostly on production engineering.
The rapid development of computer-aided (CA) techno-
logies, which completely transformed production engi-
neering, also left an indelible mark on dental prosthetics.
Striving towards its primary goal – primum non nocere
(in English, ’Above all, do not harm!’), the area of dental
prosthetics has introduced numerous novel technologies
and methods that allow the manufacture of precision,
custom-made, optimal dental replacements. During the
past decade, efforts have been concentrated towards an
advancement of the modelling and manufacture of dental
replacements by introducing modern CA systems,
state-of-the-art materials and machining technologies, as
opposed to the traditional way of manual manufacture,
which is prone to numerous subjective errors.1 Amongst
modern CA systems that have found broad application in
this area, the most widely used are 3D-digitization
systems, CAD and reverse engineering (RE), CAE,
CAM, rapid manufacture (RM) (or additive manufacture
that become the adopted term in the sector) and rapid
prototyping (RP). The development and implementation
of such technologies and systems have paved the way
towards a significant advancement in conventional
modelling, manufacture and the inspection of dental
replacements.1–6
Different dental substrates may have special
requirements related to their modelling and manufacture.
Removable partial dentures (RPDs) represent a special
type of denture, designed for partially edentulous
patients who cannot have a fixed partial denture, i.e., a
bridge. This type of prosthesis is referred to as
removable, as patients can remove and reinsert it when
required without professional help. Traditionally, RPD
frameworks are manufactured through the so-called
lost-wax technique, where a wax pattern burns out in a
preheating unit followed by an immediate casting of the
melted alloy. Though in use for decades, this technique is
sensitive and prone to human-induced errors.1,6
Materiali in tehnologije / Materials and technology 46 (2012) 2, 123–129 123
UDK 577.1:616.314-76/-77 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)123(2012)
In this research the focus is on an approach that
combines RE as a modelling technique, and RM i.e.,
selective laser melting (SLM), as the manufacturing
technology, with a special emphasis on material selection
in the fabrication of RPD frameworks.
The virtual design of dental restorations today almost
always requires the application of RE modelling. RE, a
modelling technique widely used in different engineering
fields, has been increasingly applied in the field of
prosthodontics during the past several years, mainly
because of the rapid development of dental 3D
digitization systems and the corresponding modelling
software2,3. RM is no exception and its results in the field
of prosthodontics significantly depend on RE modelling.
Realizing the benefits of RE and RM, recently there have
been several research works related to the possible use of
these technologies in the design and manufacturing of
RPD frameworks.1,3,7–9
SLM, an RM technique, is based on a layer-wise
material addition that allows the generation of complex
3D parts by selectively melting successive layers of
metal alloy powder on top of each other. As presented by
Eyers and Dotchev in5 and Vandenbroucke and Kruth
in,10 SLM is very suitable for dental and medical
applications, due to the complex geometry of the
produced parts. A pilot study presented by Williams et
al.9 showed that an RPD framework produced by SLM
techniques was comparable to conventional frameworks
in terms of accuracy, quality of fit and function.
However, this conclusion is based on a single study and
much work needs to be completed before a final
conclusion can be reached.
A very important issue in prosthodontics is the
material used, i.e., the alloy.10–13 This issue is even more
important in RPD framework manufacturing as the
application is limited to cobalt-chromium (Co-Cr) alloys
due to the low rigidity of titanium (Ti) alloys, as
described by Aridome et al. in14, and the unfavour-
able characteristics of gold (Au) alloys, that are
otherwise widely used in the fabrication of other dental
restorations. The application of SLM in the manufac-
turing of RPD, brings some additional material
requirements related to mechanical properties10,14 and
biocompatibility.15,16 Although a dental prosthesis
fabricated by SLM showed the potential of SLM as
manufacturing technique, there are still very few reports
of SLM application in the manufacture of RPDs from
Co-Cr alloys. Though Co-Cr alloys have been used in
dentistry for years, little is known about the influence of
the SLM process on the alloys’ biocompatibility and
mechanical behaviour.
This paper, with regards to the above discussion,
focuses on the applicability and possible benefits of the
application of RE and RM techniques in the design and
manufacture of RPD frameworks. Moreover, special
attention in this research has focused on the material
properties related to SLM application in the manufacture
of RPD frameworks. The paper also presents the results
of an analysis of biocompatibility conducted with an
MTT eluate test of the F75 Co-Cr dental alloy.
2 RE/RM IN DESIGN AND MANUFACTURING
OF RPD FRAMEWORKS
The RE and RM combination has been recognized as
fully compatible and very effective. Potential advantages
include: a decrease of manufacturing time, an inherent
repeatability, and the achievement of high quality
through eliminating operator variations that are usually
connected with the conventional (manual) design and
manufacture of an RPD (Figure 1).
The application of RM, i.e., SLM in an RPD frame-
work fabrication implicates the workflow presented in
Figure 2. This workflow clearly presents three main
phases:
1. RE modelling (of the patient’s cast),
2. virtual design of the RPD framework,
3. RM of RPD framework.
The first two phases are frequently unified and
denominated in references as the CAD phase, while the
third is often described as the CAM phase.1,3,5,7–10
The RE phase starts with the 3D digitization of the
patient’s cast. This usually includes acquiring a dental
D. P. JEVREMOVI] et al.: AN RE/RM APPROACH TO THE DESIGN AND MANUFACTURE ...
124 Materiali in tehnologije / Materials and technology 46 (2012) 2, 123–129
Figure 1: Conventional (manual) design of RPD
Slika 1: Konvencionalno (ro~no) na~rtovanje RPD
impression and extra oral scanning of a gypsum model
produced from the impression (Figure 1).1,6 However,
the process could replace the need for an impression by
the application of intra-oral scanning2 or CT.6,17,18 The
point cloud obtained almost always needs to be
pre-processed in order to insure a high-quality surface
reconstruction, i.e., a credible CAD model. Regarding
the applied 3D digitization technique/system, the
pre-processing step can include different processes, such
as noise filtering, data reduction, segmentation of the
point cloud parts or assembling.19 The obtained surface
model (the reference model or the "buck") is usually
exported to an STL file format based on triangular
polygons, which is a suitable format for virtual dental
surveying and virtual sculpting environments.8
The initial step of the RPD framework-design phase
is the virtual dental surveying that is needed to identify
areas of undercut present on the CAD model of the
patient’s teeth and soft tissues (Figure 3a). Unwanted
undercuts have to be removed in order to ensure an
unobstructed withdrawal of the RPD from the patient’s
orifice.1,8 However, there are some useful undercuts that
need to be retained and their identification and
measurement are important as they serve as secure
holders of flexible clasps that provide reliable retention.1
The next step is the modelling of reliefs – the parts of a
model that prevent the RPD framework from resting on
the surfaces of soft tissues (Figure 3b).1 After the reliefs
have been added, virtual sculpting of the RPD frame-
work elements (1-occlusal rest, 2-polymeric retention
frame, 3-lingual bar, 4-acrylic line, 5-non-active clasp,
6-guide plate) may begin (Figure 3c).
The virtual sculpting stage is based on software tools
enabling analogous work to that used in physical
sculpting. This is enabled through a haptic interface that
incorporates positioning in 3D space and allows rotation
and translation in all axes, transferring hand movements
into the virtual environment (Figure 4). Moreover, haptic
systems enable the operator to feel contact with the
object that is the subject of the modelling. Besides this
usage of the haptic arm in a freehand manner, the
process of virtual sculpting also allows the application of
standard CAD parametric features based on sizes,
shapes, relations and positions.1,8
Once the CAD model of the RPD framework is
obtained, it can be passed to the SLM after its
preparation in appropriate software. This preparation
primarily involves the creation of an adequate support
(Figure 5) that acts as a firm base for the RPD
framework to be built onto and which also conducts heat
away during the sintering processes. As the supports
need to be removed after the solidification of the part, it
is important to avoid placing them on the fitting surface
of the RPD.8
D. P. JEVREMOVI] et al.: AN RE/RM APPROACH TO THE DESIGN AND MANUFACTURE ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 123–129 125
Figure 2: The typical workflow of an RE/RM approach in RPD
framework fabrication
Slika 2: Zna~ilen tok RE/RM pribli`ka v okviru RPD-izdelave
Figure 3: RPD framework design phase – the main steps: a) Identification of undercuts, b) Relief modelling, c) RPD framework elements
Slika 3: Okvir faze RPD-na~rtovanja – glavni koraki: a) identifikacija spodnjih prerezov, b) modeliranje reliefa, c) elementi okvirja RPD
During the SLM process, powdered material is
spread by a hopper and wiper mechanism. To
accommodate a new layer of the material, the build
platform has to move down by one layer thickness.
Subsequently, the powder is deposited incrementally on
top of each solidified layer, and the process is repeated
(Figure 6).
The manufactured RPD framework needs to be
finished and polished and this is performed using
traditional dental laboratory procedures.8,9 Finally, the
finished RPD framework has to be evaluated on the
patient’s cast. This is performed by a prosthodontic
expert, who will assess the quality of fit according to
recommended practice.8
3 BIOCOMPATIBILITY TESTING OF THE SLM
Co-Cr ALLOY F75 (BY MTT ELUATE TEST)
One of very important issues that need to be
investigated is the biocompatibility of the specific Co-Cr
alloy used for SLM, since – to the best of authors’
knowledge – there are no known published conclusions
about biocompatibility. Though the basic chemical
elements in alloys used for SLM (F75) and conventional
investment casting, also known as the lost-wax technique
(Remanium 380+) generally match, they differ by a
small percentage due to the specific requirements needed
by the process (Table 1). However, it has been shown
that a modification in composition and pre-treatment can
influence the cytotoxicity of an alloy on a large scale.20,21
The previous discussion motivated the authors to start
research related to the biocompatibility testing of the
specific Co-Cr alloy used for SLM.
In biocompatibility evaluations of alloys, cell culture
studies are the usual starting point as they enable an
investigation of the toxicity in a simplified system that
minimizes the effect of confounding variables.21 Thus,
within this study a murine fibroblast cell line (L929) was
used in accordance with the requirements of the ISO
standard 7405 (ISO 2008).15 The cytotoxicity determi-
nation of the Co–Cr alloy used for the fabrication of an
SLM RPD framework was based on the MTT eluate test
method.
3.1 Sample preparation
The investigation included the fabrication of two
groups of test samples – the first from the SLM and the
second from conventional investment casting (Figure 7).
D. P. JEVREMOVI] et al.: AN RE/RM APPROACH TO THE DESIGN AND MANUFACTURE ...
126 Materiali in tehnologije / Materials and technology 46 (2012) 2, 123–129
Table 1: Composition of the Remanium GM 380+ and Sandvik Osprey F75 alloys in mass fractions, w/%
Tabela 1: Sestave zlitin Remanium GM 380+ in Sandvik Osprey F75 v masnih dele`ih, w/%
Ingredients, w/% Co Cr Mo Si Mn N C Fe Ni
Remanium GM 380+ 64.6 29 4.5 <1 <1 <1 <1 / /
Sandvik Osprey F75 Balance 27–30 5–7 <1 <1 / <0.35 <0.75 <0.5
Figure 6: Principle of the SLM process
Slika 6: Princip SLM-procesa
Figure 4: Virtual design of RPD frameworks
Slika 4: Virtualno na~rtovanje podlag RPD
Figure 5: Support in SLM manufacturing of RPD frameworks
Slika 5: Podlaga pri SLM-izdelavi RPD-podlage
Figure 7: Test samples obtained by SLM (left) and by vacuum casting
(right)
Slika 7: Preizku{anci SLM (levo) in vakuumskega taljenja (desno)
The first group of samples was manufactured by the
SLM system Realiser MTT-Group (Figure 8) and the
software was Magics 9.5, Materialise NV. The Co-Cr
layer thickness was 0.075 mm, the laser’s maximum scan
speed was 300 mm/s, and the beam diameter was
0.150–0.200 mm. The F75 Co-Cr alloy (Sandvik Osprey
Ltd., UK) used in this study is a spherical powder with a
maximum particle size of 0.045 mm (particle size range
0.005–0.045, mean size approx. 0.030 mm). After the
SLM process was completed and specimens’ supports
were removed, the discs were finished by polishing
according to the usual dental laboratory procedure.
The second group of samples – four disc specimens
with a radius of 5 mm and thickness of 1 mm – were
fabricated from a non-precious Co-Cr alloy Remanium
GM380+ (Dentaurum, Ispringen, Germany) containing
no Ni, Be or Fe, and widely used for RPD framework
casting. The discs were obtained from wax patterns,
invested in Rema dynamic (Dentaurum, Ispringen,
Germany) investment, and vacuum casted in a Nautilus
CC (Bego, Bremen, Germany) system (Figure 9). After
casting, the discs were divested and blasted with 100-μm
aluminium oxide particles, and then polished with silicon
carbide papers in the sequence 320, 400, 600, 1200,
1500 and 2000. The final polishing was performed using
oxide pastes.
3.2 MTT eluate testing
The test performed was the MTT (tetrazolium colori-
metric assay) eluate test, widely used for the quantitative
evaluation of cell proliferation and survival.16,20 The
assay depends on the cleavage of tetrazolium salt
3-[4,5–dimethylthiazol–2–yl]-2,5-diphenyltetrazoliumbr
omide (MTT) to purple formazan crystals by mito-
chondrial dehydrogenases in viable cells.21 The assay
detects living but not dead cells and the rate of MTT
reduction to formazan products. It is also dependent on
the degree of cell activation. Therefore, the assay is
suitable for measuring cytotoxicity, proliferation and
activation.16
Eluates of both the CM and SLM disc samples were
prepared. The samples were extracted with 10 mL of
Dulbecco’s modified Eagle’s medium (DMEM) without
serum for 48 h. The extraction was performed in an
atmosphere of 5 % CO2 and 95 % air at 37 °C. All the
extracts were filtered for sterilization and used as a
culture medium for L-929 cells. MTT assays with L-929
cells treated with different eluates were completed after
48 h of incubation. The experiment was repeated twice
(thus eight CM and eight SLM samples were tested in
two independent experiments).
The cells (L929) were cultured in Petri dishes
containing eluates of the CM or SLM alloy discs (Figure
10). They were incubated for (3, 5, 7 and 9) d at 37 °C in
95 % air and 5 % CO2. The control samples contained a
regular culture medium. After the incubation, the cells
were detached using enzymatic digestion and counted in
a counting chamber using trypan blue. Briefly, 5 × 103
cells were seeded to a 96-well plate and cultured for 48 h
at 37 °C in 95 % air and 5 % CO2.
After the incubation period, 20 μL of MTT solution
were added to each well and incubated for another 3 h.
The purple formazan product was dissolved in 100 μL of
D. P. JEVREMOVI] et al.: AN RE/RM APPROACH TO THE DESIGN AND MANUFACTURE ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 123–129 127
Figure 10: MTT eluate testing
Slika 10: MTT preizkus eluiranja
Figure 9: Nautilus CC system for vacuum casting
Slika 9: Nautilusov sistem za vakuumsko taljenje
Figure 8: SLM system – Realiser MTT-Group, UK
Slika 8: SLM-sistem Realiser MTT-grupe, VB
0.04-M hydrochloric acid in isopropanol. The reduced
MTT was then measured spectrophotometrically in a
dual-beam, microtiter plate reader Multiscan MCC/340
at 540 nm with a 690-nm reference. The optical density
values of the experimental groups were divided by the
control and expressed as a percentage of the control.
3.3 Statistical results of the MTT eluate testing
A statistical analysis was carried out using the
Statgraphics Centurion program. The data were
evaluated statistically using the Student’s t-test and a
value of p < 0.05 was considered to be statistically
significant. The results of the MTT eluate testing of the
disc samples that were taken after an extraction period of
(3, 5, 7 and 9) d are listed in Table 2, and graphically
presented in Figure 11.
Of particular interest was the confidence interval
evaluation for the difference between the means. Since
all four intervals contain the value 0, there is no
statistically significant difference between the means of
the CM and SLM samples at the 95 % confidence level.
Furthermore, a t-test was used for testing a specific
hypothesis about the difference between the means of the
populations from which the two samples come. In this
case, the test was constructed to determine whether the
difference between the two means equals 0.0 (null
hypothesis: mean1 = mean2) versus the alternative
hypothesis that the difference does not equal 0.0 (mean1
 mean2). Since the computed P-values are not less than
0.05 in all four cases, the null hypothesis cannot be
rejected.
4 DISCUSSION
Publications investigating the corrosion of dental
alloys give information firstly about the release of
potentially harmful ions from the dental device.
Cell-culture tests give an insight into whether the
released ions in a cell-culture medium imitating the oral
environment could have a negative effect on the biolo-
gical system. The results of the MTT eluate testing with
the SLM samples do not show significant cellular
damage potential. Statistical analyses carried out showed
that the alloys did not release harmful material that could
cause acute effects against L929 cells under the given
experimental conditions. Furthermore, the MTT test
showed no permanent damage to the cell function. The
viability was much higher than 50 % after all the
extraction periods for both the CM and SLM alloy.
Replication during an extended contact period with
D. P. JEVREMOVI] et al.: AN RE/RM APPROACH TO THE DESIGN AND MANUFACTURE ...
128 Materiali in tehnologije / Materials and technology 46 (2012) 2, 123–129
Table 2: Results of statistical analysis of MTT eluate testing of CM and SLM disc samples
Tabela 2: Rezultati statisti~ne analize MTT-preizkusa eluiranja CM- in SLM-vzorcev
Period of cell
incubation in d 3 5 7 9
D
es
cr
ip
tiv
e
st
at
is
ti
cs
Technology CM SLM CM SLM CM SLM CM SLM
Count 8 8 8 8 8 8 8 8
Average % of
relative cell No. 104.655 103.639 107.604 108.364 113.73 112.694 119.778 120.525
Standard
deviation 3.62557 4.56325 3.43234 2.94701 3.34183 2.92113 3.45744 3.58149
Coeff. of
variation,% 3.4643 4.40303 3.18979 2.71955 2.93839 2.5921 2.88655 2.97158
Minimum 100.05 98.64 102.04 101.54 109.78 109.78 114.01 115.61
Maximum 109.45 111.23 111.97 111.02 119.74 117.64 124.31 125.64
C
om
pa
ri
so
n
of
m
ea
ns
95 % CIM* 104.655
+/–
3.03106
103.639
+/–
3.81498
107.604
+/–
2.86951
108.364
+/–
2.46377
113.73
+/–
2.79385
112.694
+/–
2.44213
119.778
+/–
2.8905
120.525
+/–
2.99421
95 % CIDM** 1.01625
+/–
4.41952
–0.76
+/–
3.43048
1.03625
+/–
3.36576
–0.7475
+/–
3.77485
t 0.493186 –0.475165 0.660339 –0.424715
P-value 0.629527 0.641997 0.519754 0.6775
* CIM – Confidence Interval for Means, ** CIDM – Confidence Interval for the difference between the Means assuming equal variances
Figure 11: MTT test results for CM and SLM alloy after different
incubation periods – graphical review of the confidence intervals
Slika 11: Rezultati MTT-preizkusov CM- in SLM-zlitine po razli~nih
dobah inkubacije – grafi~na predstavitev intervalov zanesljivosti
potential toxic substances, however, showed good
biocompatible properties of the chosen SLM alloy.
Additionally, the negative effect decreased with time for
both the examined substances. Therefore, both alloys can
be rated as non-cytotoxic.
It has, however, to be noted that SLM, as a complex
thermo-physical process, produces a variation in the final
product depending on several factors, such as the
material, laser, scan and parameters of the environment
used.10 Changeable variables include: laser power, layer
thickness, scan speed and hatch spacing. With the current
settings, as can be seen from the study, the final product
complies with the required biocompatibility standards,
showing no potentially harmful effect. However, those
values can be adjusted accordingly, optimizing some
aspects that can have a negative effect on the materials’
properties, such as porosity. For example, for a low
energy input, successive scan tracks may not be fully
molten, leaving large pores along the scan lines, as seen
in the mentioned study. If so, a combination other
parameters might change the surface properties, which
might also result in changes in the ion release and
therefore require separate biocompatibility studies. This
study has, however, showed that the initial screening
gave positive results and the F75 SLM alloy can be
subjected to further tests.
The study concerning the ion release from the cast
and SLM samples, presented in,10 revealed the more
favourable behaviour of the SLM specimens. The main
ion detected was cobalt, since the corrosion of the alloy
is determined by the main component, and the
passivating effect of chromium. The SLM test specimens
showed lower emissions than the cast specimens,
probably because the laser-melted material is more
homogeneous, contains fewer pores and has a finer
microstructure. This also highlights the importance of
the finishing procedure, which still has to be conducted
manually.
Another physical factor that might influence the
biocompatibility šin vivo’ is the surface roughness21. The
changes in this parameter can be explained by the
so-called stair effect, inherent to the layer-wise
production of SLM. In the oral environment, this might
increase plaque retention, leading to the formation of
acidic micro-fields that might change the metallic ion
release unfavourably. This effect can be reduced by
decreasing the layer thickness or by increasing the
sloping angle10.
5 CONCLUSION
This paper shows that a complete RE/RM procedure
for RPD framework fabrication should bring significant
advantages both to practitioners and patients. Special
attention was focused on the biocompatibility analysis of
the dental alloys used with the SLM. Moreover, the
paper presents a biocompatibility evaluation of the F75
SLM dental alloy using the MTT eluate test. On the basis
of the obtained results, within the limitations of the
study, it can be concluded that the RE/RM procedure
showed a promising potential in RPD framework
fabrication, as well as that the F75 alloy used for SLM
manufacturing showed positive initial results regarding
its biocompatibility. However, further studies, including
in vivo tests and tests of mechanical properties, have to
be conducted before the final release of the alloy for
mass production.
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D. P. JEVREMOVI] et al.: AN RE/RM APPROACH TO THE DESIGN AND MANUFACTURE ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 123–129 129

P. PE^LIN et al.: INFLUENCE OF TRANSIENT RESPONSE OF PLATINUM ELECTRODE ...
INFLUENCE OF TRANSIENT RESPONSE OF PLATINUM
ELECTRODE ON NEURAL SIGNALS DURING
STIMULATION OF ISOLATED SWINISH LEFT VAGUS
NERVE
VPLIV PREHODNEGA ZNA^AJA PLATINASTE ELEKTRODE NA
@IV^NI SIGNAL MED STIMULACIJO IZOLIRANEGA @IVCA
VAGUSA SVINJE
Polona Pe~lin1, Franci Vode2, Andra` Mehle1, Igor Gre{ovnik3, Janez Rozman1
1ITIS, d. o. o., Ljubljana, Centre for implantable technology and sensors, Lepi pot 11, 1000 Ljubljana, Slovenia
2Institute of metals and technology, Lepi pot 11, 1000 Ljubljana, Slovenia
3Faculty of Natural Sciences and Mathematics, University of Maribor, Koro{ka cesta 160, 2000 Maribor, Slovenia
polona.peclin@gmail.com
Prejem rokopisa – received: 2011-07-29; sprejem za objavo – accepted for publication: 2011-12-16
The main aim of the work was to measure transient response characteristics of interface between platinum stimulating
electrodes and isolated swinish left cervical vagus nerve (segment), when electrical stimulating pulses are applied to preselected
locations along the segment and elicited neural signals, also described as compound action potentials (CAPs), are recorded from
particular compartments of the nerve.
The stimulating system was manufactured as a silicone self-coiling spiral cuff (cuff) with embedded matrix of ninety-nine
rectangular electrodes (0.5 mm in width and 2mm in length), made of 45 μm thick annealed platinum ribbon (99.99 % purity),
and a geometric surface of 1 mm2.
For electrical stimulation, a current quasitrapezoidal, asymmetric and biphasic pulses with frequency of 1 Hz, were used. To test
an influence of stimulating pulses having different parameters and waveforms on elicited CAPs, various degree of imbalance
between an electric charge (charge) injected in cathodic phase as well as charge injected in anodic phase of a biphasic
stimulating pulse, were deployed and compared.
To identify the differences in elicited CAPs however, an integral of the CAP cathodic phase as well as integral of the CAP
anodic phase of stimulating pulse, were calculated and compared.
Results showed a strong component superimposed in the CAPs, considered as an ensemble artefact which greatly obscured the
components of the CAPs, and various components did overlap.
Results also showed that stimulating pulses, having preset certain degree of imbalance between charge injected in cathodic and
charge injected in anodic phase, elicited a slight change in a positive waveform deflection of CAP manifested under a cathodic
phase as well as slight change in a negative waveform deflection of CAP manifested under an anodic phase. Furthermore, slight
difference was observed in a CAP, expressed as integral cathodic positive deflection and as integral anodic negative deflection.
However, it could be concluded that measured CAPs are not greatly influenced by the imbalance between a charge injected in
cathodic and anodic phase of quasitrapezoidal, asymmetric and biphasic stimulating pulses.
Keywords: electrical stimulation, platinum electrodes, left vagus nerve, electrochemistry, electrical charge
Glavni namen dela je bil izmeriti prehodni zna~aj na prehodu med platinasto stimulacijsko elektrodo in delom izoliranega levega
vratnega pra{i~jega `ivca vagusa (segment) med dovajanjem stimulacijskih impulzov na izbrana mesta vzdol` `ivca in hkratnim
merjenjem `iv~nega signala, imenovanega sestavljeni akcijski potencial (CAP) na dolo~enih predelih `ivca.
Stimulacijski sistem je bil izdelan v obliki silikonske spiralne objemke (cuff) z vdelano matriko devetindevetdestih pravokotnih
elektrod ({irina 0,5 mm in dol`ina 2 mm), izdelanih iz `arjenega platinastega traku (~istost 99,99 %) in geometrijsko povr{ino
1 mm2.
Za elektri~no stimulacijo so bili uporabljeni tokovni kvazitrapezni, asimetri~ni in izmeni~ni impulzi frekvence 1 Hz.
Za preizku{anje vpliva stimulacijskih impulzov z razli~nimi parametri in oblikami na izzvani CAP je bila med katodno in
anodno vneseni elektri~ni naboj uvedena dolo~ena stopnja neuravnote`enosti. Katodno in anodno vnesena naboja sta bila za
izbrane impulze med seboj primerjana.
S ciljem ugotavljanja razlik v izzvanem CAP-u pa sta bila izra~unana in med seboj primerjana integrala tako CAP-a, prisotnega
pod katodno fazo, kot CAP-a, prisotnega pod anodno fazo zgoraj omenjenega stimulacijskega impulza.
Rezultati so pokazali mo~no komponento, polo`eno na CAP, ki je sestavljena motnja in ki znatno zamegli ter prekrije
posamezne komponente CAP-a.
Rezultati so tudi pokazali, da zgoraj omenjeni stimulacijski impulzi s prednastavljeno dolo~eno stopnjo neravnote`ja med
nabojem, vnesenim v katodni fazi, ter nabojem, vnesenim v anodni fazi, izzovejo majhne spremembe pri pozitivnem odklonu
CAP-a, prisotnega pod katodno fazo, kakor tudi majhne spremembe v negativnem odklonu CAP-a, prisotnega pod anodno fazo.
Nadalje so bile opa`ene tudi majhne razlike v CAP-ih, izra`ene kot integral pod katodnim pozitivnim odklonom in kot integral
pod anodnim negativnim odklonom.
Kon~no je mogo~e skleniti, da neravnote`je med katodno in anodno vnesenim nabojem ni znatno vplivalo na izmerjene CAP-e.
Klju~ne besede: elektri~na stimulacija, platinaste elektrode, levi `ivec vagus, elektrokemija, elektri~ni naboj
Materiali in tehnologije / Materials and technology 46 (2012) 2, 131–137 131
UDK 621.357:591.18 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)131(2012)
1 INTRODUCTION
In the past few decades, vagus nerve stimulation
(VNS) has been the subject of considerable research with
the goal to be used as method to treat a number of
nervous system disorders, neuropsychiatric disorders,
eating disorders, sleep disorders, cardiac disorders,
endocrine disorders, and pain, among others.1–3 In
practically all studies in humans, VNS refers to
non-selective stimulation of the cranial nerve X, known
as the left vagus nerve, using specific electrode devices
which development was based on different models
provided by various research groups.The frequent result
of non-selective stimulation however, is the occurrence
of undesirable side effects.4–6
Peripheral nerve stimulation however, often requires
the development of electrode systems that stimulate
selectively a certain group of fibers in a nerve trunk
without excitation of other nerve fibers.
However, the long-term use of such electrical
stimulation requires that it is applied selectively and
without tissue injury. Tissue injury and the corrosion of
the stimulating electrode are both associated with high
charge density stimulation.7,8 For this reason, long-term
stimulation of the nervous tissue requires the absence of
irreversible electrochemical reactions such as electrolysis
of water, evolution of chlorine gas or formation of metal
oxides.
For a given electrode, there is a limit to the quantity
of charge that can be injected in either anodic or cathodic
direction with reversible surface processes. This limit
depends upon the parameters of the stimulating wave-
form, the size of the electrode, and its geometry.9–11
Namely, at the electrode-electrolyte interface there are
capacitive mechanisms (charging and discharging of the
electrode double layer, no electron transfer) and Faradaic
mechanisms (chemical oxidation or reduction, reversible
or irreversible).12–14 If the voltage across the electrode-
tissue interface is kept within certain limits, then
chemical reactions can be avoided, and all charge
transfer will occur by the charging and discharging of the
double-layer capacitance. However, in many instances,
the electrode capacitance is not sufficient to store the
charge necessary for the desired excitation without the
electrode voltage reaching levels where reactions could
occur.11,15
The principal approach to control the interface
voltage has been the use of charge-balanced biphasic
stimuli that have two phases that contain equal and
opposite charge. However, even with charge-balanced
stimulating pulses it is possible that the interface voltage
may reach levels where electrochemical reactions can
occur.16,17
Platinum is among commonly used stimulating
electrode materials that are capable of supplying
high-density electrical charge to effectively activate
neural tissue.11 However, stimulation with a high charge
density, pH shifts causing irreversible changes in tissue
proteins, metallic dissolution products, gross hydrogen
and oxygen gas bubbles, and oxidized organic and
inorganic species, could occur18. Therefore, some
platinum toxicity interactions in the body, actually not
conclusively proven in human, could be expected.
Namely, tests on laboratory mammals showed that
soluble platinum compounds are much more toxic than
insoluble ones while solid platinum wire or foil is
considered to be biologically inert. Complexes with
other dangerous metals and chemicals in the human body
such as platinum salts, can cause several health effects,
such as: DNA alterations, cancer, allergic reactions of
the skin and the mucous membrane, damage to organs,
such as intestines, kidneys and bone marrow and hearing
damage.19,20
In the area of Functional Electrical Stimulation (FES)
cuffs have been used in neuroprosthetic applications as
stimulation electrodes as well as electrodes for the
recording of the electroneurogram (ENG) for more than
35 years.21 Twenty years later, this method was used for
the first time in a chronic implantation in human subjects
for the recording of the ENG for the use of feedback
signal in a system for the correction of foot-drop.22
While the time was passing, theoretical considerations
and different models have stimulated and accompanied
the development of cuffs23,24 promoting them as the most
successful biomedical electrodes for selective stimu-
lation of different superficial regions of a peripheral
nerve.25,26 However, the long-term effectiveness and
potential harmful effects of the cuffs on neural tissue in
various applications is still not completely defined.
The present study addressed the mechanisms that
could be involved in the modulation of recruitment
properties of nerve fibres, and thus, in the modulation of
CAP induced by both, the cathodic and anodic charge
injected during selective stimulation of the isolated
vagus nerve with developed cuffs and imbalanced
quasitrapezoidal stimulating pulses having different
parameters.
2 METHODS
The cuff was designed taking into consideration the
results of histological examination of the swinish left
vagus nerve, the model of selective electrical stimulation
of particular superficial regions of the nerve and the
model of selective stimulation of nerve fibers with
different diameters.27–29 The cuff and physical dimen-
sions of the cuff were actually devised so to induce as
low as possible radial pressure when installed on the
nerve. Therefore, minimum mechanically induced nerve
damage might be expected.
The cuff was manufactured by bonding two 0.05 mm
thick silicone sheets together (Medical Grade Silicone
Sheeting, Non-Reinforced, 6” × 8” × 0.002” Matt,
SH-20001-002, BioPlexus Corporation, 1547 Los Ange-
P. PE^LIN et al.: INFLUENCE OF TRANSIENT RESPONSE OF PLATINUM ELECTRODE ...
132 Materiali in tehnologije / Materials and technology 46 (2012) 2, 131–137
les Avenue #107, Ventura, California 93004. USA.). At
room temperature one sheet, stretched and fixed in that
position, was covered with a layer of adhesive (RTV
Adhesive, Acetoxy, Implant Grade, Part Number 40064,
Applied Silicone Corporation, 270 Quail Court, Santa
Paula, CA 93060, USA). A second un-stretched sheet
was placed on top of the adhesive and the composite was
compressed to a thickness of 0.15 mm until the whole
curing process was completed. In normal laboratory
conditions, the curing process was completed within 24
h. When released, the composite curled into a spiral tube
as the stretched sheet contracted to its natural length.26,27
As a result, the composite is soft self-sizing and flexible
self-coiling spiral tube. When instaled on the nerve, the
cuff wraps around the nerve and, because of its
self-coiling property, adjusts automatically its inner
diameter to the size of the nerve.
Ninety-nine rectangular electrodes with a width of
0.5 mm and length of 2 mm (geometric surface g = 1
mm2, real surface area  1.4 mm13, made of 45 μm thick
annealed platinum ribbon (99.99 % purity), were then
under microscope mechanically mounted on the third
silicone sheet with a thickness of 0.05 mm. They were
arranged in nine parallel groups each containing eleven
electrodes, thus forming a matrix of ninety-nine elec-
trodes.28,29
Afterwards, the electrodes were connected indivi-
dually to the high frequency miniature and highly
flexible isolated, multi-stranded and enameled finest
copper wires (CU-lackdraht DIN 46 435,  12 × 0.04
mm, Elektrisola, Reichshof-Eckenhagen, Germany). For
experimental purpose, the junctions between platinum
electrodes and multi-stranded wires were implemented
using a special tin alloy. The multi-stranded wire was
used, since it has the same average fatigue life as their
individual constituent strands but the variance of that life
is smaller. To maximize service life, it was concluded
that wire strands should be manufactured at the smallest
diameter possible (without introducing structural flaws).
It was assumed that multi-stranded wires to the
stimulating electrodes if routed carefully would play a
minimal role on rotation of the cuff around the
logitudinal axis and on translation in a longitudinal
direction. Therefore, to ensure that the multi-stranded
wires would not be the possible source of mechanical
damage to the nerve, special care should be taken during
instalation to route them so that enough slack would be
left to avoid mechanical tensions being transmitted to the
cuff. Afterwards, a self-coiling tube was mechanically
opened and the silicone sheet with the matrix of
electrodes was adhered onto an inner side of the tube.
In fabricted cuff, when the matrix was spirally rolled
up, the longitudinal separation between nine parallel
groups of electrodes was 2 mm and the circumferential
separation between electrodes was 0.5 mm.
The dimensions of the nerve considered in cuff
design were the:
d – nominal diameter of the nerve: 2.5 mm
c – circumference of the nerve: 7.85 mm
l – total length of the cuff: 38 mm
w – approximate width of opened cuff: 12 mm
Figure 1 shows a finished ninety-nine-electrode cuff
of 44 mm in total length and 2.5 mm in diameter (inner
diameter of the first layer) and had 2.25–2.75 turns,
snugly fitting the nerve in its resting position.
From the electrochemical point of view, a very
important factor considered in the design of the cuff
from different metals, was the stability of electroche-
mical potentials and the galvanometric behavior of
stimulating electrodes in physiological media.
To supervise the electrode-electrolyte interface, the
parameters of the stimulating waveform, injected via the
stimulating electrode, were interchanged so to inten-
tionally exceed the limits for reversible charge injec-
tion.9,10,18
P. PE^LIN et al.: INFLUENCE OF TRANSIENT RESPONSE OF PLATINUM ELECTRODE ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 131–137 133
Figure 1: A perspective illustration of the the nerve segment (a)
instaled into the 99-electrode cuff (b) and a position of specific
electrodes within an arbitrary chosen longitudinal row of platinum
electrodes (c). A-C-A represents triplets of electrodes within the
stimulating section, B-represents blocking electrodes and R–R
represents bipolar couples of electrodes for measurement of a CAP.
Slika 1: Prostorska risba segmenta `ivca (a) vstavljenega v
99-elektrodno objemko (b) in polo`aj posamezne platinaste elektrode
v naklju~no izbrani vzdol`ni vrsti elektrod (c). A-C-A je troj~ek
elektrod v stimulacijski sekciji, B sta "bloking" elektrodi in R–R je
bipolarni par elektrod v sekciji za merjenje CAP-a.
Figure 2: Parameters and waveform of the stimulating pulse
Slika 2: Parametri in oblika stimulacijskega impulza
Precisely, the absence or presence of both, reversible
as well as irreversible electrochemical reactions, was
controled exclusively via the imbalance between
cathodically and anodically injected charge by the
biphasic stimulating waveform.7
The stimulating pulse used in the study and shown in
Figure 2, was current, biphasic, charge balanced and
asymmetric pulse consisting of a precisely determined
quasi-trapezoidal cathodic phase with a square leading
edge with intensity ic, a plateau with width tc and
exponentially decaying phase texp, followed by a wide
rectangular anodic phase ta/μs of a magnitude ia.
To stimulate a determined group fibres within a
particular compartment of a segment, stimulating pulses
at frequency of 1 Hz were applied via stimulating
cathode to preselected location.
An isolated, about 8 cm long segment of a swinish
mid-cervical left vagus nerve, was installed within the
cuff and mounted into the experimental chamber (Figure
3), according to the protocol approved by the ethics
committee at the Veterinary Administration of the
Republic of Slovenia, Ministry of Agriculture, Forestry
and Food (VARS).
To maintain simulated physiological thermal con-
ditions, the body of a measuring chamber machined out
from Plexiglas, was heated to 37 °C using precision
water circulator with range of control: ±0.003 °C .
To prevent extensive drying of the segment and
maintain a natural "wet" surrounding of the nerve, the
segment was occasionally flooded by the simulated
cerebral perfusion fluid consisting of (in mM): MgCI2 2,
CaCI2 2, KCI 2.5, NaCl 126, glucose 10,
NaH2PO4·H2O 1.25, NaHCO3 26. At the same time, an
interface between the stimulating electrode and neural
tissue was maintained to closely mimic the physiological
conditions.
The experiment consisted of four tests referred to as
Test1-4, where given quasitrapezoidal stimulating pulses
with preset parameters and waveform were delivered
from a single channel precision custom designed
stimulator to the triplet 5 in the stimulating section
(ACA) of the cuff. A specific waveform and parameters
of the individual pulses (1 Hz), were chosen by manual
manipulation of the dials on the stimulator.
In the Test 1, the parameters and waveform were the
following: ic = 1.84 mA, tc = 185 μs,
texp = 100 μs, exp = 35 μs and ia = 0.79 mA.
In the Test 2, the parameters and waveform were the
following: ic = 1.71 mA, tc = 65 μs,
texp = 100 μs, exp = 35 μs and ia = 0.74 mA.
In the Test 3, the parameters and waveform were the
following: ic = 4.07 mA, tc = 155 μs,
texp = 105 μs, exp = 60 μs and ia = 1.77 mA.
In the Test 4, the parameters and waveform were the
following: ic = 3.9 mA, tc = 265 μs,
texp = 105 μs, exp = 35 μs and ia = 1.67 mA.
In the first three out of four stimulation tests, the
CAP was measured simultaneously from the right end of
the segment with with the couple of electrodes in the
recording section (R–R) of the cuff, having the same
longitudinal position as appointed triplet5 within the
stimulating section (A-C-A).30,31 In the Test 4 however, a
selected recording couple was located at circumfe-
rentially opposite site according to an appointed triplet5.
In measurements of CAPs, signals recorded with an
appointed couple of electrodes were delivered to a
custom designed differential amplifier and amplified (A
= 100).
The analogous signals of both, stimulating pulses and
measured CAP signals, were digitized via an analogue-
digital conversion board (DEWE-43, high performance
data acquisition system designed and manufactured by
the company DEWESOFT using data acquisition
software DEWESoft 7.0.2 and stored on a Lenovo T420
portable computer).
To supervise the electrode-electrolyte interface
established upon intentionally exceeded limits for
reversible charge injection by different values of para-
meters and waveforms of selectively delivered stimu-
lating pulses, a charge Qc injected in cathodic phase as
well as charge Qa injected in anodic phase within
precisely defined stimulus as shown above, were
calculated and compared to each other. For this purpose,
an integral of the ic under a cathodic phase Qc as well as
the integral of the ia under an anodic phase of the
stimulus Qc, were calculated. By doing this, the
influence of the different stimuli on the offsets in the
measured CAPs that might be elicited due to an
imbalance between Qc as well as Qa, could be identified.
Since all the stimuli were current pulses expressed in
milliamperes, the corresponding charges Qc and Qa were
expressed in nAs.
However, to identify the differences in CAPs, elicited
by electrode-electrolyte interface established upon
intentionally exceeded limits for reversible charge
P. PE^LIN et al.: INFLUENCE OF TRANSIENT RESPONSE OF PLATINUM ELECTRODE ...
134 Materiali in tehnologije / Materials and technology 46 (2012) 2, 131–137
Figure 3: An isolated nerve segment (a), installed within the cuff (b)
and multi-stranded copper wires (c), mounted into the experimental
chamber
Slika 3: Izolirani segment `ivca (a), vstavljen v spiralno objemko (b),
in bakrene pletenice (c), zmontirani v poskusno celico
injection by different values of parameters and wave-
forms of selectively delivered stimulating pulses, an
integral of the CAP in cathodic phase as well as integral
of the CAP in anodic phase of stimuli were calculated
and compared to each other. Since all the CAPs were
voltage signals expressed in milivolts, the corresponding
integrals were expressed in nV s.
All the offline signal analyses were performed on a
Lenovo T420 portable computer using the Matlab
R2007a programming tool.
3 RESULTS
Figure 4 shows measured CAPs while the segment
was stimulated using quasitrapezoidal stimulus output
waveforms and parameters, namely ic, tc, texp, tau, ta and
ia, specifically preset in the abovementioned four tests.
As could be seen in recorded CAPs, the Figure 4 shown
waveforms and values of the CAP were slightly obscured
by stimulus artefacts and the transient response
characteristics of an electrode/neural tissue interface and
that of an inherent capacitance of the segment.
Table 1, however shows numerical values of calcu-
lated charge Qc injected in cathodic phase as well as
charge Qa injected in anodic phase of precisely defined
quasitrapezoidal stimuli selectively delivered to the
triplet5 in Tests1–4. Table 1 shows also values of
calculated integral of corresponding CAPs in cathodic
phase as well as in anodic phase of selectively delivered
stimulating pulses.
A Test 4 however, was considered only in the sense
of charge calculations while in a sense of integral
calculation it was not considered. Namely, in the Test 4,
corresponding CAP was measured using the couple of
electrodes localed at circumferentially opposite site
according to an triplet 5. Therefore, measured CAP
could not contain action potentials of nerve fibres
activated with an appointed triplet 5.
Table 1: Values and differences of cathodic Qc and anodic Qa charges
and values and differences of Integrals1 of the CAP manifested under
a cathodic phase and Integrals2 of the CAP manifested under an
anodic phase of the stimulating pulses in Tests1-4
Tabela 1: Vrednosti in razlike katodnih Qc in anodnih nabojev Qa ter
vrednosti in razlike Integrala 1 CAP-ov, zmontiranih pod katodno fazo
in Integrala 2 CAP-ov, izra`enih pod anodno fazo stimulusa pri testih
1–4
Variable Test 1 Test 2 Test 3 Test 4
Qc /(nA·s) 376.30 151.93 799.13 1107.82
Qa /(nA·s) 376.42 352.63 832.27 803.42
Q /(nA·s) 0.12 200.7 33.14 -304.4
Integral 1 /(nV·s) 250.57 113.01 665.22 300.38
Integral 2 /(nV·s) 59.85 61.50 76.15 44.95
Integral /(nV·s) 190.72 51.51 589.07 255.43
As result, Table 1 shows the difference Q of a
cathodic Qc and of an anodic charge Qa as well as the
difference of Integral1 of the CAP for cathodic phase
and Integral2 of the CAP for anodic phase of the
stimululating pulse for Tests1-4. Regardingly, in the
Test2, for instance, a charge Qc = 151.93 nAs was
injected in cathodic phase and a charge Qa = 352.63 nA s
was injected in anodic phase, yielding a positive
difference Qdiff = 200.7 nA s. This positive difference,
being relatively high, could expectedly elicit some
positive offset in the recorded CAP. In the Test4
however, a charge Qc = 1107.82 nA s was injected in
cathodic phase and a charge Qa = 803.42 nA s was
injected in anodic phase, yielding a negative difference
Qdiff = –304.4 nA s. This negative difference, being also
relatively high, could expectedly elicit some negative
offset in the recorded CAP. Fortunately, it could be seen
in Figure 4, that offset in all four measured CAPs, which
could arise as a consequence of predefined imbalance in
injected Qc and Qa, was not significant.
From the electrochemical point of view, it seems that
all the reactions that occured at the cathode (A in Figure
1) due to a charge Qc, injected via an ic in the cathodic
phase within a time tc, were reversed, in part or in full,
by the charge Qa, injected via the anodic phase ia within a
time ta. At each of the two triplet anodes in the same
longitudinal row of electrodes (A–A in Figure 1),
however, both the current and charge density would be
equal to one fourth of the current and charge density
occurring at the cathode. Namely, according to the model
not presented in the paper, both of the triplet anodes and
the two corresponding blocking electrodes (B–B in
Figure 1), were electrically connected. Therefore, the
electrochemical reactions that would occur at the single
anode could not be of the irreversible type.
However, this pattern would inevitable worsen in
case of stimulation in clinical practice where trains of
repetitive stimulating pulses are applied. In this case,
P. PE^LIN et al.: INFLUENCE OF TRANSIENT RESPONSE OF PLATINUM ELECTRODE ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 131–137 135
Figure 4: Specific waveform and values of CAPs measured in
aforementioned four tests: a) CAP measured in the Test 1; b) CAP
measured in the Test 2; c) CAP measured in the Test 3 and d) CAP
measured in the Test 4
Slika 4: Oblika in vrednosti CAP-ov, izmerjenih v zgoraj omenjenih
{tirih preizkusih: a) CAP, izmerjen v testu 1; b) CAP, izmerjen v testu
2; c) CAP, izmerjen v testu 3 in d) CAP, izmerjen v testu 4
excursions of potential of electrodes within stimulating
section due to predefined charge imbalance are additive
in each stimulating pulse and the resulting excursion and
consequently arised offset coud be significant.
4 DISCUSSION
An aim of the work was to contribute to the
development of models and multi-electrode cuffs to be
used for efficient and safe selective stimulation of
autonomous peripheral nerves and for selective recording
of CAPs at the same time.
The key developments in this technology were a cuff
that can expand and contract to provide a snug yet
non-compressing fit to the nerve, and a distributed
matrix of platinum stimulating electrodes which made
the performance of the cuff independent of it's
positioning around the nerve. This design has strong
potential for applications in neuro-prosthetic technology
in future9. Namely, it would be very desirable to control
different internal organs such as cardio-vascular system
in patients with heart failure or atrial fibrillation by only
one implanted system, e.g. on the lef cervical vagus
nerve.
However, it is unavoidable to understand the response
of peripheral nervous system elements to stresses that
may occur in the complex interactions that take place
between electrode and nerve secondary to VNS.
From CAP recording poinf of view, clinical use of
implanted electrodes is hampered by a lack of reliability
in chronic recordings, independent of the type of
electrodes used. Namely, persistent presence of the
electrode close to the neural tissue, causes a progressive
local neurodegenerative disease-like state surrounding
the electrode and is a potential cause for chronic
recording failure.
However, from stimulation point of view, nerve fibers
are located close to the stimulating electrode and also at
a certain distance from it, the electrode should be able to
inject enough charge to activate these fibers.25,27
However, for multielectrode stimulating systems,
containing miniature stimulating electrodes working at
relatively high charge densities, it is very important that
they are safe and electrochemically stable. Namely, to
avoid harm to the vagus nerve in clinical use of the cuff,
an inevitable requirement is the absence of irreversible
electrochemical reactions such as electrolysis of water,
evolution of chlorine gas or formation of metal oxides
that could cause severe tissue injury associated with high
charge density stimulation.16,17
Regarding both points of view, changes in the
complex impedance of stimulating and recording
electrodes in the cuff, chronically instaled onto a vagus
nerve, should be characterized in a series of animal
experiments.32
One weakness of a cuff manufacturing was a
technically demanding and a time consuming process.
Another weakness of a cuff was the use of a tin alloy at
the junctions between platinum electrodes and
multi-stranded wires. As this solution is not appropriate
for a clinical practice, a mechanical connection as a more
appropriate solution for further development of cuffs is
in preparation. Beside, a perfect electrical isolation of all
metals except electrode material is crucial for the life
time of the system, othervise the mentioned reactions
could occur.
Directions that our further work would be the
following:
• Further development of the cuff for given clinical
applications and accomplishment of electrochemical
measurements under realistic conditions using the
method of cyclic voltammetry, implementing the
Ag/AgCl reference electrode.
• Development of the strategies to enhance a capability
to obtainreliable long-term bipolar recordings of a
CAP from a particular couple of platinum electrodes
within the cuff.
It could be expected that mentioned directions, when
performed, will lead to an additional enhancement of the
cuff efficiency and to an overall more efficient and
effective implantable devices.
5 CONCLUSIONS
The most important findings of the present study are
the following:
• The strong component superimposed in the CAP was
an ensemble artefact which came exclusively from
the stimulating pulse via the transient response
characteristics of an electrode/neural tissue interface
and in part from an inherent capacitance of the
segment.
• One could speculate that stimulus artefact traveled
exclusively along the nerve surface via a capacitive
nature of an interface, while recording electrodes
measured an elicited voltage drop at the surface on
the nerve.
• Single stimulating pulses, having preset certain
degree of imbalance between a charge injected in
cathodic phase Qc and charge injected in anodic
phase Qa, elicited a slight change in a positive
waveform deflection of CAP manifested under a
cathodic phase as well as slight change in a negative
waveform deflection of a CAP manifested under an
anodic phase of the stimulating pulse.
• Measured CAPs are not greatly influenced by the
imbalance between a charge injected in cathodic and
anodic phase of quasitrapezoidal, asymmetric and
biphasic stimulating pulses.
• The reactions that occured at the cathode due to an
injected charge Qc were reversed, in part or in full, by
the charge Qa.
• The electrochemical reactions that occured at the
single anode could not be of irreversible type.
P. PE^LIN et al.: INFLUENCE OF TRANSIENT RESPONSE OF PLATINUM ELECTRODE ...
136 Materiali in tehnologije / Materials and technology 46 (2012) 2, 131–137
Acknowledgements
This study was supported in part by the Ministry of
Higher Education and Science, Republic of Slovenia,
Research Program P3-0171 and in part by the ITIS, d. o. o.,
Ljubljana, Centre for Implantable Technology and
Sensors.
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P. PE^LIN et al.: INFLUENCE OF TRANSIENT RESPONSE OF PLATINUM ELECTRODE ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 131–137 137

M. NASSER et al.: EFFECT OF THE ANTIMONY THIN-FILM DEPOSITION SEQUENCE ON COPPER-SILICON ...
EFFECT OF THE ANTIMONY THIN-FILM DEPOSITION
SEQUENCE ON COPPER-SILICON INTERDIFFUSION
VPLIV ZAPOREDJA NANOSA TANKIH PLASTI ANTIMONA NA
INTERDIFUZIJO BAKER-SILICIJ
Menni Nasser1, Boudissa Mokhtar1, Benkerri Mahfoud2, Reffas Mounir2,
Zekkar Fouzia1, Benazzouz Chaouki3
1ENMC Laboratory, F. Abbas University, Algeria
2LESIMS Laboratory, F. Abbas University, Algeria
3Nuclear research Center, 16000Algiers, Algeria
nacermenni@yahoo.fr
Prejem rokopisa – received: 2011-08-22; sprejem za objavo – accepted for publication: 2011-08-27
In this work we present a study of the effect of an antimony layer on the interdiffusion and formation of copper silicides while
inverting the sequence of Cu and Sb deposition on Si(111) substrates. Thermal evaporation was used to deposit Cu/Sb and
Sb/Cu bilayers on a Si(111) substrate heated at 100 °C, without breaking the vacuum. XRD and RBS analysis showed, for
samples heat treated at 200 °C and 400 °C, a segregation of the three elements (i.e., Cu, Sb and Si) to the surface and diffusion
in bulk ending in the formation of a layer made of a mixture containing the three elements at the samples’ surface. After 200 °C
annealing, in the Cu/Sb/Si system, we observed the formation of only Cu2Sb, and for the Sb/Cu/si system, there is the formation
of the Cu3Si and Cu2Sb phases; after 400 °C annealing, the Cu-Sb-Si mixture is formed by the cohabitation of the Cu3Si silicide
and the Cu2Sb intermetallic compound in the Cu/Sb/Si sample and only Cu2Sb in the Sb/Cu/Si sample. For the Cu/Sb/Si sample,
annealed at 400 °C, the SEM micrographs exhibit compound formation with crystallites that have a trapezoidal shape.
Keywords: thin film, PVD, diffusion, copper-antimony compound, copper silicide, Rutherford backscattering, scanning electron
microscopy, X-ray diffraction
V delu predstavljamo {tudijo vpliva plasti antimona na interdifuzijo in formiranje bakrovih silicidov pri inverziji zaporedja
nanosa Cu in Sb na Si(111)-podlago. Nanos plasti Cu/Sb in Sb/Cu je bil izvr{en s termi~nim izhlapevanjem na Si(111)-substrat
pri 100 °C brez prekinitve vakuuma. XRD- in RBS-analize vzorcev, `arjenih pri 200 °C in 400 °C, so pokazale segregacijo treh
elementov (Cu, Sb in Si) na povr{ini podlage in difuzijo masivnega materiala s tvorbo plasti zmesi treh elementov na povr{ini
podlage. Po `arjenju sistema Cu/Sb/Si pri 200 °C je bil opa`en nastanek Cu2Sb, v sistemu Sb/Cu/Si pa nastanek faz Cu3Si in
Cu2Sb. Po `arjenju pri 400 °C je bilo ugotovljeno sobivanje silicida Cu3Si in intermetalne spojine Cu2Sb pri vzorcu Cu/Sb/Si in
samo Cu2Sb pri vzorcu Sb/Cu/Si. Pri vzorcu Cu/Sb/Si, `arjenem pri 400 °C, SEM-posnetki ka`ejo nastanek spojin s kristali
trapezoidne oblike.
Klju~ne besede: tanke plasti, PVD, difuzija, spojina baker-antimon, bakrov silicid, Ruthefordovo povratno sipanje, vrsti~na
elektronska mikroskopija, difrakcija rentgenskih `arkov
1 INTRODUCTION
Antimony and copper are important elements for the
design of silicon-based electronics devices; indeed,
copper is predominant in interconnect metallization
material for deep sub-micrometer technology due to its
relatively low resistivity and electromigration resist-
ance1–3, and antimony is most frequently used as an
n-type doping element during silicon crystal growth3,4.
Unfortunately, the extensive diffusion of Cu in Si is
detrimental to the electrical performance of devices,
because Cu can form recombination-generation centers
in the active regions of Si. Additionally, it reacts with
silicon at very low temperatures, even below 150 °C,
forming Cu-Si silicides5–10 hence, diffusion barriers (e.g.,
an Sb layer) are necessary between the metallization
copper and Si substrate. To the best of our knowledge, no
study on the Cu-Sb-Si thin film system has been
published in the literature so far, though a study of the
Cu-Sb thin-film system deposited on glass substrate,
annealed up to 600 °C, has shown that during annealing
at up to 300°C, only the Cu2Sb compound is formed.
During subsequent heat treatments from 350 °C to 600
°C, the Cu9Sb2 phase nucleates and grows at the expense
of the Cu2Sb phase11. In a previous work12,13, the effect of
temperature and doping-element redistribution on atomic
interdiffusion between a thin copper layer and
monocrystalline silicon, implanted and Sb+ unimplanted
with different doses, was investigated. The results show
the growth and formation of Cu3Si and Cu4Si silicides
under crystallites shape dispatched on the sample
surface, independently of the implantation dose. On the
other hand, it was established that the copper layer is less
and less consumed as the antimony dose increases,
resulting in an accumulation of Sb ions at the silicide/Si
interface and in the silicide layer close to the surface.
However, the low antimony quantity in the presence has
not led to an understanding of the reactions which could
take place between the three elements and the inter
diffusion mechanism involving the Sb and Cu layers and
the Si substrate.
In the present study, we focus on copper, antimony
and silicon atoms’ interdiffusion in the Cu-Sb-Si thin-
layers system, while inverting the sequence of evapo-
Materiali in tehnologije / Materials and technology 46 (2012) 2, 139–144 139
UDK 681.7.026.6:539.23:669.75 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)139(2012)
ration of the Cu and Sb thin films on monocrystalline
(111) Si. Indeed, from a metallurgical point of view, it is
very interesting to study the Sb/Cu/Si and Cu/Sb/Si
sequences in order to obtain some information about the
influence of Sb on the diffusion process and on the
formation of the different phases.
2 EXPERIMENTAL PROCEDURES
N-type monocrystalline(111)-oriented silicon wafers
with a resistivity of 1–3  cm were used as substrates. In
order to eliminate as far as possible the residual layer of
silicon oxide, the wafers were etched in 10 %
hydrofluoric (HF) acid for 20 s and rinsed in de-ionized
water prior to loading into the vacuum system. Copper
and antimony thin films were thermally evaporated
alternatively, Cu/Sb as well as Sb/Cu, onto the Si(111)
substrate heated at 100 °C for a better adhesion of the
antimony layer. These evaporations were performed
without breaking the vacuum of 5.23·10–7 mbar obtained
with a turbomolecular pump. An in-situ quartz crystal
oscillator allows measurement of the thickness of the
copper and antimony layers, which were 70 nm and 30
nm, respectively, in the Cu/Sb systems, and 100 nm and
30 nm, respectively, in Sb/Cu layers. In order to promote
diffusion, conventional annealing treatments were
performed at temperatures of 200 °C and 400 °C, for 45
min, in a quartz tube, under a vacuum of about 2.66·10–7
mbar. It is important to note that the first steps range of
the reaction at the different interfaces happen at these
temperatures of annealing.
The surface morphology characterization of the
samples, before and after the heat treatments, was carried
out using a JEOL JSM-6460 scanning electron micro-
scope with an accelerating voltage of 20 kV and an EDX
analyzer. The primary energy of the electron beam is
chosen to be equal to 10 keV in order to limit the
analyzed depth to about 300 nm into the surface. A
Siemens D5000 diffractometer, in –2 mode, was used
to identify the formed compounds. The concentration
profiles of the Cu, Sb and Si were determined with the
help of spectra simulated with the RUMP program by
fitting the experimental RBS spectra recorded at (2 MeV,
4He+) with the detector positioned at 160° relative to the
beam.
3 RESULTS AND DISCUSSION
Typical X-ray diffraction patterns corresponding to
the Sb/Cu/Si(111) and Cu/Sb/Si(111) systems, as
deposited, are shown in Figures 1a and 2a, respectively.
The nominal preheating of the substrate at 100 °C was
used during evaporation in order to enhance the
in-surface adhesion. On the XRD diagram corresponding
to th Sb/Cu bilayers deposited on the Si(111), there is
M. NASSER et al.: EFFECT OF THE ANTIMONY THIN-FILM DEPOSITION SEQUENCE ON COPPER-SILICON ...
140 Materiali in tehnologije / Materials and technology 46 (2012) 2, 139–144
Figure 3: Rutherford backscattering spectra of Cu/Sb bilayers on
silicon substrate heated at 100 °C: as-deposited and annealed at 200
°C and 400°C
Slika 3: Spektri Ruthefordovega povratnega sipanja za Cu/Sb-plasti
na Si(111)-podlagi pri 100 °C: naneseni in `arjeni pri 200 °C in 400
°C
Figure 1: X-ray diffraction patterns of Sb/Cu bilayers on Si(111)
substrate heated at 100 °C: a) as-deposited and annealed at b) 200 °C
and c) 400 °C
Slika 1: Rentgenski uklonski spektri Sb/Cu-plasti na Si(111)-podlago
pri 100 °C: a) naneseni in `arjeni pri b) 200 °C in c) 400 °C
Figure 2: X-ray diffraction patterns of Cu/Sb bilayers on Si(111)
substrate heated at 100 °C: a) as-deposited and annealed at b) 200 °C
and c) 400 °C
Slika 2: Rentgenski uklonski spektri Cu/Sb-plasti na Si(111)-podlago
pri 100 °C: a) naneseni in `arjeni pri b) 200 °C in c) 400 °C
clear evidence of the main (111)Cu and (200)Cu
reflection lines, and the absence of that of the evaporated
antimony layer, which could be amorphous or formed
from nanometric grains. In the Cu/Sb/Si sample, in
addition to the high texturization of the antimony layer,
we can see, before any annealing, the apparition of the
Cu2Sb intermetallic compound. RBS spectra, for both
as-deposited samples, shown in Figure 3 and Figure 4,
reveal that the different interfaces are not abrupt. This
means that the preheating of the substrate during
evaporation has already led to atomic interdiffusion at
the different interfaces. Particularly in the Cu/Sb/Si
sequence case, where the RBS spectrum shows a slight
shift of the silicon signal towards the high-energies side,
with a 30 %. Si concentration in the outer layer. This is
synonymous with the silicon atoms segregation to the
sample surface, which has as a consequence the
beginning of the asperities formation shown on the
corresponding micrograph, (Figure 5a). On the other
hand, in the Sb/Cu/Si system, we can see the displace-
M. NASSER et al.: EFFECT OF THE ANTIMONY THIN-FILM DEPOSITION SEQUENCE ON COPPER-SILICON ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 139–144 141
Figure 6: SEM micrographs from the surfaces of Sb/Cubilayers on
silicon substrate heated at 100 °C: a) as-deposited and heat treated at
b) 200 °C and c) 400 °C
Slika 6: SEM-posnetki povr{ine plasti Sb/Cu na Si(111)-podlagi pri
100 °C: a) naneseni in `arjeni pri b) 200 °C in c) 400 °C
Figure 5: SEM micrographs from the surfaces of Cu/Sb bilayers on
silicon substrate heated at 100 °C: a) as-deposited and heat treated at
b) 200 °C and c) 400 °C
Slika 5: SEM-posnetki povr{ine plasti Cu/Sb na Si(111)-podlagi pri
100 °C: a) naneseni in `arjeni pri b) 200 °C in c) 400 °C
Figure 4: Rutherford backscattering spectra of Sb/Cu bilayers on
silicon substrate heated at 100 °C: as-deposited and annealed at 200
°C and 400 °C
Slika 4: Spektri Ruthefordovega povratnega sipanja za Sb/Cu-plasti
na Si(111)-podlagi pri 100 °C: naneseni in `arjeni pri 200 °C in 400
°C
ment of the Cu RBS spectrum to lower energy, due to its
diffusion into the silicon substrate. Again, a shoulder
appears on the right-hand side of the Cu spectrum, which
means that an appreciable amount of the mole fraction of
Cu (x = 68 %) has already segregated to the surface
through the antimony top layer. In this latter case, the
diffusion of copper atoms is favored. In this pre-reac-
tional diffusion, the displacement of Sb atoms appears
timorous because of the low diffusion coefficient of Sb
into Si. Indeed, according to Fahey et al.14, the Sb
diffusion coefficient in Si, under equilibrium conditions,
is 1.2·10–19 cm2/s. However, in the solid state, antimony
is partially soluble in copper, with the solubility
decreasing with temperature, and, it is more plausible
that Cu atoms dissolve in the antimony layer rather than
the opposite, because the atomic radius of Cu (0.128 nm)
is smaller than that of Sb (0.145 nm).
After 200 °C annealing (Figure 1b), X-ray diffrac-
tion patterns for the Sb/Cu/si system show that in the
inner copper layer the copper reacted with Si and Sb to
give reflection lines corresponding to the Cu3Si and
Cu2Sb phases. Whereas in the outer copper layer system
(Cu/Sb/Si) we see the formation of only Cu2Sb, which
leads to a decrease of the (111) and (100)Cu reflection
lines’ intensities, as indicated in Figure 2b). For this
latter case, the surface morphology shows that the
asperities of the as-deposited samples have served as
nucleation sites for the first diffusion steps on the sample
surface (Figure 5b). According to the RBS signals of the
silicon, for both samples at 200 °C annealing, the
evolution of an enhanced silicon concentration in the
surface is very easily seen. When copper is the first
deposited layer onto the silicon, an important quantity of
copper (x(Cu) = 63 %) has diffused in the bulk as a
consequence of its exceptionally fast diffusivity in the Si
substrate. Copper is known as the fastest diffusing
element in silicon among all the transition metals, for
depths of a few micrometers. For instance, the Cu
diffusion coefficient in Si is 1.4·10–6 cm2/s at 400 °C and
copper atoms move interstitially in the lattice to form an
interstitial solid solution15,16. For the Cu/Sb/Si sample
heat treated at 200 °C, (Figure 3), the RBS signal of
antimony seems to decrease in terms of yield intensity
versus the diffusion of silicon and the presence of copper
atoms at the surface, to form the mole fractions 39 % Cu,
15 % Sb and 45 % Si mixture layer that is 140 nm thick.
This phenomenon is more pronounced during 400 °C
annealing, for both samples, where the heights of the
copper and antimony signals have drastically decreased
due to the transport of the three elements, (Figures 3 and
4). Besides, the rear edges for both metallic peaks have
extended to lower channel numbers, while the front edge
of the silicon signal has extended to the higher channel
numbers. This clearly indicates that intermixing has
occurred across the different interfaces and in the
surface. In both cases, the formed alloy is evaluated from
the slopes of the rear edge of the metal layers from the
metallic signals, and it is evident from the RBS spectra
of the Sb/Cu/Si sample, that the slopes of the rear edges
of the Cu and Sb are more important than in the
Cu/Sb/Si sample. These rear edges of the Cu and Sb
RBS signals were simulated while taking the 12 % Cu, 5
% Sb, 83 % Si and 26 % Cu, 6 % Sb, 68 % Si for Cu/Sb
and Sb/Cu bilayers, respectively. For both systems, the
diffusion of all the elements is clearly visible and is also
dramatic on the signals’ shape with a perturbation of the
silicon substrate over about 1300 nm in depth. This
Cu-Sb-Si mixture is really formed by the cohabitation of
the Cu3Si silicide and the Cu2Sb intermetallic compound
in the Cu/Sb/Si and only the Cu2Sb in the Sb/Cu/Si
sample, after 400 °C annealing Figures 1c and 2c. In this
latter case, the disappearance of Cu3Si silicide to the
benefit of the Cu2Sb phase is foreseeable owing to the
fact that the growth and formation of the second
compound is energetically more favorable, as will be
argued later in this paper. The persistence of the two
compounds in the Cu/Sb/Si sample means that the
reaction is less pronounced than in the Sb/Cu/Si sample.
Equilibrium phase diagrams of the Sb-Si2, Cu-Sb4
and Cu-Si17 binary alloys, which bind the ternary
Cu/Sb/Si system, have been well known for a long time.
Among the binary systems with an important segre-
gation, Cu–Sb and Cu-Si are particularly interesting for
their strong tendency to form ordered compounds And it
is known that Sb/Si is a simple eutectic system, which
induces a low limited mutual solubility of Sb and Si and,
with no intermediate phase formation because Sb has a
high mass and low diffusivity in Si. Indeed, the solid
solubility of Sb in Si is very low: the maximum equili-
brium solubility of Sb in Si is 0.l % during an
equilibrium processes, whereas that of silicon in Sb is
negligible. In the solid state, antimony is partially
soluble in copper, with the proportion decreasing with
temperature. The solid solubilities of 1.5 % and 3 % Sb
in Cu are reached at temperatures of 250 °C and 300 °C,
respectively. A maximum solubility of 5.8 % antimony
in a copper matrix is reached at a temperature of 645
°C8. On the other hand, the copper atoms’ diffusion in Si
is essentially interstitial with an activation energy of 0.43
eV and a limited solubility of about 1·1015 cm–3 at 600
°C18,19. In addition, Cu-Sb and Cu-Si couples are com-
pletely miscible with the formation of ordered
compounds, such as Cu2Sb, Cu3Sb, 	Cu11Sb2, Cu9Sb2
and Cu3Si, Cu4Si, Cu0.83Si0.17, Cu5Si, respectively,
depending on the composition in the solid solution.
However, it is reported as the growth and formation, as
the first phases of Cu2Sb and Cu3Si in their corres-
ponding binary systems, respectively. This formation of
Cu3Si and Cu2Sb phases in our case is in conformity with
results reported in the literature11,20 and is in agreement
with both the prediction21 and the thermodynamic
minimization of the free enthalpy H. From the Cu-Sb
system, the reported free energies of formation of the
Cu2Sb, Cu9Sb2, 	Cu11Sb2 intermetallic compounds are
M. NASSER et al.: EFFECT OF THE ANTIMONY THIN-FILM DEPOSITION SEQUENCE ON COPPER-SILICON ...
142 Materiali in tehnologije / Materials and technology 46 (2012) 2, 139–144
(–4.23, –0.54 and –0.29) kJ/mol, respectively22, whereas
from the Cu-Si system, those of the Cu3Si, Cu4Si, and
Cu5Si silicides are (–4.1, –3.4, and –2.9) kJ/mol,
respectively23. Thermodynamically, the formation of the
first stable compound at the interface requires the lowest
free enthalpy and is consequently favored. In other
words, the Cu2Sb and Cu3Si compounds have the lowest
heat of formation of their mother compounds and are
energetically more favorable. While adopting the same
argumentation, if these two compounds are in competi-
tion, it is clear that the formation of Cu2Sb is more
favorable than that of Cu3Si. Indeed, these results
confirm that copper atoms are the dominant diffusion
species in both compounds, which is in agreement with
the rule that stipulates that the majority of atoms in the
formed phase should be more mobile than the minority
atoms24. The observed compounds’ compositions (XRD
spectra) in the equilibrated mixture (RBS spectra) show
that an equilibrium exists between the Cu3Si and Cu2Sb
phases, starting from 200 °C. This temperature of
formation, surprisingly low compared to those leading to
the formation of silicides and intermetallic compounds
of refractory metals, can be attributed to the high
diffusivity of the copper in both the antimony and silicon
matrices25.
Figure 5 illustrates the evolution of the surface
morphology with temperature, examined by scanning
electron microscope, for the Cu/Sb/Si(111) sample. The
first micrograph (magnif. 2000-times) of the as-depo-
sited sample shows some asperities, probably due to the
preheating of the silicon substrate; this contrasts with the
presupposed uniform aspect that the evaporated metallic
layer on the cold substrate should have. Starting from
200 °C, as seen in Figure 5b, crystallites sprinkled on
the samples’ surface, begin to take shape according to the
same oblique lines. After the 400 °C annealing (Figure
5c) we can see large white crystallites with a micro-
metric dimension of the trapezoidal shape, grown at
surface as consequence of the great interdiffusion
between the three elements. This is due to the segre-
gation and coalescence of the elements above the sample
surface, leading to the growth and formation of Cu2Sb
and Cu3Si crystallites with heights of over one micro-
meter, which is equivalent to the simulated thickness of
the coalesced metal layers. Particularly for the Cu3Si
silicide, a similar epitaxial growth has already been
reported in n octahedral shape on Si(100)26 and different
shapes with and without an intermediate metallic layer9.
In this study, such oriented crystallites’ formation was
not observed for the Sb/Cu/Si(111) sequence, in spite of
enlargement of 4000-times, (Figure 6). Indeed, in this
case, according to X-ray diffraction diagram, there is no
formation of the Cu3Si phase at 400 °C.
4 CONCLUSION
The XRD, RBS and SEM techniques have been used
to investigate atomic interdiffusion in Cu/Sb and Sb/Cu
bilayers deposited on monocrystalline silicon with a
(111) orientation. In the Cu/Sb/Si sample, in addition to
the high texturization of the antimony layer, we already
se, before any annealing, the appearance of Cu2Sb inter-
metallic compound due to substrate heating at 100 °C.
Heat treatments of 200 °C and then 400 °C, reveal a
strong Cu-Sb-Si intermixing between the three elements
for both systems, resulting in compounds formation.
After the 200 °C annealing, in the Cu/Sb/Si system, we
see the formation of only Cu2Sb, and for the Sb/Cu/si
system, there is the formation of the Cu3Si and Cu2Sb
phases. After 400 °C annealing, the Cu-Sb-Si mixture is
formed by the cohabitation of the Cu3Si silicide and the
Cu2Sb intermetallic compound in Cu/Sb/Si and only
Cu2Sb in the Sb/Cu/Si sample. For both systems, the
diffusion of all elements is clearly visible and is also
dramatic on the RBS signals’ shape with a perturbation
of silicon substrate of about 1300 nm in depth. Indeed, at
400 °C, SEM micrographs for Cu/Sb/Si sample show the
formation of Cu2Sb and Cu3Si crystallites with heights of
over one micrometer, which is comparatively equivalent
to the simulated thickness of the coalesced metal layers.
In this study, such oriented crystallites formation were
not observed for the Sb/Cu/Si(111) sequence in spite of
the SEM micrograph enlargement of 4000-times.
5 REFERENCES
1 S. M. Murarka, I.V. Verner, R. J. Gitmann, Copper – Fundamental
Mechanisms for Microelectronic Applications, Ed., John Wiley and
Sons, New York, 2000, chapter I
2 J. F. Gui Llaumond, Étude de la résistivité et de l’é lectromigration
dans les interconne xions destiné es aux technologies des nœuds 90
nm – 32 nm Ph.D. Thesis, Universite Joseph Fourier – Grenoble 1,
2005
3 A. R. Powell, R. A. A. Kubiak, S. M. Newstead, C. Parry, N. L.
Mattey, D. W. Smith et al., Elemental boron and antimony doping of
MBE Si and SiGe structures grown at temperatures below 600 °C,
Journal of Crystal Growth, 111 (1991) 1–4, 907–911
4 A. Csik, G. A. Langer, G. Erdélyi, D. L. Beke, Z. Erdelyi, K. Vad,
Investigation of Sb diffusion in amorphous silicon, Vacuum Volume
82, Issue 2, 29 October 2007, Pages 257–260 Proceedings of the 11th
Joint Vacuum Conference (JVC-11)
5 M. Uekubo, T. Oku, K. Nii, M. Murakami, K. Takahiro, S. Yama-
guchi, T. Nakano, T. Ohta, WNx diffusion barriers between Si and
Cu, Thin Solid Films, 286 (1996) 1–2, 170–175
6 V. V. Jain, Microstructure and properties of copper thin films on
silicon substrates. M. S. dissertation, Texas A&M University, 2007
7 N. Benouattas, A. Mosser, A. Bouabellou, Surface morphology and
reaction at Cu/Si interface – Effect of native silicon suboxide,
Applied Surface Science, 252 (2006), 7572
8 N. Benouattas, L. Osmani, L. Salik, C. Benazzouz, M. Benkerri, A.
Bouabellou, R. Halimi, Epitaxial growth of copper silicides by
bilayer technique on monocrystalline silicon with native SiOx,
Materials Science and Engineering B, 132 (2006), 283
9 C. Benazzouz, N. Benouattas, A. Bouabellou, Competitive diffusion
of gold and copper atoms in Cu/Au/Si and Au/Cu/Si annealed
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230 (2005), 571
10 C. Benazzouz, N. Benouattas, S. Iaiche, A. Bouabellou, Study of
diffusion at surface of multilayered Cu/Au films on monocrystalline
silicon, Nuclear Instruments and Methods in Physics Research B,
213 (2004), 519
11 R. Halimi, A. Merabet, Kinetics of compound formation in Cu-Sb
thin films, Surface Science, North-Holland, Amsterdam 223 (1989),
599–606
12 M. Benkerri, R. Halimi, A. Bouabellou, N. Benouattas, Effect of Sb+
implantation on copper silicides formation and morphology after
annealing of Cu/Si structures, Materials Science in Semiconductor
Processing, 7 (2004), 319–324
13 M. Benkerri, R. Halimi, A. Bouabellou, Antimony dopant redistri-
bution during copper silicide formation, International Journal of
Inorganic Materials, V3(8) (2001), 1299
14 P. M. Fahey, P. B. Griffin, J. D. Plummer, Point Defects and Dopant
Diffusion in Silicon, Review of Modern Physics, 61 (1989) 2,
289–384
15 Y. G. Celebi, Copper Diffusion in silicon, Ph.D dissertation Texas
Tech University USA 1998
16 T. Heiser, A. Mesli, Determination of the copper diffusion coefficient
in silicon from transient ion-drift, Applied Physics A: Materials
Science & Processing, 57 (2004) 4, 325–328
17 M. Hansen, Constitution of Binary Alloys, Ed. McGraw-Hill, New
York, 1958
18 H. Bracht, Copper related diffusion phenomena in germanium and
silicon, Materials Science in Semiconductor Processing, 7 (2004) 3,
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Cu2Sb and Cu9Sb2 as anode materials in Li-ion batteries, Electro-
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Nuclear Instruments and Methods in Physics Research, 209–210
(1983) 1, 97–105
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don, 1976
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Vol. 6, ed. by NG Einspruch, Materials and Process Characterization,
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1983, Chap 7
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and related structures, J. Appl. Phys., 66 (1989) 3, 1163
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deposition on Si(100) oriented wafers, Journal of Crystal Growth,
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144 Materiali in tehnologije / Materials and technology 46 (2012) 2, 139–144
A. L. GAMEIRO et al.: LIME-METAKAOLIN HYDRATION PRODUCTS: A MICROSCOPY ANALYSIS
LIME-METAKAOLIN HYDRATION PRODUCTS: A
MICROSCOPY ANALYSIS
PRODUKTI HIDRACIJE APNO-METAKAOLIN: MIKROSKOPSKA
ANALIZA
André Leal Gameiro1, Antonio Santos Silva1, Maria do Rosário Veiga1, Ana Luísa Velosa2
1National Laboratory of Civil Engineering, Av. do Brasil, 101, Lisbon, Portugal
2Department of Civil Engineering, Geobiotec, University of Aveiro, Aveiro
agameiro@lnec.pt
Prejem rokopisa – received: 2011-10-01; sprejem za objavo – accepted for publication: 2011-12-02
Metakaolin (MK) is nowadays a well-known pozzolanic material being used in cement-based materials, like mortars and
concretes. The reaction of MK with calcium hydroxide yields cementitious products, being calcium silicate hydrate (CSH),
stratlingite (C2ASH8) and tetra calcium aluminium hydrate (C4AH13) the main phases formed at ambient temperature. The
transformation of stratlingite and C4AH13 into hydrogarnet at long term is an important issue that may result in an increase in the
porosity and a loss of compressive strength that can induce a complete material degradation.
With the objective of studying the compounds formed in lime/MK pastes and their stability during time, blended pastes were
prepared with several substitution rates (in weight) of lime by MK, and maintained at RH > 95 % and 23 ± 2 °C. XRD and
TGA-DTA were used to follow the kinetics of the lime/MK hydration as well the reaction products. Microscopic tests
(SEM-EDS) results are performed and compared with the thermal and mineralogical data. The results obtained show that the
quantity of the hydration products formed changes with the lime replacement, being the aluminum and calcium silicates more
abundant in the higher MK content pastes, and C4AH13, C4ACH11 and C2ASH8 the major phases formed up to 90 days of curing.
Keywords: microscopy, XRD, TGA-DTA, lime, metakaolin
Metakaolin (MK) je danes dobro poznan pocolanski material in se uporablja v gradivih na podlagi cementa, npr. malta in beton.
Reakcija MK s kalcijevim hidroksidom ustvari cementne proizvode: kalcijev silikatni hidrat (CSH), stratlingit (C2ASH8) in tetra
kalcij aluminijev hidrat(C4AH13) kot glavne faze, ki nastanejo pri temperaturi okolice. Transformacija stratlingita in C4AH13 v
hidrogarnet po dolgem ~asu je proces, ki lahko pove~a poroznost in zmanj{a tla~no trdnost ter lahko povzro~i popolno
degradacijo materiala.
S ciljem raziskave spojin, ki so nastale v zmeseh apno-MK, in njihove ~asovne stabilnosti smo pripravili me{ane mase z ve~
dele`i zamenjave (v masi) apna z VK in zorili pri RH > 95 % in 23 ± 2 °C. XRD in TGA-DTA so bile uporabljene za
spremljanje kinetike hidracije apna-MK in reakcijskih produktov. Rezultate mikroskopskih opazovanj (SEM-EDS) smo
primerjali s termalnimi in mineralo{kimi podatki. Dobljeni rezultati ka`ejo, da koli~ina hidracijskih produktov z nadomestitvijo
koli~ine apna koli~ina aluminijevih in kalcijeviih silikatov raste z vsebnostjo MK v masi in so C4AH13, C4ACH11, C2ASH8
glavne faze, nastale po 90 dneh zorenja.
Klju~ne besede: mikroskopija, XRD, TGA-DTA, apno, metakaolin
1 INTRODUCTION
During the past decade, metakaolin (MK), a ther-
mally activated aluminosilicate material (Al2O3·2SiO2)
obtained by calcination of clay or soil rich in kaolinite
(Al2(OH)4Si2O5), within the temperature range of
700–850 °C 1, has been the objective of several research
studies, mainly due to its capacity to react vividly with
calcium hydroxide (pozzolanicity). Recent studies2
showed that MK is a very effective pozzolan, altering the
pore structure of the lime and cement pastes and greatly
improving its resistance to the transport of water and
diffusion of harmful ions through the matrix, supporting
the idea of its beneficial addition in blended mortars,
cement pastes and concrete.
The reaction between MK, calcium hydroxide and
water results in the formation of hydraulic products. At
ambient temperature the main phases/products formed
are calcium silicate hydrate gel (CSH), stratlingite
(C2ASH8) and tetra calcium aluminium hydrate
(C4AH13). According to literature3, it is possible to
assume that these hydration phases are linked to
lime/MK ratio, temperature, and also to the presence of
various activators. However, reference to variations
regarding these phase’s stability has been shown in
literature4. As stated by P. S. Silva and Glasser, transfor-
mation of stratlingite and C4AH13 into hydrogarnet
(C3AH6) at long term may lead to a volume reduction,
producing an increase in porosity and a loss of micro-
structural compactness, i.e., less mechanical strength.
Having the aim of studying the compounds obtained
in lime/MK pastes and their stability overtime, it is
important to discuss a possible conversion of metastable
hexagonal hydrates (C2ASH8 and C4AH13) to stable cubic
phase (siliceous hydrogarnet with variable composition,
C3ASxH6–2x), when samples are submitted to ambient
curing temperatures at long term ages, negatively
influencing the performance of MK blended matrixes,
especially on durability5.
Most of the works published up to now on lime/MK
systems are focused on reaction kinetics and on its
interaction with Portland cement3,5,6. However, there is
Materiali in tehnologije / Materials and technology 46 (2012) 2, 145–148 145
UDK 691.5:691.26:620.187 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)145(2012)
little information on microstructure and mechanical
characteristics of the lime/MK pastes6.
In order to study the development of hydration
phases of the lime/MK systems, pastes with different
mixing mass ratios of lime/MK were prepared and
afterwards cured at 23 °C and RH > 95 %.
At certain curing ages, XRD and TGA-DTA were
used to follow the kinetics of the blended MK pastes as
well the reaction products formed, while microscopic
tests (SEM-EDS) results are performed and compared
with the thermal and mineralogical data.
2 EXPERIMENTAL
The mix procedure consisted in mixing the amount of
lime with the total amount of water, which was stirred
for about 3 min using an external mixer, after which MK
was added slowly, maintaining the mixing for further 20
min.7 The pastes were then stored in open plastic
containers and introduced in a sealed chamber at RH >
95 % and 23 ± 2 °C, to maintain on-going hydration
reactions. The hydration was stopped after each
predetermined curing time, subjecting the samples to
acetone for complete removal of the free water, after
which they were then dried at 40 °C, in order to be tested
by XRD, TGA-DTA and SEM-EDS. The metakaolin
used was ARGICAL M1200S from IMERYS with the
mass fractions SiO2  55 % and Al2O3  39 % regard-
ing chemical composition, while the lime used was a
commercial Portuguese (Lusical H100) hydrated lime
with a chemical composition of Ca(OH)2  93 and MgO
 3, with classification CL90, according to the NP EN
459-1(2002) standard.
The evolution of the kinetic reactions as well as the
phases formation was evaluated by X-ray diffractometry
(XRD), thermogravimetric and differential thermal
analysis (TGA-DTA) and scanning electron microscopy
with X-ray microanalysis (SEM-EDS). The experimental
conditions used are previously published8.
3 RESULTS AND DISCUSSION
In order to illustrate the main results obtained two
pastes were selected with lime/MK mass ratios of 1/1
and 1/0.2.
3.1 X-ray diffraction analysis (XRD)
XRD patterns are illustrated in Figures 1a and b. A
peak attributed to stratlingite (C2ASH8) is noted for paste
MK1. The stratlingite tends to increase with the curing
time, becoming the dominant phase, whereas significant
amounts of monocarboaluminate (C4ACH11), quartz,
calcite, calcium silicate hydrates (CSH) and calcium
aluminate hydrates (C4AH13) are observed. Although in
paste MK02 stratlingite is not detected by XRD, it is
possible that it may be present in very low, undetected
quantity. An interesting result was also observed in paste
MK1 up to 28 d of curing regarding the presence of
calcium aluminates hydrates (C4ACH11 and C4AH13),
whereas at further ages (56 d and 90 d) only traces of
these compounds are identified, presumably signifying
that a decomposition of these phases may have occurred.
Traces of crystallized CSH were detected in higher MK
mixes.
In MK02 paste high amounts of portlandite (CH) are
observed, however a maximum peak is noticed for 56 d
and 90 d of reaction, possibly due to the decomposition
of C4AH13 and monocarboaluminate, liberating more
portlandite to the system, also observed for paste MK1.
The XRD confirms also up to 90 d the non-formation of
hydrogarnet phase for both pastes, which can be
explained due to the fact that this compound is
associated with higher curing temperatures. Reports of
the absence of hydrogarnet until 270 d of curing time are
shown by other authors6.
An important reference must be attributed to the fact
that for both pastes the calcite (CaCO3) content tends to
increase up to 90 d of curing.
Several studies have reported the appearance C4AH13
or C4ACH11 and C2ASH8 and CSH at 20 °C5 as the reac-
tion products of the lime/MK hydration reaction. In other
studies4, C2ASH8 and CSH are considered the main
phases formed, however in this research, the formation
A. L. GAMEIRO et al.: LIME-METAKAOLIN HYDRATION PRODUCTS: A MICROSCOPY ANALYSIS
146 Materiali in tehnologije / Materials and technology 46 (2012) 2, 145–148
Figure 1: XRD patterns for a) MK1 and b) MK02; Cc – calcite; P –
portlandite; MC – monocarboaluminate; St – stratlingite; CSH –
calcium silicate hydrate; C4AH13 – tetracalcium aluminium hydrate
Slika 1: XRD-spektri za a) MK1 in b) MK02; Cc – kalcit, P –
portlandit, MC – monokarboaluminat, St – stratlingit, CSH – kalcijev
silikat hidrat, C4AH13 – tetrakalcij aluminijev hidrat
of CSH, C4AH13, C4ACH11 and C2ASH8 have been
identified.
3.2 TGA-DTA analysis
Figures 2a and b present the DTA curves of MK1
and MK02. The sharp peak observed at about 110 °C can
be attributed to the presence of CSH and C4AH13, being
sharper for 28 d, disappearing after 90 d of reaction. The
characteristic peak of stratlingite, which appears as a
sharp peak at about 190 °C, can be observed for MK1,
however, a broader peak after 90 d of curing. In the same
range of temperature (160–220 °C) the dehydration of
C4ACH11 (monocarboaluminate) occurs.
For paste MK1 a small broad peak at about 480 °C is
present at 1 d reaction, indicative of the existence of free
portlandite, being almost consumed after 1 d of curing
(Figure 2a). Instead, for paste MK02, a sharper band can
be seen at approximately 500 °C (Figure 2b), due to free
portlandite in the reaction.
As referred above at about 950 °C, a sharp exother-
mic peak was only observed for paste MK1, attributed to
the formation of high-temperature phases such as mullite
and cristobalite, as described by Bakolas, appearing due
to the existence of "free" MK. These peaks do not appear
in paste MK02 due to the total amount of MK being
rapidly consumed.
3.3 SEM results
Figures 3a, b, c and d show the main results of the
SEM-EDS. Paste MK1 at 28 d of reaction presents
(Figure 3a) a well-crystallized matrix, with large
amounts of C4AH13; instead at 90 d of curing (Figure
3b) the paste matrix is more "densified" and less
crystalline with a large presence of stratlingite. In paste
MK02 (Figures 3c and d), the microstructure "evolu-
tion" does not vary like in MK1 from 28 d to 90 d,
remaining as a crystalline microstructure, revealing high
amounts of calcite.
A. L. GAMEIRO et al.: LIME-METAKAOLIN HYDRATION PRODUCTS: A MICROSCOPY ANALYSIS
Materiali in tehnologije / Materials and technology 46 (2012) 2, 145–148 147
Figure 3: SEM micrographs of MK1 and MK02 at ages 28 d and 90 d:
a) SEM image of MK1 paste at 28 d where the presence of hexagonal
C4AH13 is visible; b) SEM image of MK1 at 90 d where is visible the
predominance of C2ASH8; c) SEM image of MK02 at 28 d of curing;
revealing the presence of CaCO3; d) SEM image of MK02 paste at 90
d of curing where is visible an increase in the paste carbonation rate
Slika 3: SEM-posnetki MK1 in MK02 po zorenju 28 d in 90 d: a)
SEM-posnetek MK1-mase po 28 d, kjer je viden heksagonalni
C4AH13, b) SEM-posnetek MK1 po 90 d, kjer je vidna prevlada
C2ASH8, c) SEM-posnetek MK02 po 28 d, ki dokazuje prisotnost
CaCO3, d) SEM-posnetek MK02, kjer se v masi vidi rast hitrosti
karbonacije
Figure 2: DTA curves for a) MK1 and b) MK02; P – portlandite; CM
– cristoballite and mullite; St – stratlingite; CSH – calcium silicate
hydrate; Cc – calcite; MC – monocarboaluminate; C4AH13 – tetra-
calcium aluminum hydrate
Slika 2: DTA-spektri za a) MK1 in b) MK02; P – portlandit, CM –
kristobalit in mulit, St – stratlingit, CSH – kalcijev silikat hidrat, Cc –
kalcit, MC – monokarboaluminat, C4AH13 – tetrakalcijev aluminijev
hidrat
4 CONCLUSIONS
The influence of the lime/MK ratio was studied at
ambient temperature and RH > 95 %. According to the
results obtained the products formed are the same for all
blended mixes, while the amount formed changes with
the MK content, being the aluminum (C4AH13, C4ACH11)
and calcium silicates (C2ASH8) more abundant in the
higher MK content pastes. SEM results show that for
MK1 at 28 d the microstructure consists essentially of
C2ASH8 and C4AH13 and C4ACH11, while at 90 d of
curing there is a significant decrease of the presence of
calcium aluminate hydrates, and a predominance of
stratlingite. On the contrary, the MK02 paste reveals a
highly crystalline microstructure, with predominance of
calcite. Up to 90 d of curing time, no hydrogarnet pre-
sence is detected, in any of the blended pastes studied. A
decrease in the amount of C4AH13 and C4ACH11 with
curing time is verified, not followed by a decrease in
microstructural porosity, that may influence for mixtures
with low MK content a decrease in the mechanical
resistance of the blended lime/MK mixes. The results
show the formation of hydration products that may
confer mechanical resistance to lime/MK mortars.
Acknowledgments
The authors wish to acknowledge the Fundação para
a Ciência e Tecnologia (FCT) for the financial support
under project METACAL (PTDC/ECM/100431/2008)
and to the enterprises Lusical and IMERYS for the
supply of the lime and metakaolin used in this work.
5 REFERENCES
1 C. S. Poon, L. Lam, C. S. Kou, Wong, Y. L., Wong, Ron, Rate of
pozzolanic reaction in high-performance cement pastes, Cement and
Concrete Research, 31 (2001), 1301–1306
2 S. S. Sabir, S. Wild, J. Bai, Metakaolin and calcined clays as
pozzolans for concrete: a review, CEm. Concr. Compos., 23 (2001),
441–454
3 J. Cabrera, M. Frías, Influence of MK on the reaction kinetics in
MK/lime and MK-blended cement systems at 20 °C, Cement and
Concrete Research, 31 (2001), 519–527
4 P. S. De Silva, F. G. Glasser, Phase relation in the system
CaO-Al2O3-SiO2-H2O relevant to metakaolin-calcium hydroxide
hydration, Cem. Concr. Res., 23 (1993) 3, 627–639
5 M. Frías Rojas, Study of hydrated phases present in a MK-lime
system cured at 60 °C and 60 months of reaction, Cement and
Concrete Research, 36 (2006), 827–831
6 A. Bakolas, E. Aggelakopoulou, A. Moropoulou, S. Anagnosto-
poulou, Evaluation of pozzolanic activity and physico-mechanical
characteristics in metakaolin-lime pastes, Journal of Thermal
Analysis and Calorimetry, 84 (2006) 1, 157–163
7 A. Moropoulou, A. Bakolas, E. Aggelakopoulou, Evaluation of
pozzolanic activity of natural and artificial pozzolans by thermal
analysis, Thermochimica Acta, 420 (2004), 135–140
8 A. Gameiro, A. Santos Silva, R. Veiga, A. Velosa, Phase and micro-
structural characterization of lime-MK blended mixes, Proc. of the
VI International Materials Symposium MATERIAIS 2011, April
2011; 6 pages
A. L. GAMEIRO et al.: LIME-METAKAOLIN HYDRATION PRODUCTS: A MICROSCOPY ANALYSIS
148 Materiali in tehnologije / Materials and technology 46 (2012) 2, 145–148
S. RANDJELOVI] et al.: THE IMPACT OF DIE ANGLE ON TOOL LOADING IN THE PROCESS ...
THE IMPACT OF DIE ANGLE ON TOOL LOADING IN
THE PROCESS OF COLD EXTRUDING STEEL
VPLIV KOTA MATRICE NA OBREMENITEV ORODJA PRI
HLADNI EKSTRUZIJI JEKLA
Sa{a Randjelovi}1, Miodrag Mani}1, Miroslav Trajanovi}1, Mladomir Milutinovi}2,
Dejan Movrin2
1University of Ni{, Faculty of Mechanical Engineering, Serbia
2University of Novi Sad, Faculty of Technical Science, Serbia
sassa@masfak.ni.ac.rs
Prejem rokopisa – received: 2011-10-20; sprejem za objavo – accepted for publication: 2011-12-08
This paper presents an analysis of tool loading in the technology of the cold forward extrusion of steel. In the process of plastic
deformation it is necessary to know the contact stress as a prerequisite for a more accurate analysis of the stress and strain on the
internal structure of the continuum. In this way, accurate boundary conditions at the contact surfaces are obtained for the
achieved conditions of deformation, which represent the starting values for generating numerical approximations of the
plasticity parameter changes within the deformable volume. In the process of the forward extrusion of steel the workpiece
material is exposed to all-round pressure during the entire process. Due to the high surface pressure at the head of the punch and
the solid walls of the die, the material flows in the direction of the opening of the exchangeable conical surfaces of the die.
During the extrusion process, the greatest resistance occurs in the direction of the axis displacement, i.e., the head punch, while
the walls of the tools suffer considerably smaller loads. However, this has crucial importance for the accuracy and the quality of
the finished part.
Keywords: cold extrusion, angle die, contact stress, FEM
V ~lanku je opisana analiza obremenitve orodja pri hladni ekstruziji jekla v smeri naprej. Pri procesu plasti~ne deformacije je
treba poznati kontaktno napetost, ki je prvi pogoj za bolj natan~no analizo napetosti in deformacije na notranjo strukturo
kontinuuma. Tako je mogo~e dose~i natan~ne mejne razmere na kontaktnih povr{inah pri dolo~enih pogojih deformacije, ki so
za~etne vrednosti za generiranje numeri~nih pribli`kov spremembe parametrov plasti~nosti v deformabilnem volumnu. Pri
ekstruziji materiala v smeri naprej je pre{anec med celotnim procesom izpostavljen okoli{njemu pritisku. Zaradi visokega
povr{inskega pritiska na ~ela trna in trdne stene valjastega orodja material te~e v smeri premika osi, torej ~ela trna, medtem ko
je obremenitev stene orodja mnogo manj{a, vendar je zelo pomembna za natan~nost in kakovost iztiskanca.
Klju~ne besede: hladna ekstruzija, kot matrice, kotaktna napetost, FEM
1 INTRODUCTION
Analyses of the uni-directional extrusion process
have been made by many authors. First of all, this is a
process of volume deforming in which the workpiece
material is subjected to overall pressure throughout the
entire process. Due to the high surface pressure on the
extruder head and on the rigid matrix walls, the material
flows towards the opening on the conical matrix
surfaces. During the extrusion process the greatest forces
occur in the extrusion axis direction, i.e., on the punch
head and on the matrix walls.
Many examples from industry show, in a clear
manner, that realistic lifetime calculations for cold-
forging tools should be focused on the accurate
prediction of the number of load cycles until the
appreciable continual damage or the initiation of fatigue
cracks.
In addition, the results gained from previous investi-
gations show that lifetime predictions only based on
finite-element analyses are characterised by intolerable
inaccuracies1.
2 EXPERIMENTAL SET-UP
The satisfying experimental matrix rigidity and
annulling of high pressure in the radial direction can be
achieved in two ways. One of them is to increase the
matrix-wall thickness up to a certain limit, thus obtaining
the required matrix rigidity. The other solution is
installing the clamping-ring application, thus bringing
the matrix body into a pre-stress state of the opposite
sign with respect to the stresses occurring during the
extrusion process itself.
If, in the first case, we regard the receiver as a fairly
thick pipe of outer diameter D1, under high inner
pressure loading, then its walls are subjected to a radial
stress of pressure Rr and tangential extension stress Rt
whose greatest value is on inner receiver diameter D0.
R pr = − , R
a
a
p C pt =
+
−
⋅ = ⋅
2
2
1
1
(1)
where a = D1/D0.
By superposition these two stresses, according to the
plastic yield hypothesis, the following overall stress is
obtained:
Materiali in tehnologije / Materials and technology 46 (2012) 2, 149–154 149
UDK 621.77 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)149(2012)
R R R R R p C Cu r t r t= + − = ⋅ + +
2 2 21 (2)
Assuming that, for instance, the outer receiver
diameter is four times larger than the inner diameter, i.e.,
a = 4, we find that Ru = 1.85  p.
Along with the further increase of the outer diameter,
the overall stress reduction in the receiver walls is not
adequate. In the case when a = 10, which is not quite
justified in real exploitation conditions, C = 1.02 and Ru
= 1,75p is obtained, i.e., the overall stress reduction is
9.5 % along with an outer receiver dimension increase of
2.5 times. Since the receiver must remain in the elasticity
range, i.e., since there must be no plastic deformation (Ru
< Re) throughout the process, we can approximately
determine the greatest value of the working pressure in
the receiver made of alloyed tool steel submitted to heat
treatment:
Re  2 000 MPa, pmax = 1 100 MPa
The above-presented analysis has been used as the
basis for experimental tool design for the forward-
extrusion procedure. Unlike the exploitation tools, these
tools, Figure 1, enable an extrusion force measurement
on the punch as well as that of the radial forces in the
lower part of the tool on the receiver wall.
The extrusion force on the punch can be measured in
many ways. In this investigation, the choice of measure-
ment procedure is made by means of the universal force
transducer (2 000 kN), which can be placed either in the
upper or the lower part of the tool, as is our case, through
which the overall loading is transmitted along the
extrusion axis. For the measurement of the radial forces,
the measuring pin load-cell method (Figure 2), ascribed
Plancak et al.2, is the most suitable for this kind of plastic
metal-deforming process.
The loading due to the contact between the work-
piece and three measuring pins is transmitted to the
measuring capsule (Figure 3) placed on the outer matrix
wall.
The extension of the measuring wall of the capsule
which is 1.5 mm thick, due to the radial forces in the
receiver tool, actually yields the loading magnitude on
the matrix walls. On each capsule wall there are two
measuring bands HBM (measurement and compensation
ones) glued and joined into a semi-bridge (Wheatstone)
necessary to carry out their calibration with a known
loading and thus set up a relation between the capsule
wall elongation and the force being transmitted.
In order to obtain complete information about the
magnitude and the kind of stress throughout the
extrusion process, the measuring pins distribution is
defined by the matrix geometry as well as the workpiece
size. For these reasons, there are three measuring pins
S. RANDJELOVI] et al.: THE IMPACT OF DIE ANGLE ON TOOL LOADING IN THE PROCESS ...
150 Materiali in tehnologije / Materials and technology 46 (2012) 2, 149–154
Figure 4: Experimental tool, receiver with three measuring pin
load-cells
Slika 4: Eksperimentano orodje: prejemnik s tremi merilnimi trni
Figure 2: Measuring pin load-cell for radial forces
Slika 2: Merilna celica s trni za radialne sile
Figure 1: Experimental tool, receiver
Slika 1: Eksperimentalno orodje: prejemnik
Figure 3: Measuring capsule
Slika 3: Merilni valj
placed in radial way in the extrusion matrix body at an
angle of 120° to the measuring pin that is in direct
contact with the workpiece during the extrusion process.
In order to obtain complete information about the
loading magnitude during the process, the measuring
pins are placed at various heights in the material receiver
(Figure 4).
In order to provide for variants of the presented tool
solution, the very deformation focus (conical matrix
part) and calibration zone are made in special dies
introduced into the matrix body. The characteristic
conical tool surface, on the given matrices, has three
values of the angle, i.e., 60°, 90° and 120° (Figure 5),
which will directly affect both the extrusion forces and
the radial forces in the tool.
The material of the workpiece was low-carbon steel
Ck 10 (DIN) for cold forging3. The flow stress at room
temperature was modeled by the strain hardening
function K = 285 + 539.7·0.304 MPa, obtained from the
Rastegaev compression test according reference3. The
Young’s modulus and the Poisson’s ratio were 210 GPa
and 0.3, respectively. FEM analysis was performed on a
constant friction model, with the friction factor m = 0.5
K.4
3 SIMULATION RESULTS AND FEM ANALYSES
The coordinate system for presenting the results
comprises a time x-axis with the number of readings
equal to 800 with a step, the time interval between two
signals of 0.003 s as well as the y-axis with the force in
kN on the x-axis one part is singled where the workpiece
extrusion process and the extruded-part ejection process
are marked.
The measurement results show a certain regularity
(Figures 6, 7 and 8) and similar effects can be noticed in
all the extrusion processes. The force upon the punch
shows, before the very extrusion process, a marked
instability as well as a very high increase, after which it
drops to its minimal value throughout the overall
deforming process. The force instability, as well as its
increase, can be explained by considering the most
favorable initial position of the workpiece in the tool and
by the needed – relatively high – force for the very
beginning of the material flow on the matrix insert walls.
A marked force drop shows that the first phase is
completed, that the material filled up the input part of the
cone; after that the force starts to increase rapidly to a
maximum, after which the material flow on the conical
tool parts begins. The maximum value has a marked
increase along with the matrix-angle increase.
In all diagrams the radial forces are denoted by the
relative height at which they were measured. Namely, as
there is a difference regarding the height of the place at
which the pin contact to the workpiece; at every 2 mm
from the upper edge of the matrix insert, at mutual
matrix angles of 120°, there is a different increase in the
given forces recorded. In the beginning of the process,
the maximum radial force is achieved at the highest pin
with respect to the matrix insert at the moment when the
extrusion force achieves its maximum value, i.e., when
the initial unstable phase is completed and the material
S. RANDJELOVI] et al.: THE IMPACT OF DIE ANGLE ON TOOL LOADING IN THE PROCESS ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 149–154 151
Figure 7: Force distribution at a die angle of 90°
Slika 7: Sile pri kotu matrice 90°
Figure 5: Extrusion matrices with various cone angles
Slika 5: Ekstruzijske matrice z razli~nimi koti
Figure 6: Force distribution at a die angle of 60°
Slika 6: Sile pri kotu matrice 60°
flowness has already started. This is explained by the
very workpiece itself at this particular moment
(barrel-like form in the receiver) and immediately after
it, when the workpiece material filled up the whole
material receiver volume and when its "maximum
diameter slides" along the matrix walls. After reaching
its maximum value as well as its retention, this force
drops to almost a zero value. A less distinct maximum is
reached by the radial force on the second pin, at the
height of 4 mm from the die insert, but in an almost
identical period of time when the first radial force
reaches its maximum. The lowest pin with respect to the
matrix pickup records almost the same magnitude of
radial force as that on the second pickup, but it can
clearly be seen that it preserves this value, with some
slight decline, until the end of the extrusion process,
since the non-extruded volume of the material from the
receiver also remains at its height.
The extrusion force on the extruder has values
ranging from 560 kN to 600 kN in the matrix with the
smallest cone angle of 60°, i.e., 660 kN to 690 kN for the
matrix with a cone angle of 90°, or to 720–790 kN in the
S. RANDJELOVI] et al.: THE IMPACT OF DIE ANGLE ON TOOL LOADING IN THE PROCESS ...
152 Materiali in tehnologije / Materials and technology 46 (2012) 2, 149–154
Figure 11: Contact stress at a die angle of 90°: a) on the punch and b)
on the receiver wall in the radial direction
Slika 11: Kontaktne napetosti pri kotu 90°: a) na batu in b) na steni
prejemnika v radialni smeri
Figure 9: Field of efective plastic strain at three die angles: 60°, 90°
and 120°
Slika 9: Polje efektivnih plasti~nih deformacij pri treh kotih matrice
60°, 90° in 120°
Figure 8: Force distribution at a die angle of 120°
Slika 8: Sile pri kotu matrice 120°
Figure 10: Contact stress at a die angle of 60°: a) on the punch and b)
on the receiver wall in the radial direction
Slika 10: Kontaktne napetosti pri kotu 60°: a) na batu in b) na steni
prejemnika v radialni smeri
matrix with a cone angle of 120°. Reduced to the
cross-sectional area of the workpiece, over which this
force is transmitted, working pressures of 1900–2500
N/mm2 occur in the extrusion process. The radial force
on the matrix wall, depending on the cone angle, moves
in the interval from 50 kN to 230 kN for 60°, i.e., from
70 kN to 230 kN for 90° and 100 kN to 350 kN for 120°.
The force increase follows the height of the measure-
ment place on the receiver wall. In the radial direction
there are considerably smaller pressures and they move
within the limits from 945 N/mm2 to 1742 N/mm2.
On the right-hand side of all the diagrams, there is a
marked instability of all the forces associated with the
ejection phase of the extruded piece from the die and the
matrix itself.
A numerical 2D Finite-Element Method (FEM) ana-
lysis of the investigated models of forward extrusion was
performed using Simufact. Forming 10.0 software
package5. The commercial FEM package enabled the
entire forming process to be simulated, while simul-
taneously predicting a large number of parameters at
both the workpiece and the tool6. In this paper the FEM
is employed to predict stress-strain state, the forming
load and the geometry of the workpiece. To simulate the
process, a model of elastic-plastic material for the work-
piece was chosen, as the die and punch are considered to
be rigid bodies. Due to the axial-symmetry of the
deformation process, only one-half of the workpiece was
modeled. The displacement of the punch is defined to be
16 mm and the punch velocity as 0.1 mm/s.
The workpiece model was initially meshed with
advancing front quad elements with a size of 0.3 mm, the
total number of which was 2 220. In the simulation, the
remeshing of the starting elements was executed in the
most highly deformed zones of the workpiece. The
remeshing procedure was performed at every five
increments in order to minimize the effect of the tool
penetration through the elements due to large workpiece
deformations.
Stress-strain components within the workpiece vol-
ume obtained by the FE analysis are shown in Figures 9,
10, 11 and 12. It is significant that the stress-strain state
is very heterogeneous7.
The FEA-simulation-predicted load-stroke diagram
closely resembles the ones obtained experimentally
(Figure 13). Initially, the load increases quickly up to
1.8 mm (120°), 2.2 mm (90°) and 3.4 mm (60°) of the
punch stroke. As the punch stroke progresses further, the
load continues to increase gradually. The final phase is
marked by a noticeable load decrease.
4 CONCLUSIONS
The analysis of the tool loading points to the order of
the loading magnitude as well as the force effect
distribution over time during the process. The tool
loading is of a variable character and the order of
magnitude directly depends on the workpiece diameter,
the finished part diameter and the extrusion angle in the
deformation focus. The measurement itself aims at
pointing to the loading magnitude at the contact surfaces,
while, in further work, this could serve as input data for
solving the stress-distribution equations with respect to
the volume of the extruded part.
A special set of interchangeable tools, with three
different angle dies, condition different values and levels
of change in the radial force, as well as the degree of
damage to the tools, which directly affect its service life,
S. RANDJELOVI] et al.: THE IMPACT OF DIE ANGLE ON TOOL LOADING IN THE PROCESS ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 149–154 153
Figure 13: Loading distribution during the working stroke
Slika 13: Razporeditev obremenitev med delovnim ciklom
Figure 12: Contact stress at a die angle of 120°: a) on the punch and
b) on the receiver wall in the radial direction
Slika 12: Kontaktne napetosti pri kotu 120°: a) na batu in b) na steni
prejemnika v radialni smeri
is used. The change of radial force in time points to the
changeable shape of the workpiece during the process
and to the surface of contact within the receiver tool.
Acknowledgement
This paper is part of the project III41017 Virtual
human osteoarticular system and its application in
preclinical and clinical practice, funded by the Ministry
of Education and Science of the Republic of Serbia
(http://vihos.masfak.ni.ac.rs).
5 REFERENCES
1 B. Falk, U. Engel, M. Geiger, Fundamental aspects for the evolution
of the fatigue behaviour of cold forging tools, Journal of Materials
Processing Technology, (2001), 158–164
2 M. Plancak, A. N. Bramley, F. H. Osman, Some observations on
contact stress measurement by pin load cell in bulk metal forming,
Journal of Materials Processing Technology, (1996), 339–342
3 Steels for cold forging – their behaviour and selection, ICFG
document No.11/01, ISBN 3-87525-148-2, 2001
4 K. H. Jung, H. C. Lee, J. S. Ajiboye and Y. T. Im, Characterization of
Frictional Behavior in Cold Forging, Tribology Letters, 37 (2010),
353–359
5 Simufact.Forming. 10.0. Documentation
6 J. Kusiak, M. Pietrzyk, J. L. Chenot, Die Shape Design and Eva-
luation of Microstructure Control in the Closed die Axisymetric
Forging by Using FORGE2 Program, ISIJ International, 34 (1994) 9,
755–760
7 Z. Peng, T. Sheppard, A study on material flow in isothermal
extrusion by FEM simulation, Modelling and simulation in materials
science and engineering, 12 (2004), 745–763
S. RANDJELOVI] et al.: THE IMPACT OF DIE ANGLE ON TOOL LOADING IN THE PROCESS ...
154 Materiali in tehnologije / Materials and technology 46 (2012) 2, 149–154
J. [TÌTINA et al.: FINAL-STRUCTURE PREDICTION OF CONTINUOUSLY CAST BILLETS
FINAL-STRUCTURE PREDICTION OF CONTINUOUSLY
CAST BILLETS
NAPOVED KON^NE MIKROSTRUKTURE KONTINUIRNO ULITIH
GREDIC
Josef [tìtina, Lubomír Klime{, Tomá{ Mauder, Franti{ek Kavi~ka
Brno University of Technology, Technická 2, 616 69 Brno, Czech Republic
stetina@fme.vutbr.cz
Prejem rokopisa – received: 2011-10-20; sprejem za objavo – accepted for publication: 2012-01-09
In steel production, controlling and monitoring quality, grade and structure of final steel products are very important issues. It
has been shown that the temperature distribution, the magnitude of temperature gradients, as well as the cooling strategy during
the continuous steel casting have a significant impact on material properties, the structure and any defect formation of cast
products. The paper describes an accurate computational tool intended for investigating the transient phenomena in continuously
cast billets, for developing the caster control techniques and also for determining the optimum cooling strategy in order to meet
all quality requirements. The numerical model of the temperature field is based on the finite-difference implementation of the
3D energy-balance equation using the enthalpy approach. This allows us to analyse the temperature field along the entire cast
billet. Since the steel billets are produced constantly 24 hours per day, the transient temperature field is being computed in a
non-stop trial run. It enables us to monitor and investigate the formation of the temperature field in real time within the mould,
as well as the secondary and tertiary cooling zones, where the observed information can be immediately utilized for the
caster-control optimization with respect to the whole machine or just an individual part. The application of the presented model
is demonstrated with two examples including the steelworks in Tøinec, Czech Republic, and in Podbrezová, Slovakia. To
consider different operational conditions, the influences of the secondary-cooling setting on the surface and the inner defects
formation, and on the final structure of the 150 × 150 mm billet are also discussed.
Keywords: concast billet, numerical model, solidification
Pri proizvodnji jekla je pomembno izvajanje kontrole kvalitete, vrste in mikrostrukture kon~nih proizvodov. Prikazano je, da
ima razporeditev temperature, razpon temperaturnega gradienta, kot tudi strategija ohlajanja med kontinuirnim ulivanjem
pomemben vpliv na lastnosti materiala, mikrostrukturo in mo`nost nastanka napak v litem proizvodu. V ~lanku je predstavljeno
natan~no ra~unalni{ko orodje za preiskave prehodnih pojavov v kontinuirno uliti gredici. Orodje je namenjeno razvoju kontrolne
tehnike in tudi za dolo~anje optimalne strategije ohlajanja za doseganje zahtevane kakovosti. Numeri~ni model temperaturnega
polja temelji na uporabi kon~nih diferenc 3D-ena~be energijskega ravnote`ja z uporabo entalpije. Omogo~a analizo
temperaturnega polja vzdol` celotne lite gredice. Ker je proizvodnja gredic stalna, 24 ur na dan, je bilo prehodno temperaturno
polje izra~unano za neprekinjeno preskusno obratovanje. Omogo~ena je kontrola in preiskava nastanka temperaturnega polja v
realnem ~asu v kokili, v sekundarni in terciarni hladilni coni. Ugotovljena informacija se lahko neposredno uporabi za
optimiranje celotne livne naprave ali pa samo posameznega dela. Uporaba predlaganega modela je prikazana za dva primera iz
`elezarne Tøinec na ^e{kem in `elezarne Podbrezová na Slova{kem. Prikazan je tudi vpliv nastavitve sekundarnega ohlajanja na
nastanek povr{inskih in notranjih napak na gredici 150 mm × 150 mm pri razli~nih obratovalnih razmerah.
Klju~ne besede: kontinuirno ulite gredice, numeri~ni model, strjevanje
1 NUMERICAL MODEL OF THE TEMPE-
RATURE FIELD OF A CONCAST BILLET
The presented in-house model of the transient
temperature field of the blank from a billet caster
(Figure 1) is unique. In addition to being entirely 3D, it
can operate in real time. It is possible to adapt its
universal code and use it for any billet caster. The
numerical model covers the temperature field of the
entire length of a blank (i.e., from the meniscus inside
the mould all the way down to the cutting torch) with up
to one million nodes.
The solidification and cooling of a blank and the
simultaneous heating of the mould is a case of the 3D
transient heat and mass transfer in a system comprising a
blank-mould ambient and, after leaving the mould, only
a blank ambient1. If mass transfer is neglected and if
only conduction is considered as being decisive, then the
heating up of the mould is described by the Fourier
Equation (1). The solidification and the cooling of a
blank is described by the Fourier-Kirchhoff Equation (2),
Materiali in tehnologije / Materials and technology 46 (2012) 2, 155–160 155
Figure 1: A billet caster
Slika 1: Shematski prikaz kontinuirnega ulivanja gredic
UDK 621.74.047:536.421.4:519.61/.64 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)155(2012)
which contains the components describing the heat flow
from the melt flowing with a velocity v, and the
component including the internal source of latent heats
of phase or structural changes Qsource .


⋅ = ⎛⎝
⎜ ⎞
⎠
⎟ +
⎛
⎝
⎜
⎞
⎠
⎟ + ⎛⎝
⎜ ⎞c
T
x
k
T
x y
k
T
y z
k
T
z
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂ ⎠
⎟ (1)


⋅ = ⎛⎝
⎜ ⎞
⎠
⎟ +
⎛
⎝
⎜
⎞
⎠
⎟ + ⎛⎝
⎜ ⎞c
T
x
k
T
x y
k
T
y z
k
T
z
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂ ⎠
⎟ +
+ ⋅ + +
⎛
⎝
⎜
⎞
⎠
⎟ +c u
T
x
v
T
y
w
T
z
Q
∂
∂
∂
∂
∂
∂

source (2)
Figure 2 shows the temperature balance of an
elementary volume representing the general node of the
mesh (i,j,k) inside the mould. The heat conductivities
VX, VY and VZ along the main axes are:
VX k
A
xi j k i
x
, , = Δ
VX k
A
xi j k i
x
− −=1 1, , Δ
(3a)
VY k
A
yi j k j
y
, , = Δ
VY k
A
yi j k j
y
, ,− −=1 1 Δ
(3b)
VZ k
A
zi j k k
z
, , = Δ
VZ k
A
zi j k k
z
, , − −=1 1 Δ
(3c)
The heat flows QX, QY and QZ through the elemen-
tary volume along the main axes are:
QX VX T Ti j k i j k i j k= −+, , , ,
( )
, ,
( )( )1
  (4a)
QY VY T Ti i j k i j k i j k= −+, , , ,
( )
, ,
( )( )1
  (4b)
QX VX T Ti j k i j k i j k1 1 1= −− −, , , ,
( )
, ,
( )( )  (4c)
QY VY T Ti j k i j k i j k11 1 1= −− −, , , ,
( )
, ,
( )( )  (4d)
QZ VZ T Ti j i j k i j k i j k, , , , ,
( )
, ,
( )( )= −+1
  (4e)
QZ VZ T Ti j i j k i j k i j k1 1 1, , , , ,
( )
, ,
( )( )= −− −
  (4f)
The temperature balance of the general node is:
( ), , , ,QZ QZ QY QY QX QXi j i j i j i j1 1 1+ + + + + =
=
   

 
x y z c
T Ti j k i j k
⋅ ⋅ ⋅ ⋅
−( ), ,
( )
, ,
( ) (5)
where the right-hand side expresses the accumulation
(or loss) of heat in the node i,j,k during the time step .
The unknown temperature of the general node of the
mesh inside the mould in the following instant ( + )
is therefore given by the explicit formula:
T T QZ QZ QY QYi j k i j k i j i j i j i j, ,
( )
, ,
( )
, , , ,(
 = + + + + +1 1 QX QX1+ ⋅)
⋅ ⋅ ⋅ ⋅ ⋅

   x y z c
(6)
The temperature field of the blank passing through a
radial caster of a large radius can be simplified by the
Fourier-Kirchhoff equation where only the vz component
of the velocity is considered. Equation (2) is therefore
reduced to:


⋅ = ⎛⎝
⎜ ⎞
⎠
⎟ +
⎛
⎝
⎜
⎞
⎠
⎟ + ⎛⎝
⎜ ⎞c
T
x
k
T
x y
k
T
y z
k
T
z
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂ ⎠
⎟ +
+ ⋅ ⋅ +c w
T
z
Q
∂
∂

source (7)
Equation (7) must cover the temperature field of the
blank in all three stages: above the liquidus temperature
(i.e., the melt), in the interval between the liquidus and
solidus temperatures (i.e., the so-called mushy zone) and
beneath the solidus temperature (i.e., the solid phase). It
is therefore convenient to introduce the thermodynamic
function of specific volume enthalpy Hv = cT, which is
dependent on temperature, and also includes the phase
and structural heats (Figure 3).
Heat conductivity k, specific heat capacity c and
density  are thermophysical properties that are also
functions of temperature. Equation (7) therefore takes on
the form:
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
H
x
k
T
x y
k
T
y z
k
T
z
v

= ⎛⎝
⎜ ⎞
⎠
⎟ +
⎛
⎝
⎜
⎞
⎠
⎟ + ⎛⎝
⎜ ⎞
⎠
⎟ +w
H
z
v∂
∂
(8)
The heat balance of the elementary node is:
( ), , , ,QZ QZ QY QY QX QXi j i j i j i j1 1 1+ + + + + =
=
  

 
x y z
T Tvi j k vi j k
⋅ ⋅
−( ), ,
( )
, ,
( ) (9)
J. [TÌTINA et al.: FINAL-STRUCTURE PREDICTION OF CONTINUOUSLY CAST BILLETS
156 Materiali in tehnologije / Materials and technology 46 (2012) 2, 155–160
Figure 3: The enthalpy function for steel showing the phase and struc-
tural changes
Slika 3: Entalpijska funkcija za jeklo s fazno in strukturno premeno
Figure 2: Heat balance of the general node of the mesh
Slika 2: Diagram toplotnega ravnovesja v splo{ni to~ki mre`e
where the heat flow QZi,j must now also include the
enthalpy of the incoming volume of melt:
( ( ), , , ,
( )
, ,
( )
, ,
(QZ QZ T T A w Hi j i j i j k i j k z vi j k1 1= − − ⋅ ⋅+
   ) (10)
The unknown enthalpy of the general node of the
blank in the following instant ( + ) is given by the
explicit formula, similar to Equation (6):
H H QZ QZ QY QYvi j k vi j k i j i j i j i, ,
( )
, ,
( )
, , , ,(
 = + + + +1 1 j QX QX+ + ⋅1 )
⋅ ⋅ ⋅

  x y z
(11)
Figure 3 indicates how the temperature model for the
calculated enthalpy in Equation (11) determines the
unknown temperature.
The next task is to choose a suitable coordinate
system and a mesh. This paper deals with the symme-
trical half of one cross-section of a blank from the
meniscus inside the mould down to the cutting torch.
The origin of the coordinate system is positioned on the
small radius in the centre of the width (Figure 4). This
enables all coordinates to be positive, which facilitates
the software programming. In the region of the radius,
the Cartesian coordinates are transformed into the
cylindrical ones (i.e., y is the radius and z is the angle).
The mesh is generated automatically and the model
supports all densities of the mesh introduced in Figure 4.
All the results presented in this paper are based on a
mesh of 573,594 nodes (11 in the x-direction, 21 in the
y-direction and 1861 in the z-direction) and a 7.5 mm ×
7.5 mm × 15 mm elementary volume.
All thermodynamic properties of the cast steel,
dependent on its chemical composition and the cooling
rate, enter the calculation as functions of temperature2.
This is therefore a significantly non-linear task because,
even with the boundary conditions, their dependence on
the surface temperature of the blank is considered here.
Regarding the fact that the task can be considered
symmetrical along the axis (Figure 4), it is sufficient to
deal with only one half of the cross-section. The
boundary conditions are therefore as follows:
1. T T= cast the level of the steel (12a)
2. − =k
T
n
∂
∂
0 the plane of symmetry (12b)
3. − = ⋅ −k
T
n
htc T T
∂
∂
( )surface amb inside the mould (12c)
4. − = ⋅ − + −k
T
n
htc T T T T
∂
∂
( ) ( )surface amb surface
4
amb
4
within the secondary and tertiary zones (12d)
5. − =k
T
z
q
∂
∂
 beneath the rollers (12e)
The boundary conditions are divided into the area of
the mould, the area of the secondary cooling and the area
of the tertiary cooling.
The initial condition for the investigation is the
setting of the temperature in individual points of the
mesh. A suitable temperature is the highest possible
temperature, i.e., the pouring temperature. The explicit
difference method is used for solving this problem. The
principle of this method is that the stability of the
calculation is dependent on the magnitude of the time
step. The model has incorporated a method for adapting
the time step, i.e., the time step entered by the operator is
merely a recommendation and the software modifies it
throughout the calculation.
2 HEAT TRANSFER COEFFICIENT ALONG
THE ENTIRE CASTER
The cooling by the water nozzles has the main
influence and it is therefore necessary to devote much
attention to establishing the relevant heat-transfer
coefficient of the forced convection. Commercially sold
models of the temperature field describe the heat-transfer
coefficient beneath the nozzles as a function of the
incident quantity of water per unit area. They are based
on various empirical relationships. This procedure is
undesirable. The model discussed in this paper obtains
its heat-transfer coefficients from the measurements of
spraying characteristics of all nozzles used by the caster
on the so-called hot plate in an experimental laboratory3,4
and for a sufficient range of operational pressures of
water, as well as for a sufficient range of casting speeds
of the blank (i.e., casting speed). This approach repre-
sents a unique combination of an experimental measure-
ment in a laboratory and a numerical model for
calculating the non-linear boundary conditions beneath
the cooling nozzle.
Figure 5 presents the measured values of the heat-
transfer coefficients processed by the temperature-model
software. For the nozzle configuration, there is a graph
of the heat-transfer coefficient beneath the nozzle. These
graphs are plotted for a surface temperature of 1000 °C.
The resultant heat-transfer coefficient is determined
by adding up the partial coefficients. This basically
entails the total heat-transfer coefficient because even
radiation, with the introduction of the "reduced heat-
transfer coefficient from radiation", was converted to
convection. On the areas of the blank, where the natural
convection and radiation occur, the total coefficient is
given by the sum of the reduced coefficient from radia-
tion and the coefficient of the actual natural convection.
J. [TÌTINA et al.: FINAL-STRUCTURE PREDICTION OF CONTINUOUSLY CAST BILLETS
Materiali in tehnologije / Materials and technology 46 (2012) 2, 155–160 157
Figure 4: The mesh and the definition of the coordinate system
Slika 4: Mre`a in opredelitev koordinatnega sistema
In the area beneath the nozzle, the resultant heat-transfer
coefficient is obtained as the sum of the forced-convec-
tion coefficient gained from the laboratory-temperature
measurement and the reduced heat-transfer coefficient
from radiation2.
On a specific caster, the nozzles of the secondary
cooling are divided into several independent regulation
zones, enabling the formation of the temperature field of
the blank. Figure 6 shows the 6 individual regulation
zones and Figure 7 shows the courses of the resultant
heat-transfer coefficients along the small radius of the
billet caster. On a specific caster, the nozzles of the
secondary cooling are divided into several independent
regulation zones (I, IIA, IIB, IIIA, IIIB and IV), enabling
the formation of the temperature field of the blank.5
3 EFFECT OF THE SECONDARY COOLING
The setting of the secondary cooling and its
optimization is a very complicated problem6. In a real
operation specific intensity of cooling is characterized by
the consumption of the cooling water per 1 kg of cast
steel. On the basis of the curves indicating various con-
sumptions of cooling water per unit of the mass of cast
steel varying from 9 L/kg to 18 L/kg, the temperature of
the blank was calculated and presented in Figure 8.
These cooling curves are established for the given caster
referring to six cooling zones I, IIA, IIB, IIIA, IIIB and
IV.
4 ON-LINE MODEL OF THE TEMPERATURE
FIELD
A temperature model can be considered to be
successfully implemented if it is integrated into the
existing information and control systems of a caster. The
users (i.e., technologists) can record the real-time data
from the on-line model into their off-line model of the
temperature field, carry out any necessary changes in the
input parameters (e.g., alter the secondary cooling or the
casting speed). After a simulation on the off-line model,
it is possible to determine how the temperature field will
change after the implementation of the changes. Another
application of the off-line version is in the occurrence of
defects on/in the actual slab or sheet steel. The user can
J. [TÌTINA et al.: FINAL-STRUCTURE PREDICTION OF CONTINUOUSLY CAST BILLETS
158 Materiali in tehnologije / Materials and technology 46 (2012) 2, 155–160
Figure 5: The heat-transfer coefficient for the 5065l nozzle: a) flow
through a nozzle at 5.17 L/min, b) flow through a nozzle at 10.00
L/min
Slika 5: Koeficient prehoda toplote za {obo 50651: a) pretok skozi eno
{obo pri 5,17 L/min, b) pretok skozi eno {obo pri 10,00 L/min
Figure 7: The resultant heat-transfer coefficient along the small radius
of the billet caster
Slika 7: Dobljeni koeficienti prehoda toplote vzdol` notranjega radija
naprave za ulivanje gredic
Figure 6: Positions of the nozzles along the billet caster in 6 indivi-
dual zones
Slika 6: Pozicije hladilnih {ob vzdol` naprave za ulivanje gredic v 6
obmo~jih
read the temperature field from the archive server using
the dynamic model and – using the off-line model –
analyse any likely causes of defects and prepare the
necessary measures for the defects never to occur again.
The off-line model will (in future) enable the reading of
quantities and their dependences from the application
server and, using statistical methods and the relation-
ships among these quantities and defects, will look for
the cause in the original temperature field of the con-
casting from a specific melt. However, this will be the
task of the mathematical-stochastic prediction model.
Figure 9 compares the average values of the
measured surface temperatures in the same points.
Comparing the absolute values, it is possible to see that
there are long intervals where the deviation is significant
and, on the other hand, there are intervals where the
values are identical.
5 CONCLUSIONS
This paper presents a 3D numerical model of the
temperature field (for concasting of steel) in the form of
an in-house software that has been implemented in the
operation of TØINECKÉ @ELEZÁRNY, Czech Republic
and in Podbrezová, Slovakia. The model deals with the
main thermodynamic transfer phenomena during the
solidification of concasting.
Our analysis proved the usefulness of the model for
real applications, as well as the reliability and robustness
of the used numerical methods and other software. The
model has been applied in the calculation and setting of
the constants of the caster control system, including the
simulation of the caster operation under non-standard
situations (e.g., partial failures of the secondary cooling
during an unexpected slowing down of the casting), in
the planned maintenance of the machine or its structural
improvements, in the utilization of the information that
helps the operator to make spontaneous changes to the
control of the machine, in the utilization of monitoring
and controlling the quality, in the direct control of the
casting speed and in the flow of water in individual zones
of the secondary cooling of the prediction system7.
Acknowledgments
This analysis was conducted by using a program
devised within the framework of the project GA CR No.
106/08/0606, 106/09/0940 and P107/11/1566, NETME
centre – New Technologies for Mechanical Engineering
CZ.1.05/2.1.00/01.0002, Specific research BUT
BD13102003 and the co-author, the holder of Brno PhD
Talent Financial Aid sponsored by Brno City Municipa-
lity.
Nomenclature
A/m2 – area
c/(J/kg K) – specific heat capacity
htc/(W/m2 K) – heat transfer coefficient
Hv/(J/m3) – volume enthalpy
k/(W/mK) – heat conductivity
T/K – temperature
Tamb/K – ambient temperature
Tcast/K – melt temperature
Tsurface/K – temperature in unbending part
J. [TÌTINA et al.: FINAL-STRUCTURE PREDICTION OF CONTINUOUSLY CAST BILLETS
Materiali in tehnologije / Materials and technology 46 (2012) 2, 155–160 159
Figure 9: The measured and calculated temperatures of a billet
Slika 9: Izmerjene in izra~unane temperature gredice
Figure 8: A comparison of the temperature fields with different
intensities of secondary cooling: a) cooling curve 9 L/kg, b) cooling
curve 18 L/kg
Slika 8: Primerjava temperaturnih polj pri razli~ni intenzivnosti
sekundarnega hlajenja: a) krivulja ohlajanja pri 9 L/kg, b) krivulja
ohlajanja pri 18 L/kg
q/(W/m2) – specific heat flow
QX/W, QY/W, QZ/W – heat flows
Qsource /(W/m3) – internal heat source
x/m, y/m, z/m – axes in given direction
u/(m/s), v/(m/s), w/(m/s) – casting speed in given direc-
tion
VX/(W/K), VY/(W/K), VZ/(W/k) – heat conductivity
/(kg/m3) – density
/( W/m2 K4) – Stefan-Boltzmann constant
 – emissivity
/s – time
6 REFERENCES
1 J. K. Brimacombe, The Challenge of Quality in Continuous Casting
Process, Metallurgical and Materials Trans. B, 30B (1999), 553–566
2 J. Miettinen, S. Louhenkilpi, J. Laine, Solidification analysis
package IDS. Proceeding of General COST 512 Workshop on
Modelling in Materials Science and Processing, M. Rappaz and M.
Kedro eds., ECSC-EC-EAEC, Brussels, Luxembourg, 1996
3 M. Prihoda, J. Molinek, R. Pyszko, M. Velicka, M. Vaculík, J. Burda,
Heat Transfer during Cooling of Hot Surfaces by Water Nozzles,
Metalurgija = Metallurgy, 48 (2009) 4, 235–238
4 J. Horsky, M. Raudensky, Measurement of Heat transfer
Characteristics of Secondary Cooling in Continuous Casting. In
Metal 2005. Ostrava: TANGER, 2006, 1–8
5 A. Richard, Kai Liu Harding, Ch. Beckermann, A transient
simulation and dynamic spray cooling control model for continuous
steel casting, Metallurgical and materials transactions, 34B (2003),
297–302
6 F. Kavi~ka, J. Stetina, B. Sekanina, K. Stransky, J. Dobrovska, J.
Heger, The optimization of a concasting technology by two
numerical models, Journal of Materials Processing Technology, 185
(2007) 1–3, 152
7 B. G. Thomas, R. J. O’Malley, D. T. Stone, Measurement of tempe-
rature, solidification, and microstructure in a continuous cast thin
slab. Paper presented at Modeling of Casting, Welding and Advanced
Solidification Processes VIII, San Diego, CA, TMS 1998
J. [TÌTINA et al.: FINAL-STRUCTURE PREDICTION OF CONTINUOUSLY CAST BILLETS
160 Materiali in tehnologije / Materials and technology 46 (2012) 2, 155–160
S. AVDIAJ, B. ERJAVEC: OUTGASSING OF HYDROGEN FROM A STAINLESS STEEL VACUUM CHAMBER
OUTGASSING OF HYDROGEN FROM A STAINLESS
STEEL VACUUM CHAMBER
RAZPLINJEVANJE VODIKA IZ NERJAVNEGA JEKLA
Sefer Avdiaj1,3, Bojan Erjavec2
1University of Prishtina, Faculty of Natural Sciences and Mathematics, Mother Teresa Av. 3, Prishtina 10000, Kosova
2Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
3Lotri~, d.o.o., Selca 163, 4227 Selca, Slovenia
sefer.avdiaj@uni-pr.edu
Prejem rokopisa – received: 2012-02-07; sprejem za objavo – accepted for publication: 2012-02-22
Due to outgassing from the walls of vacuum calibration chambers, the generation of the calibration pressure in primary vacuum
calibration systems operating below 10–6 Pa becomes un-accurate. Austenitic stainless steel (SS) is the most common
construction material for ultrahigh vacuum (UHV) and extremely high vacuum (XHV) chambers. Hydrogen is the predominant
residual gas at very low pressures in SS vacuum systems, i.e., in the UHV and XHV range. Therefore, the reduction of the
hydrogen outgassing rate is the most challenging problem in achieving pressures in those ranges. In this paper, a vacuum
chamber with a wall thickness of 2.62 mm and made from AISI type 304 L SS was examined with the aim of mitigating the
outgassing rate. The heat treatments were carried out at 250 °C for 380 h and at 350 °C for 140 h. After baking at 250 °C for
380 h (corresponding to a dimensionless time Fo = 3.09), an outgassing rate q = 2.86 × 10–13 mbar L s–1 cm–2 was achieved at
room temperature (RT). This RT outgassing rate was further reduced to q = 5.7 × 10–14 mbar L s–1 cm–2 after baking for another
140 h at 350 °C (Fo = 8.66, resulting in a total dimensionless time Fo = 11.75).
Keywords: hydrogen outgassing, stainless steel, dimensionless time, pressure-rise method, throughput method, gas-flow
calibration coefficient
Zaradi razplinjevanja sten vakuumskih kalibracijskih posod postane vzpostavitev kalibracijskega tlaka v primarnih
kalibracijskih vakuumskih sistemih, delujo~ih v tla~nem obmo~ju manj{em od 10–6 Pa, nenatan~na. Avstenitno nerjavno jeklo je
najprimernej{i material za gradnjo ultra visokovakuumskih (UVV) in ekstremno visokovakuumskih (EVV) posod. Pri zelo
nizkih tlakih v UVV- in EVV-podro~ju je vodik prevladujo~ rezidualni plin v vakuumskih sistemih, narejenih iz nerjavnega
jekla. Tako je zmanj{anje razplinjevanja vodika najpomembnej{e za doseganje nizkih tlakov v na{tetih podro~jih. V tem ~lanku
je predmet raziskav vakuumska posoda z debelino stene 2,62 mm, ki je narejena iz nerjavnega jekla AISI type 304 L. Postopki
pregrevanja vakuumske posode so se izvajali pri 250 °C 380 h in pri 350 °C 140 h. Po pregrevanju pri 250 °C 380 h (kar ustreza
brezdimenzijskemu ~asu Fo = 3,09) je bila pri sobni temperaturi dose`ena gostota pretoka razplinjevanja q = 2,86 × 10–13
mbar L s–1 cm–2. Z nadaljnjim pregrevanjem pri 350 °C 140 h (Fo = 8,66, kar rezultira v celotni brezdimenzijski ~as Fo =
11,75) se je gostota pretoka razplinjevanja pri sobni temperaturi zmanj{ala na q = 5,7 × 10–14 mbar L s–1 cm–2.
Klju~ne besede: razplinjevanje vodika, nerjavno jeklo, brezdimenzijski ~as, metoda nara{~anja tlaka, preto~na metoda,
kalibracijski koeficient za plinski pretok
1 INTRODUCTION
The development of the science and technology of
ultrahigh vacuum (UHV) and extremely high vacuum
(XHV) devices has been strongly coupled to the
development of increasingly larger and more sophisti-
cated devices for physics research, such as particle
accelerators, magnetic fusion devices, gravity wave
observatories and many surface analysis techniques.
Scientific advancements in the understanding of the
outgassing limits in UHV/XHV conditions are associated
with these technological developments.
Outgassing refers to the spontaneous liberation of
gases from the walls of a vacuum chamber or other
components placed inside the vacuum chamber. This gas
release results from two processes:1
• Gas diffusion from the interior of vacuum chamber
walls to their inner surface. This is followed by gas
desorption into the chamber volume that contributes
to the vacuum system outgassing.
• Release of gases or vapours previously adsorbed onto
the inner surface of the vacuum chamber walls. These
gases may have adsorbed onto the chamber inner
surface while it was exposed to the environment and
then slowly released as the pump removed the gas
from the vacuum chamber.
The performance of a vacuum system is limited
because it is impossible to eliminate all of the gas
sources. Outgassing occurs even in the best-designed
vacuum systems for UHV/XHV. The ultimate pressure in
the system is related to the magnitude of the gas load;
therefore, the measures taken to reduce outgassing are
critical for the production of UHV/XHV. These effects
extend the time required to reach UHV/XHV.
Due to the vaporisation or sublimation of atoms and
molecules from some materials with a vapour pressure
higher than or comparable to the residual pressure within
a vacuum chamber, these materials will increase out-
gassing. Thus, not all types of materials can be used to
build vacuum systems, and sometimes it is difficult to
select appropriate materials that will fulfil the require-
ments of different processes. Among technical alloys,
austenitic stainless steel (SS) offers several advantages
compared to other alloys, and it is the most widely used
Materiali in tehnologije / Materials and technology 46 (2012) 2, 161–167 161
UDK 533.5:669.14.018.8:669.788 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 46(2)161(2012)
construction material for building vacuum chambers,
instruments and components. SS is the most important
construction material for UHV/XHV systems because of
its good vacuum and mechanical properties.2 It is non-
magnetic, corrosion resistant and chemically inert. At
room temperature (RT), SS exhibits negligible vapour
pressure and negligible permeation of atmospheric
gasses.3 SS is relatively cheap, and it can be machined
and welded by standard procedures.
In the high-vacuum range, the predominant gas is
usually water vapour resulting from the exposure of the
system to humid air.4 The outgassing of hydrogen from
SS is the main gas load in UHV/XHV systems,5 parti-
cularly in large systems (i.e., accelerators and storage
rings). Thus, hydrogen outgassing is the most significant
limiting factor in reaching outgassing rates below 10–12
mbar L s–1 cm–2 in SS vacuum systems.6 By reducing the
hydrogen content in the bulk, the hydrogen outgassing
rate can be reduced.
Many techniques have been reported that mitigate
outgassing in vacuum systems, but reducing the outgass-
ing rates of SS remains a challenge. These outgassing
techniques include the following:
a) Surface treatments to reduce surface roughness, such
as electro-polishing and surface machining under
special conditions 3
b) Surface treatments to create oxide films to act as a
barrier to the diffusion of hydrogen from the bulk3
c) High-temperature bake-out (vacuum firing) to reduce
the amount of dissolved H (for SS as high as 1 000
°C)4,7
d) Baking the vacuum system to remove water vapour,
which has to be performed at 150–450 °C4
e) Degreasing and chemical cleaning4
f) Deposition of thin films to serve as a barrier layer on
inner surfaces8
g) Choosing metals with a low solubility for hydrogen
(e.g., copper)4
h) Reduction of hydrogen surface mobility by introduc-
ing surface trapping to reduce recombination8
i) Thin-film getter coatings to serve as a pump for
hydrogen diffusing from the bulk9
The typical outgassing rate of vacuum components
made from SS is on the order of 10–11 mbar L s–1 cm–2
without additional processing. An outgassing rate of
10–13 mbar L s–1 cm–2 is "routinely" achievable with a
sufficient bake-out treatment. Outgassing rates lower
than 10–14 mbar L s–1 cm–2 are achievable only with
exceptional care, particularly in thin-wall vessels.8 Until
now, several reports have been published describing the
reduction of SS outgassing rates by conventional
bake-out or vacuum firing, but the resulting outgassing
rates are not consistent: similar bake-out temperatures
and times gave significantly different numerical values
for the outgassing rates.
In 1967, Calder and Levin10 published a paper on the
reduction of outgassing of hydrogen from metals. They
proposed the diffusion-limited model (DLM), which
means that the atomic hydrogen concentration at the
surface is zero and the hydrogen atoms desorb and
recombine as molecules as soon as hydrogen reaches the
surface. However, their assumption was not correct for
lower outgassing rates, whereas their assumption was
reasonable for higher outgassing rates. The experimental
values obtained by Calder and Levin were two orders of
magnitude larger than expected according to the DLM.2
Therefore, their assumption that the surface density of
hydrogen when solving the diffusion equations is always
zero is questionable. To overcome this problem, Moore11
calculated the outgassing rate using a recombination-
limited model (RLM). Moore assumed that the surface
density can never be zero and that the boundary condi-
tions must be included when solving the diffusion
equation.
Nemani~, [etina, Bogataj and Zajec12–14 used very
thin-wall vessels instead of using thicker materials to
reduce the outgassing rate. They also introduced a
Fourier number (or dimensionless time) Fo to describe
the heat treatment intensity. They reported extremely low
outgassing rates of approximately 10–15 mbar L s–1 cm–2,
which were achieved using thinner materials.12 There-
fore, the use of thin materials is more efficient and
economically suitable because shorter baking times can
be used. However, the insufficient mechanical strength of
very thin walls may make them unsuitable for the
practical realisation of a vacuum chamber.
1.1 Measuring of outassing flux and outgassing rate
The outgassing flux of a SS sample can be measured
by two general methods4. In the first method, the
pressure increase is measured as a function of time in a
closed vacuum chamber after its evacuation to a low
pressure. This method is called the rate-of-rise (RoR)
method or the gas-accumulation method. The outgassing
flux Q is obtained as
Q V
P
t
=
d
d
(1)
where V is the chamber volume and dP/dt is the
measured rate of the pressure change in a closed
chamber at a constant temperature.
In the second method, the vacuum test chamber is
pumped through an orifice of known conductance, and
the pressure drop across the orifice is measured. This
method is known as the throughput method or the orifice
method. The outgassing flux Q is expressed as
Q C P= Δ (2)
where C is the conductance and P is the pressure diffe-
rence.
In both methods, the outgassing rate q of the sample
is obtained by dividing the measured outgassing flux Q
by the surface area A of the sample: q = Q/A. The
measured outgassing rate is the net rate of the difference
S. AVDIAJ, B. ERJAVEC: OUTGASSING OF HYDROGEN FROM A STAINLESS STEEL VACUUM CHAMBER
162 Materiali in tehnologije / Materials and technology 46 (2012) 2, 161–167
between the intrinsic outgassing rate of the surface and
the readsorption rate.15
A common problem of both methods is that the test
sample is placed in a vacuum measurement chamber
whose walls also outgas during the measurement. Thus,
the background outgassing flux of the empty chamber
must be measured and subtracted from the measured
outgassing flux of the sample in the chamber.
The presence of a hot cathode residual gas analyser
(RGA) or ion gauge in the test chamber can cause
additional outgassing, and the residual gas can react with
the hot filament. Because they exhibit a certain pumping
speed, RoR measurements can only be obtained with an
inert vacuum gauge, such as a spinning rotor gauge
(SRG).1
2 EXPERIMENTAL SETUP AND SAMPLE
DESCRIPTION
We measured the time dependence of the outgassing
rate of SS during bake-outs at 250 °C and 350 °C. Our
main goal was to determine the time needed to reach the
outgassing rate below 10–13 mbar L s–1 cm–2 at these
bake-out temperatures.
We made two simple cylindrical chambers to serve as
the outgassing samples (Figure 1). Each chamber was
assembled from a CF 16 flange and had a short
connection tube with an inner diameter of 16 mm and a
wall thickness of 1.5 mm. The circular end plates had a
thickness of 2 mm, and the cylindrical body had an inner
diameter di = 108.7 mm and a wall thickness of 2.62
mm. The only difference between the two chambers was
the cylindrical body length. The length of the body of
chamber V1 was l1 = 10 mm, and the length of the body
of chamber V2 was l2 = 210 mm.
The chambers were connected to a vacuum
measurement system, as shown in Figure 2. The system
was pumped by a turbomolecular pump and was
equipped with a quadrupole mass spectrometer (QMS), a
Bayard-Alpert gauge (BAG) and an SRG. The SRG was
connected to a small chamber V0 that was made of a
standard CF16 4-way cross and a CF16 T-piece. The
chamber V0 can be isolated from the pump by a valve
(E0). The chambers V1 and V2 were symmetrically
connected to V0 through the bake-proof CF16 all-metal
valves E1 and E2, respectively. The calibration gas was
introduced into V0 through valve E3.
For the modelling of the outgassing behaviour of SS
by DLM and RLM, the thickness of the sample is an
important parameter. The same reduction in the
outgassing rate of a plane sheet takes 4 times longer time
if the sample is 2 times thicker. The entire measurement
sample chamber was constructed from AISI type 304 L
SS, but we were unable to make the whole chamber from
the same sheet of SS. Additionally, each part had a
different thickness. However, the measurement pro-
cedure was designed so that we could easily calculate the
outgassing rate of the cylindrical body (which has a
well-defined thickness) from the separate measurements
of the outgassing flux of each chamber.
The chambers V0, V1 and V2 were placed in an oven
where they could be uniformly heated up to 350 °C.
Additional vacuum measurement instruments (BAG and
QMS) were mounted outside the oven due to the
S. AVDIAJ, B. ERJAVEC: OUTGASSING OF HYDROGEN FROM A STAINLESS STEEL VACUUM CHAMBER
Materiali in tehnologije / Materials and technology 46 (2012) 2, 161–167 163
Figure 2: A schematic of the outgassing measurement system: QMS –
quadrupole mass spectrometer; SRG – spinning rotor gauge; BAG –
Bayard-Alpert gauge; E0, E1, E2 and E3 – bake-proof vacuum valves;
V0, V1 and V2 – measuring volumes
Slika 2: Shema vakuumskega sistema za meritve razplinjevanja: QMS
– kvadropolni masni spektrometer; SRG – viskoznostni merilnik z
lebde~o kroglico; BAG- Bayard-Alpertova trioda; E0, E1, E2 in E3 –
pregrevni vakuumski ventili; V0, V1 in V2 – merilni volumni
Figure 1: A photo of two SS cylindrical vacuum chambers used for
the outgassing rate study
Slika 1: Fotografija dveh cilindri~nih vakuumskih posod iz nerjavnega
jekla, namenjenih za {tudij razplinjevanja
limitation of their operational temperature. During the
bake-out, the SRG suspension head was removed, and
only a thimble with a rotor was baked with the other
parts.
The outgassing flux measurements at ambient
temperature were performed using the RoR method. The
outgassing flux measurements while baking the vacuum
system (at temperatures of either 250 °C or 350 °C) were
performed using the throughput method, which is
described in Section 2.2.
2.1 Measuring of outassing flux using the RoR method
By closing valves E0, E2 and E3 and opening valve
E1, we measured the pressure increase in chambers V0 +
V1 with the SRG and used this reading to calculate the
outgassing flux Q1 = QV0 + QV1. With the same SRG, we
also measured the pressure increase in chambers V0 +
V2 when valve E1 was closed and E2 was opened. This
measurement gave Q2 = QV0 + QV2. In both cases, QV0
was the same because of the symmetrical configuration
of the connection with V1 and V2. By subtracting the
two measured outgassing fluxes, the outgassing flux of
V0 can be eliminated: Q2 – Q1 = QV2 – QV1. The
difference is equal to the outgassing flux from the major
part of the cylindrical body of chamber V2 having a
surface area A = di × (l2 – l1).
To calculate the outgassing flux by the RoR method
following Equation 1, we need to know the volume of
the measurement system. The volumes V1 and V2 can
be calculated from geometrical measurement because
they are composed of simple cylindrical parts. The
volume V0 has a more complicated geometrical shape
because it includes the inner volume of the CF16 valves.
Its volume can be determined by the static gas-expansion
method. The calculated surface areas of V1 and V2 are
accurate. However, we can only estimate the surface area
of V0. The volumes and the surface areas for the 3 parts
of the vacuum system are given in Table 1.
Table 1: The geometrical surface areas and volumes of particular parts
of the vacuum measurement system, including their ratios
Tabela 1: Geometrijska povr{ina A in volumen V posameznih delov
vakuumskega merilnega sistema, vklju~ujo~ njihovo razmerje
Geometrical surface area
A (cm2)
Volume
V (L)
Ratio (V/A)
(L cm–2)
V0 475 (rough estimate) 0.1206 2.54 × 10–4
V1 247 0.1046 4.23 × 10–4
V2 931 1.962 2.11 × 10–3
Vtot 1653 2.1872 1.32 × 10–3
2.2 Measuring of outassing flux using throughput me-
thod
To measure the outgassing flux of the test chambers
during baking, the use of the SRG is not possible
because the operating temperature for the measuring
head of the SRG only ranges from 10 °C to 50 °C.16 As a
consequence, the RoR method cannot be used. There-
fore, an adapted throughput measurement method that
differs from the orifice method was used.
The steady-state outgassing flux from chamber V1
can be calculated by measuring the pressure difference
P1, as measured by the BAG or QMS while opening
and closing valve E1. By knowing the effective pumping
speed S at the point where the BAG or QMS are
connected to the vacuum system, we can calculate the
corresponding outgassing flux Q1 = S × P1. Similarly,
by opening and closing E2, we can determine P2, and
the outgassing flux Q2 = S × P2. Also, the difference
between the two fluxes is the outgassing flux from the
major part of the cylindrical body of chamber V2, which
has a surface area A = di × (l2 – l1), as in the case of
RoR measurements.
We were primarily interested in the outgassing of
hydrogen from SS, so we used a QMS tuned to the mass
number 2 (H2 peak) for the P measurements. The basic
output signal of the QMS is an ion current of a particular
gas. The ion current Ii+ as a function of the partial gas
pressure Pi is given by Ii+ = gi × Pi, where gi is the
sensitivity coefficient, which depends on the gas type,
the QMS geometry and the settings of various
operational parameters. For accurate measurements, gi
must be determined experimentally (calibrated in situ)
for each instrument and each gas. Therefore, to measure
the outgassing flux of hydrogen, the sensitivity coeffi-
cient of the QMS gH2 and the effective pumping speed
SH2 had to be determined. For hydrogen, the outgassing
flux Q equals
Q
S I
g
=
+
H H
H
2 2
2
Δ
(3)
To more easily determine the outgassing rate q =
Q/A, we combined the effective pumping speed and the
sensitivity coefficient into a single calibration coefficient
KH2:
K S gH H H2 2 2= / (4)
The calibration coefficient KH2 can be determined
experimentally in a straightforward manner and is
described in the following section.
2.3 Determination of calibration coefficient
We can introduce a known flow of a calibration gas
through valve E3, which is an adjustable leak valve. By
closing valves E0, E1 and E2, the flow rate of the
calibration gas Qcal into the volume V0 can be measured
by the RoR method using the same SRG as for the
outgassing flux measurement. After the flow of cali-
bration gas is determined, we can measure the difference
in the hydrogen ion current IH2+ by opening valve E0.
Because the calibration gas flow was introduced to the
QMS from the same direction as the outgassing flux
from the measuring chambers, the calibration coeffi-
cients for the calibration gas flow and the outgassing flux
S. AVDIAJ, B. ERJAVEC: OUTGASSING OF HYDROGEN FROM A STAINLESS STEEL VACUUM CHAMBER
164 Materiali in tehnologije / Materials and technology 46 (2012) 2, 161–167
have similar values. The calibration coefficient can be
calculated by rearranging Equation 3:
K
Q
IH
cal
H
2
2
= +Δ
(5)
The measurements of IH2+ have been performed for
a wide range of calibration gas flows Qcal. Figure 3
shows that when the Qcal (measured by the SRG)
increases, the IH2+ (measured by the QMS) increases
proportionally.
From the data in Figure 3 and with Equation 5, we
have calculated the calibration coefficient. The cal-
culated values are given in Table 2. The mean value of
the calibration coefficient K of the QMS for hydrogen
was 1.12 × 105 mbar L A–1 s–1, and the relative standard
deviation was (K)/K = 3.61 %.
Table 2: The calibration coefficients of QMS for hydrogen while
measuring outgassing flux by the throughput method
Tabela 2: Kalibracijski koeficient QMS pri merjenju pretoka raz-
plinjevanja vodika s preto~no metodo
I+ (A) KH2 (mbar L A–1 s–1)
1.21 × 10–10 116998
1.10 × 10–10 118505
3.63 × 10–11 114295
2.37 × 10–11 115234
1.47 × 10–11 113737
1.42 × 10–11 110694
1.41 × 10–11 110739
5.54 × 10–12 106750
1.83 × 10–12 107483
1.24 × 10–12 107763
1.21 × 10–12 108467
1.20 × 10–12 108439
The final equation to calculate the outgassing rate is
q
K
A
I I= −( )Δ ΔV2 V1 (6)
where IV1 and IV2 are the changes in the hydrogen
ion current when closing valves E1 and E2, respectively.
3 RESULTS AND DISCUSSION
3.1 Time dependence of H2 outgassing rate during
bake-out at 250 °C
The measuring chambers were first baked at 250 °C
for 380 h. The bake-out temperature was increased to
350 °C, and baking continued for an additional 140 h.
The chambers were baked by keeping the outer side of
the chamber wall in air while the inner side was under
vacuum.
During the bake-out at 250 °C, we performed several
measurements of the outgassing rate at the bake-out
temperature using the throughput method. The heat
treatment intensity expressed as a Fourier number, the
baking time of the system and the outgassing rates
during the heat treatment are given in Table 3.
Table 3: The measured hydrogen outgassing rate of the SS vacuum
chamber during the bake-out. All of the outgassing rates were
measured at a temperature of 250 °C.
Tabela 3: Izmerjena gostota pretoka razplinjevanja vodika vakuumske
komore iz nerjavnega jekla med pregrevanjem pri 250 °C
Fo t (h) at 250 C q (mbar l s–1 cm–2)
0.20 24 2.6 × 10–8
0.77 95 1.10 × 10–8
0.77 95.2 1.12 × 10–8
0.78 96 1.19 × 10–8
0.88 108 8.20 × 10–9
0.89 110 8.40 × 10–9
0.91 112 8.80 × 10–9
1.54 190 4.20 × 10–9
1.59 196 3.90 × 10–9
1.71 211 2.90 × 10–9
1.72 211.5 2.95 × 10–9
1.72 212 3.00 × 10–9
1.84 227 2.50 × 10–9
1.86 228.5 2.40 × 10–9
1.87 230 2.30 × 10–9
2.06 254 2.00 × 10–9
2.07 255 2.05 × 10–9
2.08 256 2.08 × 10–9
2.26 278 1.80 × 10–9
2.27 279 1.85 × 10–9
2.84 350 1.00 × 10–9
2.85 351 1.05 × 10–9
3.06 376 6.90 × 10–10
3.07 378 7.30 × 10–10
The Fourier number (or dimensionless time) Fo =
4Dtd–2 is a characteristic quantity for describing and
modelling diffusion. It was used to compare the
outgassing rates for different experiments.17 To deter-
mine the dimensionless time Fo, the following must be
known: the diffusion constant D for hydrogen at tem-
perature T, the processing time t and the wall thickness d.
S. AVDIAJ, B. ERJAVEC: OUTGASSING OF HYDROGEN FROM A STAINLESS STEEL VACUUM CHAMBER
Materiali in tehnologije / Materials and technology 46 (2012) 2, 161–167 165
Figure 3: The difference in the hydrogen ion current, measured by
QMS (using the throughput method), as a function of the hydrogen
gas flow, measured by SRG (using the RoR method)
Slika 3: Razlika v ionskem toku vodika (ki je bil izmerjen s preto~no
metodo in uporabo QMS) v odvisnosti od pretoka vodika (izmerjenega
z metodo hitrosti nara{~anja tlaka in uporabo SRG)
The parameters t and d are measured directly. The
temperature dependence of D is given by
D D
E
kT
a=
−⎛
⎝
⎜ ⎞
⎠
⎟
0 exp (7)
Typical values for the diffusion pre-exponential
factor (D0 = 0.012 cm2 s–1) and the activation energy (Ea
= 0.57 eV)18 were used to calculate the Fo values.
The bake-out at 250 °C was interrupted a few times
to cool the chambers and to measure the outgassing rates
at RT. The time course was as follows: bake-out at 250
°C for 24 h, interrupt baking for the RT measurement,
bake-out at 250 °C for 73 h, interrupt baking for the RT
measurement and bake-out at 250 °C for 284 h followed
by an RT measurement. The temperature was then raised
to 350 °C, and baking was performed for additional 140
h. The system was then cooled again, and the last RT
measurement was performed. The results of the hydro-
gen outgassing rate measurements at RT, which were
performed using the RoR method, are summarised in
Table 4.
Table 4: The hydrogen RT outgassing rates of SS after the indicated
bake-out treatment
Tabela 4: Gostota pretoka razplinjevanja vodika iz nerjavnega jekla
pri sobni temperaturi po navedenem postopku pregrevanja
T (°C) Incrementaltime t (h) Fo (t) Fo
q
(mbar L s–1 cm–2)
250 24 0.19 0.19 2.10 × 10–12
250 73 0.59 0.78
250 284 2.31 3.09
350 140 8.66 11.75
A comparison of our results of the RT outgassing rate
with the results obtained by Park et al.19 is shown in
Figure 4. The Fourier numbers for the data from Park et
al. were recalculated using our activation energy for the
diffusion Ea = 0.57 eV (instead of Ea = 14.5 kcal mol–1 
0.622 eV, which was used by Park et al.). Our results
agree and are within a factor 3 to 4 of the results of Park
et al.
A graphical representation of the results (taken from
Table 3) of the outgassing rates measured at 250 °C as a
function of dimensionless time Fo is shown in Figure 5.
In Figure 6, the same measured outgassing rates as a
function of the bake-out time are presented and
compared with the calculations based on the DLM.1,6
The outgassing rate predicted by diffusion theory for t =
380 h is q = 5 × 10–11 mbar L s–1 cm–2, which is 15 times
lower than the measured value q = 7 × 10–10 mbar L s–1 cm–2.
In the early stage of bake-out, until approximately 100 h
at 250 °C, the predicted curve according to the DLM and
the measured data agree. However, discrepancies start to
emerge after 100 h. For Fo > 0.8, the atomic hydrogen
recombination on the SS surface is assumed to become a
rate-limiting step. When a hydrogen atom has moved
from the bulk to the surface/vacuum interface, it needs to
find another hydrogen atom on the surface to recombine.
S. AVDIAJ, B. ERJAVEC: OUTGASSING OF HYDROGEN FROM A STAINLESS STEEL VACUUM CHAMBER
166 Materiali in tehnologije / Materials and technology 46 (2012) 2, 161–167
Figure 6: A comparison of the DLM-predicted hydrogen outgassing
rate of SS and the measured data during the heat treatment at 250 °C
as a function of the bake-out time
Slika 6: Primerjava gostote pretoka razplinjevanja vodika iz
nerjavnega jekla, predvidenega na osnovi modela DLM, z merilnimi
rezultati v odvisnosti od ~asa pregrevanja pri 250 °C
Figure 4: A comparison of the hydrogen outgassing rates between our
measurements and the measurements of Park et al.17, which were
performed after the system was cooled to RT
Slika 4: Primerjava gostote pretoka razplinjevanja vodika iz
nerjavnega jekla med na{imi meritvami in meritvami Parka et al.17, ki
so bile izvedene po ohladitvi sistema na sobno temperaturo
Figure 5: The hydrogen outgassing rate of SS as a function of the
dimensionless time Fo during the heat treatment at 250 °C
Slika 5: Gostota pretoka razplinjevanja vodika iz nerjavnega jekla v
odvisnosti od brezdimenzijskega ~asa Fo med postopkom pregrevanja
pri 250 °C
As the number of freely diffusing hydrogen surface
atoms is reduced, recombination becomes less likely.
Thus, when the diffusion of hydrogen atoms to the
surface become small, the recombination of the hydro-
gen atoms might be the rate-limiting step, i.e., outgassing
from the surface is described by the rate at which the
hydrogen atoms can be recombined, not the bulk diffu-
sion rate.
4 CONCLUSIONS
The determination of the calibration coefficient K for
the BAG and QMS allowed us to measure the outgassing
rate during the heat treatment without stopping the
heating process. The DLM governs the initial removal of
hydrogen from SS, and it is assumes that hydrogen
surface recombination plays an important role in the
outgassing rate at lower hydrogen concentrations. To
compare the different outgassing treatments, the Fo
number appears to be a good choice because it can be
accurately calculated for any processing time and
temperature. The Fo number is a parameter that shows
the level of heat treatment of SS, and it is related to the
diffusion of hydrogen atoms in SS.
An outgassing rate q = 2.86 × 10–13 mbar L s–1 cm–2 at
RT was achieved for AISI type 304 L SS after baking the
system at 250 °C for 380 h (conversion to dimensionless
time gives Fo = 3.09). This outgassing rate was further
reduced to q = 5.7 × 10–14 mbar L s–1 cm–2 by baking for
another 140 h at T = 350 °C (Fo = 8.66, total dimen-
sionless time Fo = 11.75).
Acknowledgement
S. Avdiaj acknowledges partial financial support by
the European Union, European Social Fund. The study
was implemented in the framework of the Operational
Programme for Human Resources Development for the
period 2007–2013.
5 REFERENCES
1 K. Jousten, Thermal outgassing, Proceedings of the CERN Acce-
lerator School, Snekersten, Denmark, CERN report, edited by S.
Turner, 1999, 111–125
2 Y. Ishikava, V. Nemanic, Vacuum, 69 (2003), 501–512
3 M. Leich, Hydrogen outgassing of stainless steel, our present knowl-
edge, Proceedings of the 1st Vacuum Symposium, UK, 2010
(http://www.ss.dsl.pipex.com/rgaug/pdfs/ vs1/Leisch.pdf)
4 J. Lafferty, Foundations of Vacuum Science and Technology, John
Wiley & Sons, New York 1998, 625–652
5 P. A. Redhead, Extreme high vacuum, Proceedings of the CERN
Acceleration School, Snekersten, Denmark, CERN report, edited by
S. Turner, 1999, 213–227
6 M. Bernardini et al., J. Vac. Sci. Technol., A 16 (1998), 188–193
7 B. Zajec, V. Nemani~, Mater. Tehnol., 35 (2001) 5, 259–264
8 R. Reid, Outgassing Surface Conditioning, 2009, (www.stfc.ac.uk/
resources/pdf/ronreid25209.pdf)
9 C. Benvenuti, P. Chiggiato, F. Cicoira, V. Ruzinov, Vacuum, 50
(1998), 57–63
10 R. Calder, G. Lewin, Br. J. Appl. Phys., 18 (1967), 1459
11 B.C. Moore, J. Vac. Sci. Technol., A 13 (1995), 545–548
12 V. Nemanic, J. Setina, J. Vac. Sci. Technol., A 17 (1999), 1040–1046
13 V. Nemanic, T. Bogataj, Vacuum, 50 (1998), 431–437
14 V. Nemanic, B. Zajec, J. Setina, J. Vac. Sci. Technol., A 19 (2001),
215–222
15 P. A. Redhead, Hydrogen in vacuum systems; an Overview, Proceed-
ings of the First International Workshop on Hydrogen in Materials
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2003, 243–254, (http://adsabs.harvard.edu/abs/2003AIPC..671..
243R)
16 www.mksinst.com
17 V. Nemani~, T. Bogataj, Kovine zlitine tehnologije, 32 (1998) 3–4,
249–253
18 D. M. Grant et al., J. Nucl. Mater., 149 (1987), 180–191
19 Park et al., J. Vac. Sci. Technol., A 26 (2008), 1166–1171
S. AVDIAJ, B. ERJAVEC: OUTGASSING OF HYDROGEN FROM A STAINLESS STEEL VACUUM CHAMBER
Materiali in tehnologije / Materials and technology 46 (2012) 2, 161–167 167

M. BALCAR et al.: THE QUALITY OF SUPER-CLEAN STEELS PRODUCED AT @ÏAS, inc.
THE QUALITY OF SUPER-CLEAN STEELS PRODUCED
AT @ÏAS, inc.
KAKOVOST SUPER^ISTIH JEKEL, IZDELANIH V PODJETJU
@ÏAS, inc.
Martin Balcar1, Ludvík Martínek1, Pavel Fila1, Jaroslav Novák1, Jiøí Ba`an2,
Ladislav Socha2, Danijela Anica Skobir Balanti~3, Matja` Godec3
1@ÏAS, a. s., Strojírenská 6, 591 71 @ïár nad Sázavou, Czech Republic
2V[B – Technical University of Ostrava, 17. listopadu 15/2172, 708 33 Ostrava-Poruba, Czech Republic
3Institute of Metals and Technology, Lepi pot 11, 1000 Ljubjana, Slovenia
martin.balcar@zdas.cz
Prejem rokopisa – received: 2011-05-30; sprejem za objavo – accepted for publication: 2011-08-24
The production of Super-Clean Steels for the rotor forgings of compressors and generators for gas-turbine units started at ZDAS
with the use of secondary metallurgy processes, a ladle furnace and vacuum degassing. The development and optimization of
Super-Clean Steel production technology enables effective molten metal manufacture, conforming to the requirements for
chemical composition and micro-cleanness. According to the results of the current production, the effective production of rotor
forgings requires new technological steps in ingot casting.
Keywords: super-clean steel, steelmaking, secondary metallurgy, ingot casting
Proizvodnja super~istih jekel za odkovke rotorjev kompresorjev in generatorjev za plinske turbine se je za~ela z za~etkom
uporabe procesov sekundarne metalurgije, s ponov~no pe~jo in z vakuumsko degazacijo. Razvoj in optimizacija tehnologije
super~istih jekel omogo~ata u~inkovito izdelavo taline z upo{tevanjem kemi~ne sestave in mikro~istosti. Glede na rezultate
sedanje proizvodnje so za u~inkovito izdelavo rotorjev potrebne tehnolo{ke izbolj{ave litja ingotov.
Klju~ne besede: super~isto jeklo, izdelava jekla, sekundarna metalurgija, litje ingotov
1 INTRODUCTION
The production of rotors at ZDAS consists of
medium-weight forgings for equipment to generate
electric power, gas turbines of the type GT – 009 with a
maximum output of 11.7 MW and a gas temperature at
the outlet up to 580 °C.1
In the frame of the production of a trial series of
forgings for compressor and generator rotors in ZDAS,
samples of steel were taken during the forging of ingots
8K10.0 and 8K13.0 from the steel grade
26NiCrMoV115.
The analyses of the chemical composition and the
evaluations of the results from the viewpoint of the
achieved parameters of chemical cleanliness, as well as
from the viewpoint of the influence of casting and
solidification on the differences between the chemical
composition of the melt and the forging, make it possible
to interpret the stability of the production process. The
analyses of forging defects provided sufficient infor-
mation about the possible causes of defects.
2 CHEMICAL COMPOSITION OF A FORGING
MADE OF SUPER-CLEAN STEELS
Table 1 summarises the requirements for the
chemical composition of super-clean steel (SCS) for
Materiali in tehnologije / Materials and technology 46 (2012) 2, 169–175 169
UDK 669.18(437.3) ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 46(2)169(2012)
Table 1: Chemical composition of steel 26NiCrMoV115 in mass fractions, w/%
Tabela 1: Kemi~na sestava jekla 26NiCrMoV115 v masnih dele`ih, w/%
A C Mn Si P S Cr Ni Mo V Al Cu As Sn Sb
(w/%) (ìg/g)
min. 0.26 max. max. max. max. 1.40 2.80 0.30 max. max. max. max. max. max.
max. 0.32 0.30 0.07 0.007 0.005 1.70 3.00 0.45 0.15 0.010 0.12 100 100 50
Table 2: Chemical composition of steel 26NiCrMoV145 in mass fractions, w/%
Tabela 2: Kemi~na sestava jekla 26NiCrMoV145 v masnih dele`ih, w/%
B C Mn Si P S Cr Ni Mo V Al Cu As Sn Sb
X factor
(w/%) (ìg/g)
min. 0.26 max. max. max. max. 1.60 3.50 0.30 max. max. max. max. max. max. max.
max 0.32 0.04 0.04 0.004 0.004 1.90 3.80 0.45 0.15 0.015 0.12 80 50 20 7.0
compressor and generator rotors and Table 2 for the
discs of turbine and generator wheels.
The average contents of the alloying and tramp
elements of 87 heats of steel grade 26NiCrMoV115 (A)
and 19 heats of steel grade 26NiCrMoV145 (B) are
given in Tables 3 and 4.
On the basis of the ordinary production of forgings a
complete chemical composition was determined for 44
samples of steel from the forgings, i.e., for 44 ingots
from various heats of the steel grade 26NiCrMoV115.
Figures 1 to 4 show the distribution of the content of the
elements P, S, O and N.
For the monitored 44 heats the average content of
phosphorus is 37.1 μg/g and the standard deviation is
7.65 μg/g. The contents vary in the range from 20 μg/g to
60 μg/g.
The average content of sulphur was of 29.3 μg/g,
with the variation in the range from 10 μg/g to 50 μg/g
and a standard deviation of 12.08 μg/g.
The distribution of oxygen content is shown in
Figure 3. The average content of oxygen was 20.6 μg/g
M. BALCAR et al.: THE QUALITY OF SUPER-CLEAN STEELS PRODUCED AT @ÏAS, inc.
170 Materiali in tehnologije / Materials and technology 46 (2012) 2, 169–175
Table 3: Average content of elements in the heats of steel grade 26NiCrMoV115 in mass fractions, w/%
Tabela 3: Povpre~na vsebnost elementov v talinah jekla 26NiCrMoV115 v masnih dele`ih, w/%
A C Mn Si P S Cr Ni Mo V Al Cu As Sn Sb
(w/%) (μg/g)
AVG 0.295 0.210 0.023 0.0043 0.0028 1.588 2.913 0.390 0.106 0.0063 0.080 49.4 58.5 29.3
s 0.011 0.031 0.017 0.0008 0.0014 0.035 0.033 0.013 0.009 0.0020 0.024 7.7 15.1 2.5
Table 4: Average content of elements in the heats of steel grade 26NiCrMoV145 in mass fractions, w/%
Tabela 4: Povpre~na vsebnost elementov v talinah jekla 26NiCrMoV145 v masnih dele`ih, w/%
B C Mn Si P S Cr Ni Mo V Al Cu As* Sn* Sb*
X factor
(w/%) (μg/g)
AVG 0.293 0.032 0.012 0.0032 0.0044 1.809 3.714 0.395 0.108 0.008 0.014 31.1 32.1 <20 5.81
s 0.012 0.010 0.004 0.0006 0.0039 0.046 0.040 0.0011 0.008 0.003 0.008 18.6 17.6 – 1.10
* Concentration of elements below the detection limits
Table 5: Correlation coefficients of elements (sample of the melt / forging)
Tabela 5: Korelacija kemi~ne sestave talin in odkovkov
Elements Ni Cu Mn Si Cr V S Ca Al C Mo P N O
X (w/%) forging 0,99 0,99 0,99 0,98 0,97 0,92 0,89 0,85 0,78 0,78 0,75 0,60 0,46 0,05
Figure 3: Oxygen content – forging
Slika 3: Vsebnost kisika v odkovkih
Figure 2: Sulphur content – forging
Slika 2: Vsebnost `vepla v odkovkih
Figure 1: Phosphorus content – forging
Slika 1: Vsebnost fosforja v odkovkih
and the standard deviation was 4.56 μg/g. The oxygen
content in the forgings varied in the range from 12 μg/g
to 32 μg/g. The average content of nitrogen in Figure 4
was 64.0 μg/g and the standard deviation was 11.36 μg/g.
The nitrogen content in the forgings varied in the range
from 34 μg/g to 84 μg/g.
3 AGREEMENT OF THE CHEMICAL ANALYSIS
OF THE MELT WITH THE ANALYSIS OF THE
FORGING
The results of the chemical composition of the
samples of steel forgings were compared with the results
of the chemical analysis of the melt to verify the
agreement of both chemical analyses. The correlation
coefficients of the individual elements in descending
agreement are shown in Table 5.
The results in Table 5 suggest that the agreement of
the chemical composition of the forgings and the melts
depends on the place in the sample ingot where the
measurement was made. If we consider the position of
the analysed sample is below the ingot head, which is the
place of its biggest cross-section, and simultaneously the
latest solidification part of the ingot body, it may be
expected that due to segregations, the concentrations of
some elements may be influenced during the ingot’s
solidification.
This assumption is confirmed by the order of the
correlation coefficients of chromium, vanadium and
molybdenum, i.e., elements that form carbides. Phospho-
rus and sulphur show a high degree of segregation and
the low correlation coefficient suggests the segregation
of nitrogen and aluminium, which have a great mutual
affinity.
The lowest correlation coefficients according to
Table 5 were calculated for the gases, oxygen and nitro-
gen, while the correlation coefficient for the oxygen
concentrations is negligible. In Figures 5 to 8 the
dependence of selected elements, i.e., calcium, alumi-
nium, nitrogen and oxygen, is shown.
The disagreement in the oxygen and nitrogen
contents is apparently related to the casting process and
the sampling of metal for the analysis of both elements
M. BALCAR et al.: THE QUALITY OF SUPER-CLEAN STEELS PRODUCED AT @ÏAS, inc.
Materiali in tehnologije / Materials and technology 46 (2012) 2, 169–175 171
Figure 7: Nitrogen content – melt / forging
Slika 7: Vsebnost du{ika – talina/odkovek
Figure 5: Calcium content – melt / forging
Slika 5: Vsebnost kalcija – talina/odkovek
Figure 6: Aluminium content – melt / forging
Slika 6: Vsebnost aluminija – talina/odkovek
Figure 4: Nitrogen content – forging
Slika 4: Vsebnost du{ika v odkovkih
from the melt. The sampling occurs by taking a small
amount of melt from the flow of metal under the slide
gate into the steel ladle, from which the metal is after-
wards poured again into the ingot mould. This process
occurs with considerable contact of the melt with
surrounding atmosphere, which creates good conditions
for the saturation of the degassed melt with oxygen and
nitrogen.
For the oxygen content, we consider the concen-
tration determined by the chemical analysis of the
sample taken from the forging to be realistic one. On the
basis of the results of the analyses it is possible to
discuss the potential control of the steel’s chemical
composition, as well as possibility of verifying the
obtained individual elements concentrations already
during hot-metal production. Namely, the prediction of
oxygen and nitrogen contents in the production of hot
metal appears to be rather problematic with respect to
the final forging contents. This suggest that the existing
methodology for taking samples of melts by pouring for
a determination of the gas contents in steel of the type
26NiCrMoV115 and 26NiCrMoV145 is unsatisfactory.
The solution to this issue may be the realisation of
equipment that can take samples with the elimination of
the earlier mentioned influence of the atmosphere, i.e.,
preferably by sampling directly from the ladle at the end
of the treatment by the VD or VCD process and from the
ingot, either already during pouring or after its com-
pletion.
4 METALLOGRAPHIC CLEANLINESS OF
SUPER-CLEAN STEEL FORGINGS
The metallographic cleanliness of steel in conformity
with the standard DIN 50602 was determined according
the method K4 for 44 heats from identical samples, as
for previous examinations, and for an additional 10 heats
using samples taken in a similar manner. Thus there was
a total of 54 heats.
The distribution of micro-cleanness determined
according the standard DIN 50602 method K4 is shown
in Figure 9 interlaid with the curve of the normal
distribution with the exclusion of the extreme values of
K4 > 20. The average micro-cleanness K4 = 6.3 with a
standard deviation of 5.61 was calculated for 54 heats.
The values of K4 were in the range from 0 to 29.
From the viewpoint of the current requirements for
the cleanliness of steel the values K4 > 10 can be con-
sidered as deteriorated and K4 > 20 as unsatisfactory.
However, the limits stipulated in this manner are relative
and they are based on the assumption that the
deteriorated micro-cleanness will considerably influence
the mechanical properties, particularly the strength
characteristics and the transition temperature or the creep
resistance of the forgings.
In accordance with the defined measures, very good
micro-cleanness was found for 45 heats (83.3 %) out of
the 54 examined heats, while 6 heats (11.1 %) had worse
micro-cleanness, and an unsatisfactory micro-cleanness
was found for 3 heats, i.e., in 5.6 % of production.
In spite of the deteriorated parameters of the
metallographic purity of the steels for some heats, the
forgings passed the required tests of mechanical
properties, even without special measures concerning
their heat treatment. It is therefore possible to consider
the achieved metallographic cleanliness of super-clean
steels is acceptable. However, the objective should be to
achieve the value of K4 < 10.
The measures aimed at ensuring the required
cleanliness may be the optimisation of slag mode or its
possible modification. Due to the occurrence of
exogenous inclusions it is not possible to also exclude
the casting technology, including issues related to the
ceramics used for pouring.2
5 ANALYSIS OF THE DEFECTS IN
SUPER-CLEAN STEEL FORGINGS
Altogether, 122 shafts were produced until 2006, out
of which 18 shafts were classified as unsatisfactory due
to the occurrence of undesirable ultrasonic defects. A
M. BALCAR et al.: THE QUALITY OF SUPER-CLEAN STEELS PRODUCED AT @ÏAS, inc.
172 Materiali in tehnologije / Materials and technology 46 (2012) 2, 169–175
Figure 9: Micro-cleanness DIN50602, method K4
Slika 9: Mikro~istost po DIN50602, metoda K4
Figure 8: Oxygen content – melt / forging
Slika 8: Vsebnost kisika – talina/odkovek
total of 14.8 % of the total number of produced shafts
was rejected. Altogether, 63 pieces of shafts were made
from the ingots 8K10,0 and 59 shafts from the ingots
8K13,0, while 10 pieces of rejected shafts were made
from the ingots 8K10,0 and another 8 pieces of rejected
shafts were made from the ingots 8K13,0 3.
Defective forgings were submitted to a metallo-
graphic investigation and in the following review
documents the results of the analysis of the forging No.
447 660 of the generator rotor shaft are presented.
The shaft with a diameter of 270 mm ingot heel in
Figures 10 and 11 did not pass the ultrasonic test
performed on the roughed piece prior to drilling of
straight-through hole with a diameter of 95 mm. It was
expected that with drilling of the hole the defects will be
removed. After drilling and heat treatment an areal
defect KSR 1 to 4 mm was detected at a depth of 60 mm
to 75 mm in the central part of the piece.
A sample was taken from the forging in the
transversal direction and the exact position of the defect
was localised by repeated ultrasonic testing. A sample
for metallographic analysis was taken from the place of
the defect and after completion of the section at the
location of the defect longitudinally with respect to the
axis of the original forging continuous non-metallic
inclusions was discovered on the full length of the
sample (24 mm) of width of 1 mm. The macro-shape of
the inclusion is shown in Figure 12 and its micro-shape
in Figures 13 and 14. The steel microstructure consisted
predominantly of sorbite and bainite.
More analyses were performed in collaboration with
the Institute of Metals and Technology Ljubljana. An
identical sample was analysed by emission electron
microscope JEOL JSM 6500F and an energy-dispersive
spectroscope – EDS INSA CRYSTAL 300. In Figure 15
the points of the analyses and in Table 6 the results of
the analyses are shown.
M. BALCAR et al.: THE QUALITY OF SUPER-CLEAN STEELS PRODUCED AT @ÏAS, inc.
Materiali in tehnologije / Materials and technology 46 (2012) 2, 169–175 173
Figure 11: Detail of extent and location of the defect on the generator
rotor shaft – forging No. 447 660
Slika 11: Velikost in mesto napake na gredi rotorja generatorja –
odkovek {t. 447 660
Figure 10: Defective generator rotor shaft – forging No. 447 660
Slika 10: Defektna gred rotorja generatorja – odkovek {t. 447 660
Figure 13: Micro-shape of the large part of inclusion (500-times)
Slika 13: Mikrooblika ve~jega dela vklju~ka (pove~ava 500-kratna)
Figure 12: Forging No. 447 660. Macro-shape of the sample at the
place of defect location.
Slika 12: Odkovek {t. 447 660. Vzorec z mestom napake.
Figure 14: Shorter rows of oxides were near the large inclusion
(500-times)
Slika 14: Kraj{i oksidni vklju~ek blizu ve~jega (pove~ava 500-kratna)
The chemical composition of the non-metallic–
ceramic materials used during the production of steel
was made for a comparison with the results of the
analysis of the chemical composition of the inclusions –
see Table 7.
On the basis of a comparison of the results of the
analysis in Tables 6 and 7 and the content of the basic
elements Si, Na and K it is possible to consider the
analyses on points 2 and 4 as inclusions based on the
casting powder PC20. Spectre 1 and 3 correspond to the
slide gate sand fill Chromix 8/5. Similar conclusions
were drawn also in the other 6 cases of unsatisfactory
shafts. From the description and the set of data for the
chemical composition of the impurities found in the
forgings for the shafts of steel 26NiCrMoV115, as
determined by emission electron microscope
JEOL JSM 6500F and by energy dispersive spectroscope
EDS INSA CRYSTAL 300, it was determined that the
main cause of the unacceptable defects of the forgings
was the occurrence of non-metallic particles with a
chemical composition corresponding to the casting
powder PC 20 and to the slide gate fill sand Chromix
8/5. The determination of the real causes of the
occurrence of this combination of non-metallic materials
in ingots and forging is the subject of further tests and
investigations.
6 CONCLUSIONS
In this work the production of super-clean steels at
ZDAS from the perspective of chemical composition is
evaluated. The chemical analyses of the melts steel were
compared with the chemical composition of the forgings.
M. BALCAR et al.: THE QUALITY OF SUPER-CLEAN STEELS PRODUCED AT @ÏAS, inc.
174 Materiali in tehnologije / Materials and technology 46 (2012) 2, 169–175
Figure 15: Points of analysis of inclusion
Slika 15: Mesta analize vklju~ka
Table 6: Chemical composition in the analysed points shown in Figure 15
Tabela 6: Kemi~na sestava v to~kah, ozna~enih na sliki 15
Spectre O Al Si K Mg Na Ca Cr Ti V Mn Fe Total
(w/%)
1 32.78 8.16 0.32 – 6.78 – – 32.02 – – – 19.93 100
2 44.93 9.00 24.85 0.73 3.61 2.77 1.26 0.56 0.57 – 11.72 – 100
3 29.87 10.20 0.47 – 5.79 – – 33.77 0.57 2.26 13.49 3.56 100
4 37.17 11.24 31.08 1.09 1.65 2.57 2.23 – – – 11.56 1.40 100
Spectre 1 – order of elements: Cr > Fe > Al > Mg > Si + O
Spectre 2 – order of elements: Si > Mn > Al > Mg > Na > Ca > K > Ti > Cr + O
Spectre 3 – order of elements: Cr > Mn > Al > Mg > Fe > V > Ti > Si + O
Spectre 4 – order of elements: Si > Mn > Al > Na > Ca > Mg > Fe > K + O
Table 7: Chemical composition of non-metallic materials used during the production of steel
Tabela 7: Kemi~na sestava nekovinskih materialov, uporabljenih pri izdelavi jekla
Sample O Al Si K Mg Na Ca Cr P S Mo Ti Fe Total
(w/%)
Refining slag VD-EU2 45.2 8.2 0.9 1.3 43.0 1.3 100
Refractory shotcrete of ref.
ladle – Kalinovo 55.2 22.9 6.8 3.0 9.6 2.5 100
Sand in slide gate
Chromix 8/5
30.2 8.3 1.5 7.5 33.6 0.2 18.7 100
Pouring channel – main
gate of the system 55.1 21.0 19.5 1.7 1.0 1.8 100
Mortar for gluing of
pouring channels – Regnalit 57.7 12.5 26.0 1.7 0.6 1.6 100
Mortar for gluing of
pouring channels – @ÏAS 52.9 16.4 24.2 0.7 4.7 0.6 0.8 100
Sand SiO2 58.8 0.2 40.5 0.2 0.3 100
Sand SiO2 – recycled 58.2 0.6 38.4 0.3 1.0 1.6 100
Casting powder PC 20 51.3 13.8 20.7 2.4 0.5 2.4 1.9 0.6 1.7 1.2 3.6 100
The agreement of both is acceptable for all elements,
with the exception of the contents of nitrogen and
particularly of oxygen. It can be concluded that the
difference could be resolved by a change of methodology
of taking the samples for a determination of the contents
of gases in the hot metal.
On the basis of the evaluation of the micro-cleanness
of the steel according to the standard DIN 50 602 by
method K4, a very good micro-cleanness K4 < 10 was
assessed for 45 heats out of 54 heats, thus for 83.3 % of
the production.
The metallographic analyses of 7 rejected rotors with
use of the electron microscope showed that 6 shafts out
of 7 were unsatisfactory due to the presence of isolated
massive rows of clusters of non-metallic particles with
lengths up to 15 mm consisting of 2 phases – casting
powder and chromite sand (Cr2O3), which was used as
fill sand for refining the ladle slide gate.
The measures for ensuring the stable level of
metallographic cleanliness and for the prevention of the
occurrence of exogenous inclusions may consist of the
optimisation of slag mode or in its possible modification,
as well as of interventions into casting technology,
including the solution of the issues for ceramics used for
the pouring and filtration of steel.
Acknowledgement
The investigations were performed within the
EUREKA program of the E!3192 ENSTEEL project,
identification number 1P04EO169 and project
FR-TI1/222.
7 REFERENCES
1 M. Balcar, R. @elezný, L. Martínek, J. Ba`an, Modelling of solidi-
fication process and chemical heterogeneity of 26NiCrMoV115 steel
ingot. 7th International Symposium Materials and Metallurgy,
Croatia, [ibenik, 2006, Metalurgija, 45 (2006) 3, 229
2 J. Ba`an, L. Socha, Evaluation of corrosion of refractory materials by
molten steel. 7th International Symposium Materials and Metallurgy,
Croatia, [ibenik, 2006, Metalurgija, 45 (2006) 3, 232
3 P. Fila, M. Balcar, L. Martínek, Náhled na oblast neúplných vlastních
nákladù ve vazbì na nìkteré technologické postupy výroby
elektrooceli [Insight into incomplete factory costs in relation to some
technological processes for production of electrical steel], 22nd
National conference with foreign participants : Teorie a praxe výroby
a zpracování oceli [Theory and practice of production and treatment
of steel], 4th – 5th April 2006, Ro`nov pod Radho{tìm. Tanger spol. s.
r. o. Ostrava, 2006, 257–261
M. BALCAR et al.: THE QUALITY OF SUPER-CLEAN STEELS PRODUCED AT @ÏAS, inc.
Materiali in tehnologije / Materials and technology 46 (2012) 2, 169–175 175

L. LAVTAR et al.: SIMULATIONS OF THE SHRINKAGE POROSITY OF Al-Si-Cu AUTOMOTIVE COMPONENTS
SIMULATIONS OF THE SHRINKAGE POROSITY OF
Al-Si-Cu AUTOMOTIVE COMPONENTS
MODELIRANJE KR^ILNE POROZNOSTI Al-Si-Cu
AVTOMOBILSKIH ULITKOV
Lejla Lavtar1, Mitja Petri~2, Jo`ef Medved2, Bo{tjan Taljat1, Primo` Mrvar2
1STEEL, d. o. o., Litostrojska cesta 60, SI-1000 Ljubljana, Slovenia
2Department of Material Science and Metallurgy, Faculty of Natural Sciences and Engineering, University of Ljubljana,
A{ker~eva cesta 12, SI-1000 Ljubljana, Slovenia
lavtar@steel.si
Prejem rokopisa – received: 2011-08-17; sprejem za objavo – accepted for publication: 2011-12-01
The 3D shoot-sleeve and shrinkage-porosity simulations of a high-pressure die-casting (HPDC) process are presented using the
ProCast casting-simulation software. The porosity was studied during the casting and solidification of aluminium-silicon-copper
alloy components in an H13 steel die. Excellent agreement between the simulated and experimental results was observed.
Keywords: high-pressure die casting, aluminium-silicon-copper alloy, shrinkage porosity, ProCast software
V tem prispevku je prikazano 3D-modeliranje pomika bata in kr~ilna poroznost procesa visokotla~nega litja (HPDC) z uporabo
programskega paketa ProCast. [tudija poroznosti je prikazana na aluminij-silicij-bakrovih ulitkih, litih v orodje iz jekla H13.
Analiza je pokazala zelo dobro ujemanje med modeliranimi in eksperimentalno ugotovljenimi rezultati.
Klju~ne besede: visokotla~no litje, aluminij-silicij-bakrova zlitina, kr~ilna poroznost, ProCast
1 INTRODUCTION
To manufacture a large variety of products with high
dimensional accuracy using the process of high-pressure
die casting (HPDC) the fast and economical production
of aluminium automotive components has been deve-
loped.1 In the past two decades the rapid development of
numerical simulation methodology and the solidification
simulation of castings have been introduced as an effec-
tive tool for modeling the casting process and improving
the quality of castings.2,3 The use of simulation software
saves time and reduces the costs of the casting-system
design and of the materials used.
The physical, mechanical and esthetic properties
directly depend on the metallurgical operating conditions
during casting. The combination of the mechanical
properties of the die-cast product, such as the die
temperature, the gate metal velocity, the applied casting
pressure, the cooling rate during die casting, the
geometrical complexity of the parts and the mold-filling
capacity, all affect the integrity of the cast components.
If these parameters are not controlled properly, various
defects in the finished component are to be expected.
The applied casting pressure is crucial during the
solidification of high-integrity parts. The effects of
process variables on the quality of cast components with
in-cavity pressure sensors, delay time and casting
velocity were examined by Dargusch in 2006. He found
that the porosity decreased with increasing intensifi-
cation pressure and increased with a higher casting
velocity.1,4
The porosity of castings can be examined with
destructive testing, with visual observation after
machining and non-destructing testing, like X-ray
microscopy and image-processing technology, which can
provide more detailed information about the pores. It is
also observed that the chemical composition of the alloy
affects the porosity of the cast components, the grain
refinement and the modification.5,6 Now it is commonly
accepted that the shrinkage and the gas are the two major
causes of porosity. The shrinkage porosity is associated
with the "hot spots" in the casting. The gas porosity is
caused by entrapped air in the injection system and the
cavity, the gas generated from burned lubricants, the
water in the cavity and hydrogen. The entrapped air is
the unwanted product of the high velocity of the alloy
caused by the turbulent flow during the injection process.
The paper describes a simulation of the HPDC of an
Al-Si9Cu3 casting in an H13 steel die and the com-
parison between the simulated and the experimental
porosity.
2 EXPERIMENTAL
2.1 Material and casting system
The alloy used for the die casting was an alumi-
nium-silicon-copper alloy (Table 1). The alloy is less
prone to shrinkage and internal shrinkage cavities and
has a very good castability. The ALSI H13 chromium
hot-work tool steel was used for the die. This steel has a
higher resistance to the heat cracking and die wear
caused by the thermal shock associated with the
Materiali in tehnologije / Materials and technology 46 (2012) 2, 177–180 177
UDK 621.74.043:669.71'782'3:539.21 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 46(2)177(2012)
die-casting process.7 The casting system with a shot
sleeve and a plunger are presented in Figure 1a. The
gates and runner system with two cavities are presented
in Figures 1b and 1c. The final product is an automotive
component (Figure 1d).
Table 1: Chemical composition in mass fractions of the Al-Si9Cu3
alloy, w/%
Tabela 1: Kemijska sestava Al-Si9Cu3 zlitine v masnih dele`ih, w/%
Si Cu Fe Mn Mg Zn Ni Cr
10.38 2.73 0.82 0.25 0.34 0.82 0.04 0.04
2.2 HPDC process
The casting process can be divided into four phases:
the pre-filling, the shot, the final pressure phase and the
ejection phase. In the pre-filling phase, the molten metal
is injected by a plunger, which forces the metal with a
low velocity through a horizontally mounted cylindrical
shot sleeve up to the gate. Usually, the shot sleeve is
partially filled with molten metal, the amount of which
depends on the cast component volume. The remaining
volume is empty. Previous research work has shown that
the fluid flow and the amount of empty space are
affected by the plunger motion, the shot-sleeve dimen-
sions and the amount of metal in the sleeve.8 In the
short-shot phase the plunger is accelerated to high
velocity and so any venting of the die cavity is practi-
cally impossible. In the final pressure phase, solidifi-
cation of the casting is completed and in the ejection
phase, the moulded part is removed, the die halves are
sprayed and positioned back to repeat the cycle.
The industrial HPDC process for casting an auto-
motive component starts with a plunger that has four
L. LAVTAR et al.: SIMULATIONS OF THE SHRINKAGE POROSITY OF Al-Si-Cu AUTOMOTIVE COMPONENTS
178 Materiali in tehnologije / Materials and technology 46 (2012) 2, 177–180
Figure 2: a) Shot profile with four different plunger speeds and b)
volume fraction picture of the alloy and the empty space in the shot
sleeve
Slika 2: a) Diagram pomika bata s {tirimi razli~nimi hitrostmi in b)
slika volumenskega dele`a zlitine in atmosfere v livni komori
Figure 1: Casting system: a) shot sleeve with plunger, b) gates and
runner system, c) the two cavities left and right and d) the casting
component
Slika 1: Ulivni sistem: a) livna komora z batom, b) ulivni in dovodni
kanal, c) dve livni votlini leva in desna in d) ulitek
Figure 3: a) Shot profile with three different plunger speeds and b)
volume fraction picture of the alloy and the empty space in the shot
sleeve
Slika 3: a) Diagram pomika bata s tremi razli~nimi hitrostmi in b)
slika volumenskega dele`a zlitine in atmosfere v livni komori
different speeds, as shown on the shot profile in Figure
2a. The volume fraction in Figure 2b shows that there
was no wave and no air entrapment.
3 RESULTS AND DISCUSSION
3.1 The shot-sleeve simulation
The same industrial HPDC process was then
simulated with the FEM-based software called ProCast.
The movement of the plunger was simulated using three
different plunger speeds. The simulation is shown on the
shot profile in Figure 3a. The volume fraction in Figure
3b shows no wave and no air entrapment.
The set-up time was minimized, the plunger speed in-
creased and the industrial HPDC process was shortened
by 0.48 s.
3.2 The shrinkage-porosity simulations
The shot-sleeve simulation results were used as the
boundary conditions for the cavity-filling simulations
and the shrinkage-porosity simulations, as the basic
study in this paper was the shrinkage porosity. Figure 4
shows the simulated shrinkage porosity "red spots" in the
left and right castings were the porosity spots are marked
with numbers. After nine cycles of casting constant
conditions in the die were established and after ten
cycles in the left-side casting two red spots of simulated
shrinkage porosity were examined (Figures 5 and 6) and
in the left casting (Figures 5b and 6b) a good agreement
with the simulated results of shrinkage porosity was
found (Figures 4, 5a and 6a).
4 CONCLUSIONS
In the present work the porosity of automotive
components was analyzed with ProCast, FEM-based
software. The most important conclusions that can be
drawn are:
• The shot-sleeve simulation gives valuable informa-
tion for the final quality of the components by
minimizing the volume fraction of the empty space
during the first stage of the HPDC process. The
volume fraction shows no wave and no air entrap-
ment.
• The shot-sleeve simulation gives savings in cycle
time by minimizing the set-up time during the shot
L. LAVTAR et al.: SIMULATIONS OF THE SHRINKAGE POROSITY OF Al-Si-Cu AUTOMOTIVE COMPONENTS
Materiali in tehnologije / Materials and technology 46 (2012) 2, 177–180 179
Figure 6: Shrinkage porosity in left casting at spot 3: a) simulation, b)
cut section
Slika 6: Kr~ilna poroznost v levem ulitku na mestu 3: a) modeliranje,
b) prerez
Figure 5: Shrinkage porosity in left casting at spot 1: a) simulation, b)
cut section
Slika 5: Kr~ilna poroznost v levem ulitku na mestu 1: a) modeliranje,
b) prerez
Figure 4: Shrinkage porosity simulation of: a) left and b) right
castings
Slika 4: Simulacija kr~ilne poroznosti a) na levem in b) desnem ulitku
stage of the HPDC process. The shot stage of the
HPDC process set-up time was shortened by 0.48 s.
• The shot-sleeve simulation also gives information
about the shrinkage-porosity location in castings,
called "red spots". The shrinkage porosity on the
sections of spots 1 and 3 in the left-side casting is in
good agreement with the simulated results.
5 REFERENCES
1 M. S. Dargusch, G. Dour, N. Schauer, C. M. Dinnis, G. Savage, J.
Mater. Process. Technol., 180 (2006), 37–43
2 T. R. Vijayaram, S. Sulaiman, A. M. S. Hamuda, J. Mater. Process.
Technol., 178 (2006), 29–33
3 L. A. Dobrzanski, M. Krupinski, J. H. Sokolowski, J. Mater. Process.
Technol., 167 (2005), 456–462
4 K. J. Laws, B. Gun, M. Ferry, Mater. Sci. Eng., A425 (2006),
114–120
5 P. W. Cleary, J. Ha, M. Prakash, T. Nguyen, Appl. Math. Model., 30
(2006), 1406–1427
6 M. Petri~, J. Medved, P. Mrvar, Metalurgija, 50 (2011) 2, 127–131
7 http://www.cintool.com/catalog/mold_quality/H13. pdf
8 W. Thorpe, V. Ahuja, M. Jahedi, P. Cleary, N. Stokes, Trans. 20th int.
die casting cong. & expo NADCA, Cleveland, 1999, T99-014
L. LAVTAR et al.: SIMULATIONS OF THE SHRINKAGE POROSITY OF Al-Si-Cu AUTOMOTIVE COMPONENTS
180 Materiali in tehnologije / Materials and technology 46 (2012) 2, 177–180
E. ALTUNCU et al.: WEAR-RESISTANT INTERMETALLIC ARC SPRAY COATINGS
WEAR-RESISTANT INTERMETALLIC ARC SPRAY
COATINGS
OBRABNA OBSTOJNOST INTERMETALNIH PREVLEK,
NAPR[ENIH V ELEKTRI^NEM OBLOKU
Ekrem Altuncu1-3, Sedat Iriç2, Fatih Ustel3
1Kocaeli University, Machine-Metal Tech., Kocaeli, Turkey,
2Sakarya University, Machine Eng. Dept., Sakarya, Turkey
3Sakarya University, Metall.-Materials Eng. Dept., Thermal Spray Center, Sakarya, Turkey
altuncu@kocaeli.edu.tr
Prejem rokopisa – received: 2011-10-20; sprejem za objavo – accepted for publication: 2011-12-08
The twin-wire electrical arc spraying (TWAS) process is widely used for worn-out surface restoration and the corrosion
protection of metallic constructions. The industrial benefit of arc spray coatings is the possibility of cost-effective coating
solutions to minimize corrosion problems. However, the wear resistance of metallic (such as Al, Cu and its alloys) arc sprayed
coatings is inadequate. Alloys including Cu-Al intermetallic coatings are new candidates for use in tribological environments
because of the combination of low cost and a remarkable resistance to abrasion under different working conditions. In this study
the tribological properties of Al-Cu twin-wire arc-spray coatings are investigated in dry sliding test conditions depending on the
load and the sliding distance.
Keywords: TWAS, intermetallic coatings, wear resistance
Elekri~no napr{evanje z dvojno `ico (TWAS) se {iroko uporablja za popravilo obrabljenih povr{in in protikorozijsko za{~ito
kovinskih konstrukcij. Industrijska prednost postopka je priprava poceni prevlek za zmanj{anje korozijskih te`av. Vendar
obrabna obstojnost kovinskih (Al, Cu in zlitine) napr{enih prevlek ni primerna. Intermetalne zlitine so nove kandidatke za
uporabo v tribolo{kih okoljih, ker zdru`ujejo nizko ceno in pomembno odpornost proti abrazivni obrabi. V tem delu so opisane
tribolo{ke lastnosti prevlek, napr{enih z dvojno `ico AlCu pri drsnem preizkusu v odvisnosti od obremenitve in drsne razdalje.
Klju~ne besede: TWAS, intermetalne prevleke, obrabna odpornost
1 INTRODUCTION
Wear-resistant coatings are used to reduce the
damage caused by abrasion, erosion, cavitation, and
fretting, also potentially associated with corrosion, and
in some cases to reduce friction1–3. The optimal wear
protection of light metallic substrates can be provided by
a cost-effective thermal spray coating process and
composition, depending on the operating environment
and working conditions4–9. Intermetallic coatings, alloy
coatings or metal-ceramic composite coatings can be
obtained by wire arc spraying with cored wires or
pre-alloyed wires. Cu-Al intermetallic systems were
actively researched for applications in the aviation,
automobile, naval, construction and defense sectors. The
Cu-Al alloy system has long been used for wheel
bearings for airplanes and screws for ships because of its
resistance to abrasion, corrosion, and heat7. The purpose
of this work is to develop an economical and effective
deposition method for copper-aluminum intermetallic
coatings to improve the wear resistance of light alloys.
The sliding wear resistance of Al-Cu intermetallic arc
spray coatings was investigated depending on load and
sliding distance. The crystal structure and composition of
the alloys were studied by x-ray diffraction.
2 EXPERIMENTAL DETAILS
The Sulzer Metco smart arc spray system we used
consists of a power supply, a control unit and a
robot-controlled arc spray gun. AISI 1020 low-carbon
steel and AlSi alloys with a thickness of 3 mm were used
in this study, and all the specimens to be coated were
pretreated by grit blasting. Aluminum and copper wires
with a diameter of 1.6 mm were sprayed with air used as
an atomizing gas (Table 1). The sliding wear test
(ASTM G133) conditions were as follows: sliding
stroke, 20 mm; sliding frequency, 5 Hz; and normal
Materiali in tehnologije / Materials and technology 46 (2012) 2, 181–183 181
Table 1: Arc Spray Process Parameters
Tabela 1: Parametri napr{evanja v elektri~nem obloku
Smart ArcSpray (Sulzer)
Current
(Ampere) 205–210
Voltage (Volt) 26–28
Spray Distance
(mm) 120–150
Gas Pressure
(bar) 4
UDK 621.793:539.231 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 46(2)181(2012)
loads, 10 N to 30 N. The resulting sliding distances were
200 m to 1 600 m. The thickness loss and weight loss
were measured on all the specimens under dry
conditions. The weight loss of the specimen after the test
was measured by an electronic analytical balance with a
minimum reading of 0.01 mg. The friction coefficients of
the coatings were measured using a ball-on-disc test in a
CSM tribotester.
3 RESULTS AND DISCUSSION
3.1 Microstructure of the coatings
Surface and cross-sectional SEM micrographs of the
arc-sprayed Cu-Al intermetallic coatings are shown in
Figure 1. It is clear that the coating is a mixture of white
and gray regions, which were identified as Cu and Al,
respectively, by EDS analysis. The microhardness of the
gray regions was found to be higher than that of the
white regions.
3.2 Cu-Al intermetallic phases
In the equilibrium phase diagram of Cu and Al
(Figure 2a) there are five stable intermetallic phases,
i.e., Cu9Al4, Cu3Al2, Cu4Al3, CuAl, and CuAl2, with two
terminal solid solutions of Cu(Al), which are often
designated as Cu and Al(Cu)8. In this study different
intermetallic phases were identified from XRD patterns.
These phases are: 00-025-0012; CuAl2, 00-024-000;
Cu9Al4, 00-050-1477; Cu3Al2, 00-002-1254; Al4Cu9
(JCPDS numbers). The main phase content of Cu9Al4
and Cu3Al2 intermetallics affected the wear resistance of
the coating. After a heat treatment at 400 oC for 3 h these
phase ratios were increased.
3.3 Comparative wear resistance
The effects of porosity, flattening ratio, oxide
content, and splat-to-splat bonding strength play an
important role in the coating’s antiwear performance.
Low cohesion and high porosity generally cause a large
piece of the coating to wear away and result in a
decrease of the wear resistance
The thickness loss of the specimens was determined
by measuring the cross-sectional thickness of the sound
material after testing using an optical micrometer to
observe accurately a cross-section through the central
E. ALTUNCU et al.: WEAR-RESISTANT INTERMETALLIC ARC SPRAY COATINGS
182 Materiali in tehnologije / Materials and technology 46 (2012) 2, 181–183
Figure 3: Mass-loss diagram as a function of wear load and sliding
distance for samples: a) as-sprayed and b) heat treated
Slika 3: Izguba mase v odvisnosti od obrabne obremenitve in razdalje
za vzorce: a) napr{ene in b) toplotno obdelane
Figure 2: a) Phase diagram of Al-Cu binary system8, b) XRD pattern
of the coating
Slika 2: a) Binarni fazni diagram Al-Cu8 in b) XRD-spekter prevlek
Figure1: Surface and cross-sectional SEM micrographs of the
arc-sprayed Cu-Al intermetallic coatings
Slika 1: SEM-posnetki povr{ine in prereza interemetalnih prevlek
Cu-Al, napr{enih v elektri~nem obloku
part of the track zone. The mass losses of the coatings
are shown in Figure 3a. At a low load of 10 N a very
small mass-loss difference was observed for sliding
distances between 200 m and 800 m. When the wear
load increased, the mass-loss difference increased. The
highest wear mass loss on the coating and thickness was
observed for 30 N at 800 m of sliding distance. The wear
mass-loss change after the heat treatment of the coating
is shown in Figure 3b. The heat-treated samples showed
a lower mass loss. The microstructure and phase content
of the coating have been suggested to influence the mass
loss. In Figure 4 the coefficient-of-friction (Cof) changes
are shown for the Al-Cu coating. As can be seen the Cof
values changed in the first stage of the test after which a
steady state is observed. The Cof values were measured
between 0.47 and 0.53. The heat-treated samples
exhibited lower Cof values between 0.41 and 0.45.
Figure 5 shows the wear-track profiles of the
coatings, both the track depth and width changed with an
increase of the load. The width of the wear tracks varied
between 1 650 μm and 1 730 μm at 400 m. When the
sliding distance increased to 800 m the width varied
between 1 747 μm and 2 015 μm. In both cases, the wear
track is rougher than the initial coating surface, which
indicates that particle pull-out took place during the
sliding. The morphology of the wear tracks of the arc
sprayed Al-Cu coatings confirmed that wear primarily
arose through cracked particles. The pull-out particles
then stayed in the contact area and led to three-body
abrasive wear, which was the main wear type in these
coatings.
4 CONCLUSION
Intermetallic coatings can be produced easily using
the twin-wire arc spray process. A process optimization
is required for a better coating quality. As a result of the
heat treatment of the Cu-Al arc spray coatings,
significant amounts of Cu4Al9 and Cu3Al2 intermetallic
phases were identified by XRD analysis. These phase
contents affected the wear mass loss and wear
track-profile width. The comparative mass loss as a
function of the wear load and sliding distance for both of
the heat treated coatings and original coatings were
determined. With the heat treatment we were able to
improve the wear resistance of the coating by a factor of
two.
5 REFERENCES
1 C. Bartuli, T. Valent, F. Casadei, M Tului, Advanced thermal spray
coatings for tribological applications, Proc. IMechE Part L: J.
Materials: Design and Applications, 221 (2007), 175–185
2 H. Pokhmurska, V. Dovhunyk, M. Student, E. Bielanska, Tribolo-
gical properties of arc sprayed coatings obtained from FeCrB and
FeCr-based power wires, Surf. Coat. Tech., 151–152 (2002),
490–494
3 S. Dallaire, Hard arc-sprayed coating with enhanced erosion and
abrasion wear resistance, Journal of Thermal Spray Technology, 10
(2001) 3, 511–519
4 H. Liao, B. Normand, C. Coddet, Influence of coating microstructure
on the abrasive wear resistance of WC/Co cermet coatings, Surf.
Coat. Technol., 124 (2000), 235–242
5 J. Voyer, B. R. Marple, Tribological performance of thermally
sprayed cermet coatings containing solid lubricants, Surf. Coat.
Technol., 127 (2000), 155–166
6 T. Sato, A. Nezu, T. Watanabe, Preparation of Ti-Al intermetallic
compund by wire arc spraying, Transactions of Materials Research
Society of Japan, 25 (2000) 1, 301–304
7 S. D. Sartale, M. Yoshitake, Investigation of Cu–Al surface alloy
formation on Cu substrate, J. Vac. Sci. Technol. B, 28 (2010),
353–360
8 Binary Alloy Phase Diagrams, edited by T. B. Massalski, H. Oka-
moto, P. R. Subramanian, L. Kcprzak, ASM International, Metals
Park, OH, 1990
9 B. Q. Wang, M. W. Seitz, Comparison in erosion behavior of iron-
base coatings sprayed by three different arc-spray processes, Wear,
250 (2001), 755–761
Materiali in tehnologije / Materials and technology 46 (2012) 2, 181–183 183
E. ALTUNCU et al.: WEAR-RESISTANT INTERMETALLIC ARC SPRAY COATINGS
Figure 5: Wear-track profile views
Slika 5: Videz profila obrabnih poti
Figure 4: Cof of the coatings
Slika 4: Torni koeficient (Cof) prevlek

S. BUTKOVI] et al.: EFFECT OF SINTERING PARAMETERS ON THE DENSITY ...
EFFECT OF SINTERING PARAMETERS ON THE
DENSITY, MICROSTRUCTURE AND MECHANICAL
PROPERTIES OF THE NIOBIUM-MODIFIED
HEAT-RESISTANT STAINLESS STEEL GX40CrNiSi25-20
PRODUCED BY MIM TECHNOLOGY
VPLIVI PARAMETROV SINTRANJA NA GOSTOTO,
MIKROSTUKTURO IN MEHANSKE LASTNOSTI Z NIOBIJEM
LEGIRANGA NERJAVNEGA OGNJEVZDR@NEGA JEKLA
GX40CrNiSi25-20, IZDELANEGA Z MIM-TEHNOLOGIJO
Samir Butkovi}1, Mirsada Oru~2, Emir [ari}1, Muhamed Mehmedovi}1
1University of Tuzla, Faculty of Mechanical Engineering, Univerzitetska 4, Tuzla, BiH
2University of Zenica, Institute of Metallurgy, Kemal Kapetanovi} Zenica, Travni~ka cesta 7, 72000 Zenica, BiH
samir.butkovic@untz.ba
Prejem rokopisa – received: 2011-10-20; sprejem za objavo – accepted for publication: 2011-12-22
Properties of heat-resistant, stainless-steel parts produced by the metal-injection-molding (MIM) process depend mostly on the
sintering parameters. The effect of these sintering parameters on the densification, microstructure, hardness and tensile
properties of the niobium-modified, heat-resistant stainless steel GX40CrNiSi25-20 were investigated. The prepared feedstock
was injection molded to obtain tensile test specimens (ISO 2740). The debinding of the molded parts was performed using the
catalytic debinding method, while the residual binder was removed by thermal debinding. The sintering was performed at 1200
°C and 1310 °C, in argon (Ar), hydrogen (H2) and nitrogen (N2) atmospheres, with the sintering times between 3 h and 6 h. It
was found that sintering in a nitrogen atmosphere increased the strength and reduced the ductility. The mechanical properties
were also enhanced by a higher sintering temperature (1310 °C), due to the positive effects of the pore rounding and increased
density. The prolonged sintering time caused changes in the grain size, but had little effect on the density. Faster sintering and
improved ductility was observed for the samples sintered in hydrogen and argon atmospheres.
Keywords: metal injection molding, heat-resistant stainless steel, sintering parameters, density, microstructure, mechanical
properties, metal powder
Lastnosti izdelkov iz ognjevzdr`nega nerjavnega jekla, izdelanih s tehnologijo brizganja pra{kastih materialov (MIM) so v
najve~ji meri odvisne od parametrov sintranja. Vplivi parametrov sintranja na zgo{~evanje, mikrostrukturo, trdoto in trdnostne
lastnosti z niobijem legiranega nerjavnega ognjevzdr`nega jekla GX40CrNiSi25-20 so opisani v tem ~lanku. Iz pripravljenega
surovega materiala so bili nabrizgani vzorci za preizkuse z injekcijskim ulivanjem. Sintranje je bilo izvedeno pri temperaturah
1200 °C in 1310 °C v argonski (Ar), vodikovi (H2) in du{ikovi (N2) atmosferi v trajanju med 3 h in 6 h. Ugotovljeno je bilo, da
sintranje v du{ikovi atmosferi povzro~i utrjevanje materiala in zmanj{anje duktilnosti. Mehanske lastnosti so bile pove~ane tudi
pri vi{ji temperaturi sintranja (1310 °C) zaradi zaokro`evanja por in pove~anja gostote. Podalj{an ~as sintranja je povzro~il
spremembe v velikosti zrn, vendar je imel majhen vpliv na gostoto sintranih delov. Hitrej{e sintranje in izbolj{ana plasti~nost je
bila pri vzorcih, ki so bili sintrani v vodikovi in argonski atmosferi.
Klju~ne besede: pra{kasti materiali, ognjvzdr`no nerjavno jeklo, sintranje, gostota, mikrostruktura, mehanske lastnosti
1 INTRODUCTION
The metal-injection-molding (MIM) process
combines the advantages of polymer injection molding
with the material flexibility of powder metallurgy. MIM
technology enables the mixing of different metal
powders with different binders and allows the processing
of materials with complex mechanical, thermal, wear and
magnetic properties. It is a cost-effective process in the
high-volume production of small and complex-shaped
parts.1–3 The application of MIM technology in the
production of complex-shaped components from
heat-resistant stainless steel represent a very attractive
solution to reduce manufacturing costs and overcome the
machinability problems.
Heat-resistant stainless steels are selected for a wide
range of applications because of their superior resistance
to creep, corrosion and oxidation. Heat-resistant stainless
steel GX40CrNiSi25-20 modified with niobium belongs
to the group of precipitation-hardened steels. This high-
temperature alloy provides strength and oxidation resist-
ance at temperatures up to 1200 °C and it is generally
used in the presence of combustion gases that may be
generated from exhaust and pollution-control equipment.
There are four main steps that make up the MIM
process: preparation of feedstock (mixing of metal
powder with binder), injection molding, solvent/che-
mical debinding and sintering.4 The most complex step
of MIM technology in producing stainless-steel parts is
sintering. Sintering parameters, such as, heating and
Materiali in tehnologije / Materials and technology 46 (2012) 2, 185–190 185
UDK 621.762.5:669.14.018.8 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 46(2)185(2012)
cooling rate, sintering time, sintering atmosphere, sinter-
ing temperature, partial pressure of sintering atmosphere,
affect the mechanical and physical properties, as well as
corrosion and the heat resistance of sintered parts.5 The
wrong sintering parameters can lead to a low density, the
absorption and desorption of some elements, deteriorated
mechanical properties, reduced corrosion and oxidation
resistance and a reduced service time.5 Also, if the
optimal parameters are not selected, increased sintering
costs can be expected.
In this regard, the effect of sintering temperature,
time and atmosphere on the physical and mechanical
properties of metal-injection-molded heat-resistant
stainless steel GX40CrNiSi25-20 modified with niobium
was investigated in this work.
2 EXPERIMENTAL WORK
2.1 Material
Niobium-modified GX40CrNiSi25-20 is a heat-
resistant stainless steel with a high resistance to creep
and oxidation. A higher percent of chromium provides
superior oxidation resistance, while a higher percent of
carbon, compared to ordinary austenitic stainless steel,
provides high strength and creep resistance.6,7 The
niobium addition gives structural stability and precipi-
tation hardening. The austenitic structure provides the
strength and structural stability at elevated temperatures.
This steel is intended for high-temperature applications,
up to 1200 °C, e.g., turbine blades or furnace parts. The
production of feedstock involved the mixing of
pre-alloyed metal powder with a suitable binder. The
typical chemical composition after sintering is presented
in Table 1.
Table 1: Typical chemical composition after sintering niobium-
modified GX40CrNiSi25-20 in mass fractions, w/%
Tabela 1: Tipi~na kemi~na sestava sintranega jekla GX40CrNiSi25-20
s povi{anim niobijem v masnih dele`ih, w/%
C Cr Ni Si Nb Fe
0.2–0.5 24–26 19–22 0.75–1.3 1.2–1.5 Bal
2.2 Injection molding
The injection molding of the prepared feedstock was
made in a machine for the injection molding of metal
powders type ALLROUNDER 320 C 600–100. The
shape of the mold cavity corresponds to a standard spe-
cimen for tensile tests used in powder metallurgy (ISO
2740).
At the beginning of the injection-molding process the
material is heated to 185 °C at the nozzle of the barrel
and with a screw rotation transported at the barrel top.
The back pressure of the accumulated material was 30
bar. After the accumulation of a sufficient quantity of
material, the screw stops rotating and moves forward,
pushing the melted material into the tool cavity at a tool
temperature of 110–115 °C. The holding pressure of 900
bar and its duration of 3 s provide for complete cavity
filling. After the injection molding, the parts were cooled
for 27 s.
The average density of the injection-molded parts
was 5.5 g/cm3.
2.3 Catalytic debinding
After the injection molding the parts were debound
by a catalytic debinding method. Polyacetal, as a main
component of the feedstock, was decomposed by nitric
acid at a temperature below the melting point of
polyacetal, preventing the parts from deformation during
the debinding. Usually, a small concentration of a
backbone polymer, that is unaffected by the catalyst, is
included to hold the particles together until the material
is ready to start forming necks between the particles by
diffusion (often polyethylene).
The debinding was performed at 120 °C, where
vaporized acid exists in the atmosphere. A nitrogen flow
rate of 50 L/min was used to spread the nitric acid over
the debinding furnace chamber. The nitric acid flow rate
was 3.4 mL/min. The preheating time and the debinding
time used during the debinding were 30 min and 4 h,
respectively. A purging time of 30 min was used to
remove the residual acid to the combustion chamber,
where it was burned by propane gas.
2.4 Thermal debinding and sintering
The thermal debinding and sintering were performed
in the same furnace type MIM 3045. The parts were
gradually heated to 600 °C, where the residual binder
starts to degrade and annealed for 2 h. After the residual
binder was totally removed the parts were heated to the
sintering temperature. The sintering is a process with the
temperature, time and sintering atmosphere as influential
factors. The thermal cycle of the debinding was the same
for all the experiments (Figure 1). The sintering tem-
peratures were 1200 °C and 1310 °C and the sintering
S. BUTKOVI] et al.: EFFECT OF SINTERING PARAMETERS ON THE DENSITY ...
186 Materiali in tehnologije / Materials and technology 46 (2012) 2, 185–190
Figure 1: Thermal cycle of sintering and debinding
Slika 1: Termalna cikla sintranja in odstranjevanja veziva
times were 3 h and 6 h. The atmospheres used during the
experiments were hydrogen, nitrogen, and argon.
3 RESULTS AND DISCUSSION
3.1 Density
The density of the sintered parts was measured by
Archimedes’ immersion method. The influence of the
sintering atmosphere, temperature and time on the
density of the investigated material was performed and
analyzed using a general full factorial experiment.
Analysis of variance (ANOVA) was used to demonstrate
the significance level of the variables, as well as the
effect of the sintering variables on the sintered density.
The influential factors, their levels and average density
for a selected set of parameters are presented in Table 2.
Table 2: Sintering conditions and density of sintered alloy
Tabela 2: Pogoji sintranja in gostota sintrane zlitine
Atmosphere
(A)
Temperature,
°C (B) Time, h (C)
Average
density, g/cm3
Ar 1 200 3 6.89
H2 1 200 3 7.27
N2 1 200 3 6.59
Ar 1 310 3 7.71
H2 1 310 3 7.80
N2 1 310 3 7.75
N2 1 310 6 7.66
H2 1 310 6 7.77
Ar 1 310 6 7.72
N2 1 200 6 7.16
H2 1 200 6 7.44
Ar 1 200 6 7.17
Generally, all the sintering variables have a
significant effect on the sintered density. The ANOVA
(Table 3) showed that the sintering temperature has the
highest influence on the sintered density (72.6 %),
followed by the sintering atmosphere (9.3 %), the
sintering time (4.1 %) and the two-factor and three-factor
interactions. The influence of the sintering temperature,
time and atmosphere on the density of the sintered parts
is shown in Figure 2. It was found that sintering time has
a significant effect on the sintered density only at 1200 °C,
increasing the average sintered density from 6.91 g/cm3
to 7.26 g/cm3. A longer sintering time at 1310 °C caused
insignificant changes to the density, which has to be
taken into account during the sintering-costs analysis.
Higher sintering temperatures caused a more
intensive atomic diffusion, resulting in faster sintering
and higher resulting densities. Increasing the sintering
temperature from 1200 °C to 1310 °C resulted in the
average density increasing from 7.09 g/cm3 to 7.73 g/cm3
(average of all runs). Statistical data indicated that the
sintering atmosphere also has a significant effect on the
sintered density. Sintering in hydrogen gave a density
higher than the sintered densities achieved in the
nitrogen and argon atmospheres. This difference is
particularly emphasized at a temperature of 1200 °C and
a time of 3 h, where sintering in hydrogen gave a density
of 7.27 g/cm3, while sintering in nitrogen and argon gave
densities of 6.59 g/cm3 and 6.89 g/cm3, respectively.
Small hydrogen atoms diffuse into the metal lattice and
do not inhibit the elimination of final porosity1. Argon
and nitrogen remains in the final pores and build the
internal pressure, whereby the elimination of the porosity
in the final stage of sintering is inhibited.
For the nitrogen samples the sintering activity was
probably impeded by the absorbed nitrogen, which
S. BUTKOVI] et al.: EFFECT OF SINTERING PARAMETERS ON THE DENSITY ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 185–190 187
Figure 2: Influence of sintering temperature, time and atmosphere on
the density of the sintered parts
Slika 2: Vpliv temperature, ~asa in atmosfere sintranja na gostoto
sintranih delov
Table 3: Analysis of Variance for density
Tabela 3: Analiza variance za gostoto
Term DOF SumSqr MeanSqr P F Contribution, %
A 2 0.324213 0.162106 < 0.0001 111.5602 9.355652
B 1 2.516833 2.516833 < 0.0001 1732.064 72.62707
C 1 0.142913 0.142913 < 0.0001 98.35132 4.123965
AB 2 0.159606 0.079803 < 0.0001 54.9198 4.605679
AC 2 0.032124 0.016062 0.0019 11.05376 0.92699
BC 1 0.211313 0.211313 < 0.0001 145.4236 6.097752
ABC 2 0.060982 0.030491 0.0001 20.98351 1.759717
Pure Error 12 0.017437 0.503171
Residuals 12 0.017437 0.001453
obviously reduces the mass-transport mechanism during
sintering.5 Also, it is well known that hydrogen is the
most effective reducing atmosphere causing effective
oxide removal, which results in a longer sintering dwell
time. The highest density of 7.8 g/cm3, which is 98.7 %
of theoretical density, was achieved using a hydrogen
atmosphere and a temperature of 1310 °C.
3.2 Mechanical properties
Parameters of sintering at the final stage of the MIM
(Metal Injection Molding) process have a decisive
influence on the mechanical properties of the produced
parts. The characteristics of the residual porosity, che-
mical composition, structure after sintering and the
density of the sintered parts are factors that make heat-
resistant stainless steel very sensitive to the sintering
parameters. Tensile tests were made on standard tensile
tests samples used in powder metallurgy.
Table 4 shows the tensile and elongation results of
the samples sintered in nitrogen, hydrogen and argon
atmospheres at 1200 °C and 1310 °C. All the samples
were sintered for 3 h in an atmosphere with 400 mbar of
partial pressure.
It was found that the tensile and yield strengths
increased with an increase of the temperature. The
average tensile strength for parts sintered in argon and
nitrogen was increased from 317 MPa to 680 MPa with a
change of the sintering temperature from 1200 to 1310
°C. The tensile strength increase is the result mainly of
increased density and a significant reduction of porosity.
Also, a reduction of porosity by increasing the tempera-
ture caused a substantial improvement in the elongation
of the sintered parts (Figure 3).
The reduction of the temperature leads to a large drop
in the elongation and strength of the material. It is
evident that sintering at a temperature of 1200 °C and in
a nitrogen atmosphere gives an elongation of 2.2 %,
while sintering in argon achieved a maximum elongation
of 3.5 %. A slight improvement was observed for the
samples sintered in hydrogen atmospheres (5.3 %),
because of the higher density compared to the densities
of samples sintered in argon and nitrogen. The low
density and the sharp edges of the residual porosity of
the parts sintered at lower temperatures caused a stress
concentration during the tensile test, making the material
very brittle. Analyzing and comparing the tensile test
results, it is evident that the samples sintered in a
nitrogen atmosphere experienced strengthening during
the sintering. The best tensile strength of 777 MPa and
yield strength of 412 MPa were achieved using a nitro-
gen atmosphere and a sintering temperature of 1310 °C
Sintering in a nitrogen atmosphere caused the
absorption of nitrogen, resulting in the solid solution
strengthening of the material. Also, based on previous
research, the formation of NbCrN and precipitation
strengthening were also possible8. The strengthening of
the material was avoided using an argon atmosphere,
where the average tensile strength and yield strength of
583 MPa and 223 MPa were achieved, respectively. A
substantial elongation improvement was also observed
on samples sintered in an argon atmosphere and tem-
perature 1310 °C.
In order to see the effect of nitrogen absorption on
the mechanical properties of the sintered parts, additional
experiments were made. The main condition for
intensifying the absorption and increasing the nitrogen
content in the steel is to increase the partial pressure of
the nitrogen atmosphere. In this regard, the partial
pressure of nitrogen in the sintering furnace chamber
was increased from 400 mbar to 600 mbar. After
sintering, the hardness was measured and the results are
presented in Table 5. The increasing of the partial
pressure of nitrogen caused an increasing of the hardness
of the sintered parts (Figure 4). A higher nitrogen level
contained in the steel after sintering, as a result of the
increased partial pressure of the nitrogen atmosphere,
caused a more intensive strengthening of steel. It can be
concluded that the change in the mechanical properties
of the sintered heat-resistance stainless steel is possible
through the partial pressure of the nitrogen atmosphere.
It was also observed that the sintering temperature
has a very significant effect on the hardness of the
sintered parts. The hardness increased with a higher
sintering temperature and the average hardness increased
S. BUTKOVI] et al.: EFFECT OF SINTERING PARAMETERS ON THE DENSITY ...
188 Materiali in tehnologije / Materials and technology 46 (2012) 2, 185–190
Figure 4: Influence of nitrogen partial pressure on the hardness of the
sintered parts
Slika 4: Vpliv parcialnega tlaka du{ika na trdoto sintranih delov
Figure 3: Influence of density on the elongation of the sintered parts
Slika 3: Vpliv gostote na raztezek sintranih kosov
from 180 HV1 to 230 HV1 with a temperature increase
from 1200 °C to 1310 °C.
Table 4: Mechanical properties after sintering
Tabela 4: Mehanske lastnosti po sintranju
Tempe-
rature
/oC
Atmo-
sphere
Partial
pressure
/mbar
Average
tensile
strength
Rm/MPa
Average
yield
strength
Re/MPa
Average
elongat-
ion
A/%
1310 Ar 400 583 223 38
1310 N2 400 777 412 27
1200 H2 400 359 245 5.3
1200 N2 400 292 – 2.2
1200 Ar 400 291 217 3.5
Table 5: Comparison of the hardness for different partial pressures of
the nitrogen atmosphere and temperatures
Tabela 5: Primerjava trdote za razli~ne parcialne tlake du{ika in
temperature
Partial pressure
/mbar Temperature /°C Hardness /HV1
400 1200 180
400 1310 230
600 1310 255
3.3 Microstructure
The microstructure of the samples sintered at a
temperature of 1200 °C and an argon atmosphere
(Figure 5a) reveals an insufficient degree of sintering
with a noticeable residual porosity and a very small grain
size with a density of 6.89 g/cm3. Insufficient connection
between the grains caused a tensile and yield strength
reduction and the elongation of parts sintered at 1200 °C.
Micrograph of the parts sintered at a temperature of 1310
°C (Figures 5b, d and e) reveal grain growth and a fully
austenitic microstructure with a minimal residual
porosity and a density of 7.71 g/cm3. A small percent of
residual porosity indicates that the material reached
almost theoretical density with well connected grains
and better mechanical properties.
The parts sintered in hydrogen and argon experienced
a more intensive grain growth compared to the parts
sintered in nitrogen. Clean grain boundaries facilitate
grain-boundary movement, resulting in effective pore
absorption, higher density and larger grains compared to
the parts sintered in nitrogen. Impeded sintering activity
and reduced mass transport caused slower motion of the
grain boundaries of the parts sintered in nitrogen,
resulting in a reduced sintering density. The micro-
structure of parts sintered at 1310 °C is comparable to
the microstructure of the wrought material.
A prolonged sintering time caused a slight grain
coarsening (Figures 5c and f). Also, some of the pores
observed on the samples sintered for 6 h coarsened
maybe as a consequence of an extended sintering time
and vacancy diffusion from smaller to bigger pores.9,10
The micrographs of parts sintered in nitrogen reveal an
austenitic microstructure with Cr2N regions created as a
consequence of nitrogen absorption.
4 CONCLUSION
The density of the heat-resistant stainless steel
GX40CrNiSi 25–20 produced by the MIM process
depends mostly on the sintering temperature. Increasing
the average density from 7.09 g/cm3 to 7.73 g/cm3 was
achieved by increasing the sintering temperature from
1200 °C to 1310 °C. Sintering in hydrogen and argon
resulted in higher densities and better ductility of the
sintered parts compared to the nitrogen atmosphere.
S. BUTKOVI] et al.: EFFECT OF SINTERING PARAMETERS ON THE DENSITY ...
Materiali in tehnologije / Materials and technology 46 (2012) 2, 185–190 189
Figure 5: Microstructure of sintered parts for different conditions: a) argon, 1200 °C, 3 h, b) argon, 1310 °C, 3 h, c) argon, 1310 °C, 6 h, d)
nitrogen, 1310 °C, 3 h, e) hydrogen, 1310 °C, 3 h, f) hydrogen, 1310 °C, 6 h, glyceregia
Slika 5: Mikrostruktura sintranih kosov pri razli~nih razmerah: a) argon, 1200 °C, 3 h, b) argon, 1310 °C, 3 h, c) argon, 1310 °C, 6 h, d) du{ik,
1310 °C, 3 h, e) vodik, 1310 °C, 3 h, f) vodik, 1310 °C, 6 h, glyceregia
Insufficient density of parts sintered at a temperature of
1200 °C caused brittleness of the steel with a maximum
elongation of 5.3 %. A superior tensile and yield strength
were obtained by sintering in a nitrogen atmosphere. The
maximum tensile strength of 777 MPa and yield strength
of 412 MPa were achieved using a nitrogen atmosphere
with 400 mbar of partial pressure and a temperature of
1310 °C. Strengthening also depended on the nitrogen
partial pressure. The hardness was increased by 10 %
when the nitrogen partial pressure was changed from 400
mbar to 600 mbar. It was found that a longer sintering
time at a temperature of 1310 °C had a minor effect on
the density of the sintered parts. Also, it is very
important to emphasize that the prolongation of the
sintering time at a temperature of 1200 °C, from 3 h to
6 h, increased the sintered density from 6.91 g/cm3 to
7.26 g/cm3, which is still much less than the density
achieved at a temperature of 1310 °C. This is very
important during the optimization of the sintering profile
and indicates the importance of using a higher
temperature to reduce the sintering time and the sintering
costs.
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190 Materiali in tehnologije / Materials and technology 46 (2012) 2, 185–190