VSEBINA – CONTENTS
IZVIRNI ZNANSTVENI ^LANKI – ORIGINAL SCIENTIFIC ARTICLES
New discovered paradoxes in theory of balancing chemical reactions
Novoodkriti paradoksi v teoriji uravnote`enja kemijskih reakcij
I. B. Risteski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 503
Characteristics of creep in conditions of long operation
Zna~ilnosti lezenja pri dolgotrajni uporabi
N. A. Katanaha, L. B. Getsov . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 523
A thermodynamic and kinetic study of the solidification and decarburization of malleable cast iron
Termodinami~na in kineti~na analiza strjevanja in razoglji~enja belega litega `eleza
M. Pirnat, P. Mrvar, J. Medved . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 529
Modelling and preparation of core foamed Al panels with accumulative hot-roll bonded precursors
Na~rtovanje in izdelava Al-panelov s sredico iz Al-pen na osnovi ve~stopenjsko toplo valjanih prekurzorjev
V. Kevorkijan, U. Kova~ec, I. Paulin, S. D. [kapin, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 537
Numerical solution of hot shape rolling of steel
Numeri~na re{itev vro~ega valjanja jekla
U. Hanoglu, S. Islam, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 545
Solidification and precipitation behaviour in the AlSi9Cu3 alloy with various Ce additions
Strjevanje in izlo~anje v zlitini ALSI9CU3 pri razli~nih dodatkih Ce
M. Von~ina, S. Kores, P. Mrvar, J. Medved . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 549
Effect of change of carbide particles spacing and distribution on creep rate of martensite creep resistant steels
Vpliv spremembe razdalje med karbidnimi izlo~ki in njihove porazdelitve na hitrost lezenja martenzitnih jekel, odpornih proti lezenju
D. A. Skobir Balanti~, M. Jenko, F. Vodopivec, R. Celin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 555
Stress-strain analysis of an abutment tooth with rest seat prepared in a composite restoration
Napetostno-deformacijska analiza opornega zoba z zapornim sede`em, izdelana s kompozitnim popravilom
Lj. Tiha~ek [oji}, A. M. Lemi}, D. Stamenkovi}, V. Lazi}, R. Rudolf, A. Todorovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 561
Identification and verification of the composite material parameters for the Ladevèze damage model
Identifikacija in verifikacija parametrov kompozitnega materiala za model Ladevèze
V. Kleisner, R. Zem~ík, T. Kroupa. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 567
Evaluation of the strength variation of normal and lightweight self-compacting concrete in full scale walls
Ocena variacije trdnosti normalnega in lahkega vibriranega betona v polnih stenah
M. M. Ranjbar, M. Hosseinali Beygi, I. M. Nikbin , M. Rezvani, A. Barari . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 571
The influence of buffer layer on the properties of surface welded joint of high-carbon steel
Vpliv vmesne plasti na lastnosti povr{inskih zvarov jekla z veliko ogljika
O. Popovi}, R. Proki} - Cvetkovi}, A. Sedmak, G. Buyukyildrim, A. Bukvi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 579
Corrosion stability of different bronzes in simulated urban rain
Korozijska stabilnost razli~nih bronov v umetnem kislem de`ju
E. [vara Fabjan, T. Kosec, V. Kuhar, A. Legat . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 585
Morphology and corrosion properties PVD Cr-N coatings deposited on aluminium alloys
Morfologija in korozijske lastnosti CrN PVD-prevlek, nanesenih na aluminijeve zlitine
D. Kek Merl, I. Milo{ev, P. Panjan, F. Zupani~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 593
STROKOVNI ^LANKI – PROFESSIONAL ARTICLES
The effect of electromagnetic stirring on the crystallization of concast billets
Vpliv elektromagnetnega me{anja na kristalizacijo kontinuirno ulitih gredic
F. Kavicka, K. Stransky, B. Sekanina, J. Stetina, V. Gontarev, T. Mauder, M. Masarik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 599
Wear of refractory materials for ceramic filters of different porosity in contact with hot metal
Obraba ognjevzdr`nega materiala kerami~nih filtrov z razli~no poroznostjo v stiku z vro~o kovino
J. Ba`an, L. Socha, L. Martínek, P. Fila, M. Balcar, J. Chmelaø . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 603
ISSN 1580-2949
UDK 669+666+678+53 MTAEC9, 45(6)501–659(2011)
MATER.
TEHNOL.
LETNIK
VOLUME 45
[TEV.
NO. 6
STR.
P. 501–659
LJUBLJANA
SLOVENIJA
NOV.–DEC.
2011
The influence of the mineral content of clay from the white bauxite mine on the properties of the sintered product
Vpliv vsebnosti minerala gline iz rudnika belega boksita na lastnosti sintranega proizvoda
M. Krgovi}, I. Bo{kovi}, M. Vuk~evi}, R. Zejak, M. Kne`evi}, R. Mitrovi}, B. Zlati~anin, N. Ja}imovi} . . . . . . . . . . . . . . . . . . . . . . . . 609
Effect of pre-straining on the springback behavior of the AA5754-0 alloy
Vpliv prenapenjanja na povratno elasti~no izravnavo zlitine AA5754-0
S. Toros, M. Alkan, R. E. Ece, F. Ozturk . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 613
Heat treatment and mechanical properties of heavy forgings from A694–F60 steel
Toplotna obdelava in mehanske lastnosti te`kih izkovkov iz jekla A694-F60
M. Balcar, J. Novák, L. Sochor, P. Fila, L. Martínek, J. Ba`an, L. Socha, D. A. Skobir Balanti~, M. Godec . . . . . . . . . . . . . . . . . . . . . . 619
The tensile behaviour of friction-stir- welded dissimilar aluminium alloys
Natezne zna~ilnosti tornih pomi~nih zvarov razli~nih aluminijevih zlitin
R. Palanivel, P. Koshy Mathews. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 623
Screen-printed electrically conductive functionalities in paper substrates
Elektroprevodne oblike, pripravljene s sitotiskom na papirnih podlagah
M. @vegli~, N. Hauptman, M. Ma~ek, M. Klanj{ek Gunde . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 627
Recent growing demand for magnesium in the automotive industry
Rast povpra{evanja po magneziju v avtomobilski industriji
M. J. Freiría Gándara . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 633
Contact with chlorinated water: selection of the appropriate steel
Kontakt s klorirano vodo – izbor ustreznega jekla
L. Gosar, D. Drev . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 639
LETNO KAZALO – INDEX
Letnik 45 (2011), 1–6 – Volume 45 (2011), 1–6 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 645
I. B. RISTESKI: NEW DISCOVERED PARADOXES IN THEORY OF BALANCING CHEMICAL REACTIONS
NEW DISCOVERED PARADOXES IN THEORY OF
BALANCING CHEMICAL REACTIONS
NOVOODKRITI PARADOKSI V TEORIJI URAVNOTE@ENJA
KEMIJSKIH REAKCIJ
Ice B. Risteski
2 Milepost Place # 606, Toronto, Ontario, Canada M4H 1C7
ice@scientist.com
»Failure to distinguish carefully between mathematical
and metamathematical statements leads to paradoxes.«
M. Kac, S. M. Ulam, Mathematics and Logic,
Dover Pub. Inc., Mineola 1992, p. 126.
Prejem rokopisa – received: 2011-01-24; sprejem za objavo – accepted for publication: 2011-09-21
In this work are given new paradoxes and fallacies discovered in the theory of balancing chemical reactions. All the
counterexamples showed that so–called »chemical procedures« of balancing chemical reactions given in earlier chemical
literature are inconsistent. Balancing chemical reactions is a mathematical procedure independent of chemistry. In order to avoid
the appearance of paradoxes, chemical reactions must be considered as a formal system founded by virtue of well–defined
mathematical model. The results obtained in this work affirmed that the usage of traditional ways of balancing chemical
reactions is limited. They may be used only for balancing some elementary chemical equations. In other words, foundation of
chemistry looks for a new approach of balancing chemical reactions, which must be completely different than current »chemical
procedures«. This work is a collection and analysis of some paradoxes and fallacies which appeared in the theory of balancing
chemical reactions.
Keywords: chemical reactions, paradoxes, balancing, fallacies
Predstavljeni so novi paradoksi in zmote, odkrite v teoriji uravnote`enje kemijskih reakcij. Vsi protiprimeri so pokazali, da so
tako nekonsistentne tako imenovane kemijske procedure uravnote`enja kemijskih reakcij, navedene v zgodnjih virih.
Uravnote`enje kemijskih reakcij je matemati~na procedura, neodvisna od kemije. Da bi se izognili nastajanju paradoksov, je
treba kemijske reakcije formulirati kot formalen sistem na podlagi dobro definiranega matemati~nega modela. Rezultati v tem
delu potrjujejo, da je uporaba tradicionalnih na~inov uravnote`enja kemijskih reakcij omejena. Uporabljamo jo le za
uravnote`enje nekaterih elementarnih kemijskih reakcij. Z drugih besedami, temelj kemije i{~e nove pribli`ke za uravnote`enje
kemijskih reakcij, ki se morajo razlikovati od sedanjih kemijskih procedur. To delo je izbor in analiza nekaterih paradoksov in
zmote, ki so se pojavile v teoriji uravnote`enja kemijskih reakcij.
Klju~ne besede: kemijske reakcije, paradoksi, uravnote`enje, zmote
1 INTRODUCTION
In this section we shall discuss the balancing of
chemical reactions from scientific viewpoint. It is an es-
sential precondition for better understanding of our dis-
course about paradoxes connected with traditional ways
of balancing chemical reactions.
Before opening this discussion about balancing
chemical reactions, we would like to give a few remarks
about the name of so–called course »general chemistry«.
Why? The reply is very simple, because this course
treats the balancing of chemical reactions as its subject.
At the beginning of our exposition we want to say that
the name »general chemistry« is not appropriately cho-
sen. Generally speaking, »general chemistry« does not
exist, and on top of all it is not possible to exist, because
the principles of this particular chemistry are weak and
do not hold for all parts of chemistry. They have only
particular meaning and nothing more. In other words, it
means that its principles are not general. This is just one
thing. Another thing, chemistry is founded by virtue of
mathematical principles, but also there not exists »gen-
eral mathematics« and speaking more accurately it is not
possible to exist.
For instance, in mathematical logic in the 20th cen-
tury lots of paradoxes were discovered,1 and mathemati-
cians thought that only there are possible antinomies and
that other parts of mathematics are in safety. Reality
showed that it is not true. In mathematical analysis lots
of counterexamples were found.2,3 Also certain
counterexamples were found in probability & statistics.3,
4 In topology as a contemporary mathematical discipline
a great number of counterexamples were detected too.5 It
does not mean that other mathematical disciplines are
without contradictions. No! Just the opposite, in almost
all the branches of mathematics different kinds of
counterexamples are detected6. These facts show that
mathematical principles are not general, i. e., they hold
only in certain part of mathematics. These are the causes
why there is not »general mathematics«. If we take into
account the fact that chemical reaction is a basic issue in
chemistry and according to its definition (see: Definition
Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522 503
UDK 54 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)503(2011)
2.2, in the section 2), then it follows that chemistry is
founded on mathematical principles. If there is not »gen-
eral mathematics«, then how is possible to exist »general
chemistry«? Simply speaking, it is impossible!
In order to correct this irrational name of chemistry,
it is necessary to choose an appropriate name which will
be more suitable for that particular chemistry. For exam-
ple, the names basic chemistry and elementary chemistry
fit for that chemistry.
In chemical as well as mathematical journals there
are lots of published papers which treat the problem of
balancing chemical reactions. Specially, in chemical
journals are considered many different so–called »chemi-
cal ways« for balancing chemical reactions, but unfortu-
nately all of them offer only particular procedures for
balancing of some simple chemical reactions. These
»chemical ways« very often have negative consequences
for chemistry. For instance, they produce fallacies or ab-
surd results, because most of them work on erroneous
principles, but not on true principles.
Balancing chemical reactions is not a simple proce-
dure as some traditionally oriented chemists think or as
they want to be. In7 the author emphasized very clearly,
that balancing chemical reactions is not chemistry; it is
just linear algebra.
Balancing chemical reactions is a basic matter of
chemistry, if not one of its most important issues, and it
plays a main role in its foundation. Indeed, it is a subtle
question, which deserves considerable attention.
This topic has always excited nature as deeply as no
other question in chemistry. The balancing of chemical
reactions has acted upon human mind in such stimulating
and fruitful way as hardly any other idea, but also this
subject needs an explanation as no other concept.
The concept of balancing chemical reactions in
chemistry is tracked and transformed from traditionalism
to intuitionism in the 19th century, from intuitionism to
irrationalism in the first half of 20th century, from
irrationalism to particularism in the second half of 20th
century, while the particularism is sharpened by
Moore–Penrose matrix8, 9 concept into generalism and
transformed into formalism, which takes on the status of
a paradigm in the 21st century.
To the question of balancing chemical reactions the
mathematicians, chemists and computer engineers with
different conceptions will not always give the same an-
swer. A mathematician, a chemist and a computer engi-
neer will answer each in a different way. Some of them
will admit that perhaps the others are right in a certain
sense and will try to reinterpret the others’ procedures in
their own language. But in general everyone, more or
less, will remain convinced that, in fact, still only he is
right.
Since chemistry is not immune from contradictions,
also we found several absurd results there when balanc-
ing chemical reactions. In this work only just these ab-
surd contradictions are studied.
Sometimes the mistakes in reasoning come because
our experience with one situation causes us to assume
that the same reasoning will hold true in a related but dif-
ferent situation. This kind of mistake can occur at a very
simple level or at a more complex one. At a simple level,
the most common conclusion is that we know that we
have to reject the reasoning, although it may be difficult
to say why. At the more complex level, we may conclude
that the reasoning must be accepted even when the re-
sults seem to contradict our notion of how the real world
works.
Scientific research and experience have shown, when
the results of reasoning and mathematics conflict with
experience, then there is probably a fallacy of some sort
involved10.
In the literature11–13 there are a great number of defi-
nitions for fallacy, but we shall mention merely few of
them. For us is important only deductive fallacy. A de-
ductive fallacy is a deductive reason that is illogical.
In philosophy, the term logical fallacy suitably refers
to a formal fallacy. It is defined as a defect in the struc-
ture of deductive reason, which makes the reason invalid.
Most textbooks echo the standard treatment of fallacy, as
a reason, which seems to be valid but is not so14. Accord-
ing to Maxwell15, a fallacy leads by guile to a wrong but
plausible conclusion.
In mathematics the word fallacy could also refer to a
truthful result obtained by wrong reasoning.
In looking at wrong thinking about easy ideas, we
found some cases in which reasoning about balancing
chemical equations is wrong.
As long as we cannot recognize what the fallacy is,
the situation is a paradox. In some cases, as we shall see,
the paradox is entirely inside chemistry. For most para-
doxes that are inside chemistry, elimination of the falla-
cious reasoning produces a purified chemistry that is a
better description of the real chemistry than the contami-
nated chemistry was.
In philosophy there is a bunch of different definitions
of paradoxes.
According to Sainsbury,16 a paradox is an apparently
unacceptable conclusion derived by apparently accept-
able reasoning from apparently acceptable premises.
Rescher17 defines a paradox as a set of propositions that
are individually plausible but collectively inconsistent.
Chihara18 provides a similar definition: a paradox is an
argument that begins with premises that appear to be
clearly true, that proceeds according to inference rules
that appear to be valid, but ends in contradiction.
Quine19, for example, offers the following definition (us-
ing the term antinomy, instead of a paradox): An
antinomy produces a self–contradiction by accepted
ways of reasoning. It establishes that some tacit and
trusted pattern of reasoning must be made explicit and
hence–forward be avoided or revised. In similar spirit,
Koons20 defines paradox as an inconsistency among
nearly nonrevisable principles that can be resolved only
I. B. RISTESKI: NEW DISCOVERED PARADOXES IN THEORY OF BALANCING CHEMICAL REACTIONS
504 Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522
by recognizing some essential limitation of thought or
language.
In the next section we shall consider the paradoxes
which appeared in different parts of chemistry.
2 PRELIMINARIES
The word paradox in professional works has almost
the same meaning as the word contradiction. Chemistry
as other natural sciences is not immune of paradoxes.
Unlike other natural sciences, in chemistry paradoxes ap-
peared some time later, and it has some of them, while
other sciences are overfull of such contradictions.
Now, we shall mention the well–known paradoxes in
chemistry. They are:
 The Prussian blue paradox:21 The reaction between
ferric and ferrocyanide ions to form Prussian blue and
ferrous and ferricyanide ions to form Turnbull’s blue are
profoundly influenced by the occurrence of the ionic re-
dox equilibrium:
Fe+3 + [Fe(CN)6]
–4 L Fe+2 + [Fe(CN)6]–3,
which is largely displaced toward the right.
 Feigl’s paradoxes:22
• Hydrogen peroxide as a reducing agent (Oxidizing
agents undergo mutual reduction)
2 KMnO4 + 3 H2O2  2 KOH
+ 2 MnO2 + 2 H2O + 3 O2,
NaOCl + H2O2  NaCl + H2O + O2,
2 [Fe(CN)6]
–3 + H2O2 + 2 OH
–
 2 [Fe(CN)6]
–4 + 2 H2O + O2,
Au+3 + 3 H2O2 + 6 OH
–
 6 H2O + 3 O2 + Au.
• Sulfurous acid brings about oxidations
H2SO3 + 2 H2S  3 H2O + 3 S,
or
SO2 + 2 H2S  2 H2O + 3 S,
or
2 Ni(OH)2 + SO2·O2  NiSO4 + Ni(OH)4.
• Nitric acid is not an oxidant
HNO2 + 3 HN3  2 H2O + 5 N2.
• Oxidation of aluminum at room temperature
3 HgCl2 + 2 Al  2 AlCl3 + 3 Hg,
3 HgI2 + 2 Al  2 AlI3 + 3 Hg,
3 HgS + 2 Al  Al2S3 + 3 Hg.
• Nonvolatile oxides of tin and antimony are made to
disappear by heating them with a volatile compound
SnO2 + 4 HI  SnI4 + 2 H2O.
• An acid sets a base free from a salt
H3BO3 + 4 KF  KBF4 + 3 KOH.
• Permanganate is not capable of oxidizing oxalic acid
5 H2C2O4 + 2 KMnO4 + 3 H2SO4
 2 MnSO4 + 10 CO2 + K2SO4 + 8 H2O.
• Ammonium polysulfide brings about an oxidation
SnS + (NH4)2S2  (NH4)2SnS3.
(NH4)2SnS3 + 2HCl  2 NH4Cl
+ H2S + SnS2.
 Levinthal’s paradox:23 The length of time in which
a protein chain finds its folded state is many orders of
magnitude shorter than it would be if it freely searched
all possible configurations.
 Quantum chemistry paradoxes:24
• Preponderant configurations.
• Relevant symmetry.
• Watson effect.
• High ionization energies of the partly filled shells.
• Continua of penultimate ionization.
• Continua of translational energy.
• Questions of time–scale.
• Quantum mechanics pretends to be valid for other
systems than electrons.
• Assembly properties and repeated small entities.
 Campbell’s paradoxes:25
• A catalyst is a substance which increases the rate of a
reaction without entering into it.
• System tends to a minimum in potential energy.
• The entropy of a shuffled deck of cards is greater
than that of a new deck.
• Energy is the ability to do work.
 Paradoxes of spin–pairing energy in gadolinium
(III)26: The spin–pairing parameter D = 9E1/8 for 4f q
separates the averages of all states (S0) and (S0-1) to the
extent 2DS0 where Gd+3 has D = 0.80 eV. Hartree–Fock
(flexible radial functions) have previously been per-
formed for each of the four S, producing a value of D =
1.09 eV but the contributions to D from kinetic energy T,
electron–nuclear attraction Q, and interelectronic repul-
sion C with ratios – 1:6:(– 4) distributed Tc (3.5), Tf (–
4.5), Qc (– 7), Qf (13), Ccc (3.5), Ccf (– 7.9), Cff (0.4) in-
dexed c (closed nl shells) and f (4f). Pragmatic D = 0.80
eV corresponds to 1.84 times the calculated contribution
from Cff and to – 0.184 times the sum of C integrals. Ad-
ditional complications are expected from the correlation
energy –100 eV.
 Structure–Activity Relationship (SAR) paradox27:
Exceptions to the principle that a small change in a mol-
ecule causes a small change in its chemical behavior are
frequently profound.
 Helium paradoxes:28 The relatively high 4He/21Ne,
3He/22Ne and 4He/CO2 ratios in midocean ridge basalts
suggest that it is the midocean ridge basalt reservoir that
is He–rich and that the high ratio 3He/4He in midocean
ridge basalts is due to excess 4He, not a deficit in the
šprimordial’ isotope 3He.
I. B. RISTESKI: NEW DISCOVERED PARADOXES IN THEORY OF BALANCING CHEMICAL REACTIONS
Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522 505
 Temperature dependence of G° and the equilib-
rium constant Keq:29 The sign of S° determines the tem-
perature dependence of G°, it is H° that is responsi-
ble for the shift in Keq with temperature.
 The q/T paradox:30 Which šcontains more heat’, a
cup of coffee at 95 °C or a liter of icewater?
 pH paradox:31 pH  – lg[H+].
 Parrondo’s paradox32: A mathematical concept
known as Parrondo’s paradox is the unexpected situation
in which two specific losing strategies can, by alternat-
ing between them, produce a winning outcome.
 Parrondo’s paradox motivated the development of
many new computational models of chemical systems, in
which thermal cycling problem is studied.33 By these ki-
netics systems compare the rates of formation of prod-
ucts under temperature–cycling and steady–state condi-
tions. Also, these computational models of thermal
cycling announce new applications in chemistry, bio-
chemistry and chemical engineering. More essentially,
by these models one obtains knowledge that some simple
chemical systems might behave paradoxically, and that
forced oscillating conditions may induce an outcome.
However, these paradoxes are not alone and there are
more which appears in the theory of balancing of chemi-
cal equations. Just these paradoxes are main research ob-
ject in this work.
3 A NEW CHEMICAL FORMAL SYSTEM
Chemists must introduce a whole set of auxiliary def-
initions to make the chemistry work consistently. The
more abstract the theory is, the stronger the cognitive
power is.
What does it mean a chemical equation? The reply of
this question lies in the following descriptive definition
given in a compact form.
Definition 2.1. Chemical equation is a numerical
quantification of a chemical reaction.
Let X be a finite set of molecules.
Definition 2.2. A chemical reaction on X is a pair of
formal linear combinations of elements of X , such that
: a x b yij
j
r
j ij
j
s
j
= =
∑ ∑→
1 1
(1  i  m) (2.1)
with aij, bij  0.
The coefficients xj, yj satisfy three basic principles
(corresponding to a closed input–output static model)
• the law of conservation of atoms,
• the law of conservation of mass, and
• the reaction time–independence.
Definition 2.3. Each chemical reaction  has a do-
main
Dom = {x  X  aij > 0} (2.2)
Definition 2.4. Each chemical reaction  has an im-
age
Im = {y  X  bij > 0} (2.3)
Definition 2.5. Chemical reaction  is generated for
some x  X, if both aij > 0 and bij > 0.
Definition 2.6. For the case as the previous defini-
tion, we say x is a generator of .
Definition 2.7. The set of generators of  is thus
Dom 	 Im.
Often chemical reactions are modeled like pairs of
multisets, corresponding to integer stoichiometric con-
stants.
Definition 2.8. A stoichiometrical space is a pair (X ,
R ), where R is a set of chemical reactions on X . It
may be symbolized by an arc–weighted bipartite directed
graph G(X , R) with vertex set X 
 R , arcs x   with
weight aij if aij > 0, and arcs   y with weight bij if bij >
0.
Let us now consider an arbitrary subset A  X .
Definition 2.9. A chemical reaction  may take place
in a reaction combination composed of the molecules in
A if and only if Dom  A .
Definition 2.10. The collection of all feasible reac-
tions in the stoichiometrical space (X , R), that can start
from A is given by
R A = {  R  Dom  A }. (2.4)
In34 is proved the following proposition.
Proposition 2.11. Any chemical equation may be
presented in this form
x yj
j
r
aij
i
i
m
j
j
s
bij
i
i
m
= = = =
∑ ∏ ∑ ∏=
1 1 1 1
Y W (2.5)
where xj (1  j  r) and yj (1  j  s) are unknown ra-
tional coefficients, Yi and Wi (1  i  m) are chemical
elements in reactants and products, respectively, aij (1
 i  m; 1  j  r) and bij (1  i  m; 1  j  s) are
numbers of atoms of elements Yi and Wi, respectively, in
j–th molecule.
Definition 2.12. The nullity of the reaction matrix A
is
nullityA = n – r, (2.6)
where n is the total number of reaction molecules and
by r = rankA the rank of the matrix A is denoted.
Definition 2.13. For any chemical reaction these cri-
teria hold:
1° if nullityA = 0, then the reaction is unfeasible,
2° if nullityA = 1, then the reaction is unique, and
3° if nullityA > 1, then the reaction is non–unique.
We shall define a fallacy in this way.
Definition 2.14. A wrong result attached with a
seemingly logical explanation of why the result is correct
is a fallacy.
A new definition for a paradox should look like this.
Definition 2.15. A paradox is a seemingly true asser-
tion that leads to an inconsistency or a situation, which
resists intuition.
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506 Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522
What does it mean the term non–stoichiometric?
Briefly, it means that a substance may participate in two
different reactions simultaneously, in which case the rel-
ative amounts of two products would bear no fixed ratio
to one another.
According to this, now we can define a non–stoichio-
metric reaction as a real vector space.
Definition 2.16. A non–stoichiometric reaction is a
vector space in which a given set of vector–molecules as
reactants gives final vector–molecules as products whose
molecular proportions are variable in a continuous
sense.
4 NEW PARADOXES AND FALLACIES
In this section we shall present chronologically the
new paradoxes and fallacies, which we discovered in the
theory of balancing chemical reactions.
I. Steinbach gave this statement35: While chemical
equations may balance algebraically, they are not neces-
sarily stoichimetrically exact.
In order to illustrate the above statement, »as the cor-
rect stoichiometric equations«, he »balanced« the fol-
lowing chemical reactions
2 KMnO4 + 5 H2O2 + 4 H2SO4  2 KHSO4
+ 2 MnSO4 + 8 H2O + 5 O2, (3.1)
K2Cr2O3 + 5 H2O2 + 5 H2SO4  2 KHSO4
+ Cr2(SO4)3 + 9 H2O + 4 O2, (3.2)
NaOBr + H2O2  NaBr + H2O + O2, (3.3)
KClO + KClO2  KClO3 + KCl, (3.4)
3 HClO3  HClO4 + Cl2
+ 2 O2 + H2O, (3.5)
4 KClO3 + 16 HCl  4 KCl + 7 Cl2
+ 8 H2O + 2 ClO2. (3.6)
He said: Obviously there are an infinite number of so-
lutions for the coefficients. Only a few of the total possi-
ble solutions are given.
Next, he stated: when equations are balanced by ei-
ther the valence–change or the ion–electron methods, the
coefficients obtained are the correct stoichiometric ones.
He finished his article like this: Equations (3.1) ÷
(3.6) have no stoichiometric meaning, and it is doubtful
whether they have any real significance other than that
they contain the short–hand suggestion of the reactants
used and the products obtained. They do, however, have
a definite suggestion that further investigation would be
most desirable. It is by examples such as these that the
wide divergence between the stoichiometric equation and
the actual mechanism of a chemical reaction is so poi-
gnantly revealed.
All the statements mentioned above are paradoxical.
Why these statements are inconsistent will be explained
in the following text.
Now, we shall make a very clean distinction what is
what! The first statement is completely wrong, because
the algebraic method has not any restriction of its usage.
It holds for every chemical reaction, while other
so–called »chemical methods« hold only for some partic-
ular cases.
From the general solution of the reaction (3.1),
x1 KMnO4 + x2 H2O2 + 2x1 H2SO4
 x1 KHSO4 + x1 MnSO4
+ (3x1/2 + x2) H2O + (5x1/4 + x2/2) O2,
(∀x1, x2  )
we can see that it is a correct two–parametric stoichio-
metric reaction, but not as stated Steinbach. It is just one
thing. Another thing, he »offered« only a particular so-
lution of (3.1) for x1 = 2 and x2 = 5. Immediately, after
publication of his article35, Hall36 pointed out, the state-
ment for the reaction (3.1) that is a »good« example of
variable coefficients is absolutely wrong. It has been
known to be stoichiometric for many, many years and
was studied by C. F. Schönbein, who found that particu-
lar reaction as expressed by (3.1).
Remark 3.1. Schönbein’s particular solution is ob-
tained under certain experimental conditions and it does
not mean that for other different condition other solu-
tions will not be possible!
Chemical reaction (3.2) is balanced incorrectly! The
coefficients of the above reaction correspond to this
chemical reaction
x1 K2Cr2O7 + x2 H2O2 + 5x1 H2SO4
 2x1 KHSO4 + x1 Cr2(SO4)3
+ (4x1 + x2) H2O + (3x1 + x2)/2 O2,
(x1, x2  )
for x1 = 1 and x2 = 5. The reaction that is offered by
Hall36
K2Cr2O7 + 3 H2O2 + 5 H2SO4
 2 KHSO4 + 7 H2O + 3 O2,
is balanced wrongly too!
The general solution of the reaction (3.3) is
x1 NaOBr + x2 H2O2  x1 NaBr
+ x2 H2O + (x1 + x2)/2 O2, (x1, x2  )
and Steinbach found a particular solution for x1 = x2 = 1.
Again, we had not any limitations with the usage of the
algebraically method. Also, in this case the method
worked perfectly.
Chemical reaction (3.4) has this general solution
x1 KClO + x2 KClO2  (x1 + 2x2)/3 KClO3
+ (2x1 + x2)/3 KCl, (x1, x2  )
and the above particular solution corresponds for x1 = x2
= 1.
The general solution of the chemical reaction (3.5) is
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Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522 507
x1 HClO3  x2 HClO4 + (x1 – x2)/2 Cl2
+ (5x1 – 7x2)/4 O2
+ (x1 – x2)/2 H2O, (x1 > 7x2/5) (3.7)
For x1 = 3 and x2 = 1, as a particular case from (3. 7)
immediately follows (3. 5).
For x1 = x2, from (3.7) one obtains this elementary re-
action
2 HClO3 + O2  2 HClO4.
For x1 = 7x2/5, chemical reaction (3.7) transforms into
7 HClO3  5 HClO4 + Cl2 + H2O.
If x1 < x2, then (3.7) becomes
x1 HClO3 + (x2 – x1)/2 Cl2
+ (7x2 – 5x1)/4 O2 + (x2 – x1)/2 H2O
 x2 HClO4, (x1 < x2).
If x2 < x1 < 7x2/5, then chemical reaction (3.7) attains
this form
x1 HClO3 + (7x2 – 5x1)/4 O2  x2 HClO4
+ (x1 – x2)/2 Cl2 + (x1 – x2)/2 H2O,
(x2 < x1 < 7x2/5).
For the reactions (3.3), (3.4) and (3.5) Lehrman37
with the mentioned particular cases gave a comprehen-
sive construction of ad infinitum equations.
The last reaction (3.6), between potassium chlorate
and hydrochloric acid, has a general solution
x1 KClO3 + x2 HCl  x1 KCl
+ (– 3x1 + 5x2/2)/4 Cl2 + x2/2 H2O
+ (3x1 – x2/2)/2 ClO2, (3.8)
(6x1/5 < x2 < 6x1).
For x1 = 4 and x2 = 16 from the chemical reaction
(3.8) as a particular solution follows (3. 6).
For x1 = 5x2/6, from (3.8) one obtains this elementary
reaction
5 KClO3 + 6 HCl  5 KCl
+ 3 H2O + 6 ClO2.
For x1 = x2/6, then (3. 8) becomes
KClO3 + 6 HCl  KCl + 3 Cl2 + 3 H2O.
If x1 > 5x2/6, then (3. 8) transforms into
x1 KClO3 + x2 HCl + (3x1 – 5x2/2)/4 Cl2
 x1 KCl + x2/2 H2O
+ (3x1 – x2/2)/2 ClO2, (x1 > 5x2/6). (3. 9)
We would like to emphasize here that all the reac-
tions are balanced by the well–known algebraically
method, and all the two–parametric coefficients are cor-
rect. It is completely different from the third Steinbach’s
statement given above, which gives advantage to the va-
lence–change or the ion–electron methods.
If we take into account the very well–known rule that
equations for consecutive reactions may be added and
equations for concurrent reactions may not be added,
then obviously Steinbach brings up the old, very old er-
ror involved in adding equations for concurrent reac-
tions. It is his biggest mistake.
What happens when the reactions are not unique ac-
cording to Steinbach? In that case, he »balanced« them
as an infinite number of reactions, which do not express
actual stoichiometric relations. They are »derived« by
combining together the reactions for concurrent reac-
tions, in which coefficients in each of the two or more re-
actions are previously multiplied by different numbers.
Sure, these reactions are incorrect, as they do not corre-
spond to the real stoichiometric relations among the sub-
stances, which are involved. In some cases, there is no
constant relation among the quantities of substances ex-
pressed by his reactions. The ratio of the quantities will
depend upon chemical conditions (concentration, tem-
perature, etc.).
The above Steinbach’s reactions, in fact, are oxida-
tion–reduction reactions for which we found the general
solutions by using of algebraic method. Every one of
them in fact represents two chemical reactions, not one
as Steinbach stated. Therefore, these reactions are not
non–stoichiometric reaction! Standen in his article38
named these Steinbach’s reactions as bizarre non–stoi-
chiometric equations.
McGavock in39 gave a note on errata in the Stein-
bach’s article35.
In order to explain what a non–stoichiometric reac-
tion represents, McGavock considered an example from
his book40.
Example 3.2. It is the cracking reaction
x1 C8H18  x2 C2H4
+ x3 C3H6 + x4 CH4. (3.10)
Now, we shall show that the above reaction (3.10) is
not a non–stoichiometric reaction.
From the scheme given below
v 1
=
C
8H
18
v 2
=
C
2H
4
v 3
=
C
3H
6
v 4
=
C
H
4
C 8 2 3 1
H 18 4 6 4
follows this vector equation
x1v1 = x2v2 + x3v3 + x4v4,
i. e.,
x1 (8, 18)
T
= x2 (2, 4)
T + x3 (3, 6)
T + x4 (1, 4)
T,
or
(8x1, 18x1)
T
= (2x2 + 3x3 + x4, 4x2 + 6x3 + 4x4)
T.
From the system of linear equations
8x1 = 2x2 + 3x3 + x4,
18x1 = 4x2 + 6x3 + 4x4,
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508 Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522
one obtains x4 = x1 and x3 = (7x1 – 2x2)/3. Now, bal-
anced chemical equation has this general solution
x1 C8H18  x2 C2H4 + (7x1 – 2x2)/3 C3H6
+ x1 CH4, (x1 > 2x2/7). (3.11)
The vectors v1, v2, v3 and v4 of the molecules of the
above chemical reaction are linearly dependent and they
generate an infinite number of vector spaces V over ,
i. e., one obtains an infinite number of solutions [x1, x2,
(7x1 – 2x2)/3, x1], (x1 > 2x2/7), that means that the chemi-
cal reaction (3. 10) is non–unique.
Figure 1: Plane 7x1 – 2x2 – 3x3 = 0 in 3
However, it is not possible, either on an algebraic or
empirical basis, to exclude nonintegral values for the co-
efficients. Each point of the plane 7x1 – 2x2 – 3x3 = 0,
given on the Figure 1, represents a triad of positive,
nonintegral values. Infinity of such points corresponding
to infinity of triads of coefficients in the above equation
exists. By virtue of this finding, it is asserted that this
cracking reaction does not represent a non–stoichio-
metric reaction. It contradicts McGavock’s statement.
McGavock in39 obtained this reaction
C8H18  2 C2H4 + C3H6 + CH4. (3.12)
Actually, the reaction (3.12) is a particular solution of
(3.11), for x1 = 1 and x2 = 2. The above McGavock’s
»non–stoichiometric« reaction (3.12) according to
Standen38 can also be written as two reactions:
C8H18  C7H14 + CH4,
C7H14  2 C2H4 + C3H6.
Before the reaction (3.12) be regarded as a genuine
non–stoichiometric reaction, it must be shown that it is
one reaction. It might be a combination of these two re-
actions:
2 C8H18  7 C2H4 + 2 CH4,
3 C8H18  7 C3H6 + 3 CH4.
Also, this Standen’s counterexample shows that
chemical reaction (3.12) is not non–stoichiometric reac-
tion.
By this and the Definition 2.16, we proved that
McGavock’s presentation for non–stoichiometric reac-
tion is an ordinary fallacy.
II. Porges in his article41 wrote: After examinning a
great many chemical equations, one concludes that most
of them are of the type in which the number of com-
pounds involved exceeds the number of elements by
unity.
This Porges’ statement does not represent any crite-
rion for balancing chemical equation and it is completely
wrong. In fact, it is only an ordinary paradox! The
counterexamples given below show it.
Example 3.3. Let us consider this elementary chemi-
cal reaction
x1 BiCl3 + x2 H2O  x3 HCl + x4 BiClO.
Here, we have a case when the number of involved
compounds does not exceed the number of elements by
unity, i. e., in this case four elements are involved in four
molecules. The above chemical reaction reduces to a sys-
tem of four linear equations in four unknown variables
and the chemical reaction has a unique solution. The bal-
anced reaction has this form
BiCl3 + H2O  2HCl + BiClO.
Example 3.4. For instance, in the chemical reaction
x1 Cu2S + x2 HNO3  x3 Cu2SO4
+ x4 NO + x5 H2O,
five elements are involved in five molecules. The above
reaction has a unique solution x1 = 3, x2 = 8, x3 = 3, x4 =
8, x5 = 4.
Example 3.5. In this particular reaction
x1 K4Fe(CN)6 + x2 K2S2O3  x3 KCNS
+ x4 K2SO4 + x5 K2S + x6 FeS,
are involved six elements and same number of mole-
cules, i. e., we have six linear equations in six unknown
variables. The chemical reaction has a unique solution
x1 = 2, x2 = 12, x3 = 12, x4 = 9, x5 = 1, x6 = 2.
Example 3.6. The chemical reaction
x1 AgPF6 + x2 Re(CO)5Br + x3 CH3CN
 x4 AgBr + x5 [Re(CO)5(CH3CN)]PF6,
contains nine elements in five molecules and it reduces
to a system of nine linear equations in five unknown
variables. The chemical reaction has a unique solution
x1 = x2 =  = x5 = 1.
The mentioned four counterexamples contradict to
the above Porges’ statement. By this, we refuted his
statement.
In the same article41 there are fallacies too.
For instance, he reduced the chemical reaction
x1 FeCl2 + x2 K2Cr2O7 + x3 HCl
 x4 FeCl3 + x5 KCl
+ x6 CrCl3 + x7 H2O, (3.13)
to the following system of linear equations
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Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522 509
x1 = x4, 2x1 + x3 = 3x4 + x5 + 3x6,
2x2 = x5, 2x2 = x6, (3.14)
7x2 = x7, x3 = 2x7.
Immediately from (3.14), he obtained
x1 = 6x2, x2 = x2, x3 = 14x2, x4 = 6x2,
x5 = 2x2, x6 = 2x2, x7 = 7x2. (3.15)
Next, he said: It is evident that the general integral
solution of (3.14) is derived from x2 = k, where k is any
positive integer, and consequently although (3.14) has an
infinite number of integral solutions, all of them derive
from x2 = k, and consist merely of multiplies of values for
the respective variables established by (3.15). All solu-
tions other than for k = 1, are therefore trivial chemically
as well as mathematically.
Unfortunately, the above statement is fallacious! The
last sentence All solutions other than for k = 1, are there-
fore trivial chemically as well as mathematically, is in-
correct. Previous Porges stated that x2 = k, where k is any
positive integer, and after that he took k = 1. It is wrong!
Why? To this question a very simple answer will follow
like this. The reaction (3.13) has a unique solution. If we
substitute (3.15) into (3.13), and after that if we divide
the relation (3.13) by an arbitrary real number x2 ≠ 0, im-
mediately follows
6 FeCl2 + K2Cr2O7 + 14 HCl
 6 FeCl3 + 2 KCl + 2 CrCl3 + 7 H2O.
Actually, Porges considered the general solution
(3.15) of the system (3.14) separately of (3.13), what is
wrong. The reaction (3.13) and the general solution
(3.15) of the system (3.14) must be considered as one
whole, because they are connected with each other.
Obviously, we did not introduce any constant k, as it
was done previously by Porges. Our approach is com-
pletely different than Porges’ wrong way he used.
The same remark also holds for the second Porges’
reaction
As2S3 + 3 (NH4)2S  2 (NH4)3AsS3. (3.16)
considered in41.
The third considered reaction in41 is given by this ex-
pression
x1 HAuCl3 + x2 K4Fe(CN)6
 x3 KAu(CN)4 + x4 KAu(CN)2
+ x5 KAu(CN)2Cl2 + x6 KCl + x7 HCl
+ x8 [4Fe(CN)3·3Fe(CN)2]. (3.17)
For its »solution« the author offered these expres-
sions
x1 = 12k – 2, x2 = 7k, x3 = 2, x4 = 8k – 1,
x5 = 4k – 3, x6 = 16k + 2,
x7 = 12k – 2, x8 = k, (3.18)
where k takes on all positive integral values. Yet (3.18)
is merely a particular solution, but it is not the general
solution of (3.17).
The general solution of the reaction (3.17) is given by
this expression
14x1 HAuCl3 + 14x2 K4Fe(CN)6
 (24x2 – 14x1) KAu(CN)4
+ (7x1 + 4x2) KAu(CN)2
+ (21x1 – 28x2) KAu(CN)2Cl2
+ (56x2 – 14x1) KCl + 14x1 HCl
+ 2x2 [4Fe(CN)3·3Fe(CN)2],
(4x2/3 < x1 < 12x2/7). (3.19)
The balanced reaction (3.19) is an expression of two
parameters, but it is not one–parametric expression as it
is given in41. In fact, it is another Porges’ fallacy.
Remark 3.7. Intentionally, we omitted from consider-
ation other particular solutions of (3.19) because we had
into account that our work has a limited size.
Before ending his article41, Porges posed the follow-
ing three questions:
1° Is there, for equations of this third type which ad-
mits infinitely many distinct, not multiple solutions, a
least action principle similar to that in mechanics?
2° Of the unlimited number of ways of balancing
such an equation as (3.17), is that corresponding to the
solution in least integer of the linear equations the one
invariably indicated by the laboratory work, which is, af-
ter all, the real criterion?
3° Further, what would a minimum solution be – that
for which the square root of the arithmetic mean of the
squares of the variables is less than for any other solu-
tion?
To date we did not meet in chemical literature any re-
ply on these questions. It is a challenge for us to try to
give appropriate answers on the above questions. Sure,
the answers will be given in a rough form, because a
comprehensive replay looks for a special article dedi-
cated only on that particular subject.
The answer on the first question should be like this.
No! In chemistry there is not a least action principle sim-
ilar to that in mechanics, because the balancing chemi-
cal equation has not any tangent point with mechanics,
just it is connected with linear algebra. In fact, balanc-
ing chemical equations is not chemistry; it is just linear
algebra7. On the other hand, the term a least action as-
sociate on the dynamism of 18th century as metaphysics
was traced by Leibniz and Boscovich. Later Boyle gave
an explicit formulation of chemistry in a coherent meta-
physical scheme. A comprehensive study about meta-
physics of chemical reaction is given in42.
The second question is interesting for a discussion,
and it can be answered negatively in this way. Unlimited
number of ways of balancing such a reaction as (3.17) is
not corresponding to the solution in least integer of the
linear equations the one invariably indicated by the lab-
oratory work, which is, after all, the real criterion. This
case, in fact, boils down to the continuum problem. It is
an extremely hard problem, which is a stumbling block
I. B. RISTESKI: NEW DISCOVERED PARADOXES IN THEORY OF BALANCING CHEMICAL REACTIONS
510 Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522
for mathematicians as well as chemists. In other words,
this case can sink into the third question.
As the third question is posed, it contains a wrong
formulation for minimum solution. It is one more
Porges’ fallacy. One can reply the last question like this.
When the question is posed it was unbelievable for that
time. It was impossible because the mathematical meth-
ods needed for its proof were unknown. For instance,
Moore–Penrose pseudoinverse matrix8,9 was discovered
ten years later in 1955 and its first application in
chemistry43 appeared more than two decades later in
1978. Also, then was unknown the general problem of
balancing chemical equations44. From today viewpoint,
by using of Moore–Penrose pseudoinverse matrix the
minimal solution is obtained in45,46.
Now, a new question arises: why not look for a topol-
ogy of solutions of chemical equations, instead of finding
minimal solutions? It is a much better question, than
non–unique equation to be reduced to a minimal case.
The question is new, and it is extremely hard. Sure, it
will be a challenge for the next research.
Porges finished his article41 on this way: In the very
few equations of the third type encountered by the writer,
the balancing in least integers was always that for which
k was also least, and was identical with the laboratory
balancing; but given a general solution not minimized by
smallest admissible value of k – theoretically not impos-
sible – then what?
It represents one last Porges’ fallacy! We shall build
the reply to this question on the concept of the contin-
uum. In fact, the word continuum is recognizable as the
name used by Cantor to refer to the real line. From the
expression (3. 19), we can see that this kind of chemical
equations reduces to the Cantor’s continuum problem.
This problem is simply condensed in the following ques-
tion: How many points are there on the straight line in
Euclidean space? In other words, the question is: How
many different sets of integers do there exist?47 This
problem is neither simple nor easy; it needs a wide ex-
planation. It shows that balancing chemical equations is
a main object in Foundation of Chemistry, which lies
in an intertwined mixture of topology, abstract algebra,
linear algebra, axiomatic set theory, mathematical logic,
computability theory and proof theory. To explain a little
more fully this idea we must first discuss the concept of
a formal system. Why? Simply, chemical equation must
be treated only as a formal system, if we like to avoid ap-
pearance of paradoxes. In an opposite case we shall have
paradoxes as these mentioned in this section.
A formal chemical system consists of a finite set of
symbols and of a finite number of rules by which these
symbols can be combined into formulas or statements.
That kind of formal system is given in second section of
this work. A number of such statements are nominated as
axioms and by repeated applications of the rules of the
system one obtains an ever growing body of provable
statements.
A proof of a given statement is a finite sequence of
statements that starts with an axiom and ends with the
preferred statement. The sequence is such that every
transitional statement is either an axiom or is derivable
by the rules of the system from statements that lead it.
Thus, a statement that a sequence of formulas does or
does not represent a proof of formula is »not« a state-
ment in the formal system itself. It is a statement
»about« the system and such statements are often re-
ferred to as »metamathematical«.
For the first time a formal generalized inverse matrix
approach for balancing chemical equation is introduced
in44. Balancing chemical equations as a matrix well–de-
fined formal system is given in the works45,46,48–50.
III. Another paradox in the theory of balancing
chemical equations is the following Standen’s state-
ment38: It would seem that examples could not be found
where the number of mathematical equations actually
exceeds the number of variables; for if the mathematical
equations were inconsistent, the whole thing would be an
impossibility, while if they were consistent it would indi-
cate that the chemical equation had been appropriately
broken down into its terms.
To prove the above Standen’s absurdity, we shall use
the following counterexamples.
Example 3.8. The chemical reaction
x1 CoCl2 + x2 Na3PO4  x3 NaCl + x4 Co3(PO4)2,
contains five elements involved in four molecules, i. e.,
in this case the chemical reaction reduces to a system of
five linear equations in four unknown variables and the
chemical reaction has a unique solution. The balanced
reaction has this form
3 CoCl2 + 2 Na3PO4  6 NaCl + Co3(PO4)2.
Example 3.9. In this particular reaction
x1 Cu(NH3)4Cl2 + x2 KCN + x3 H2O
 x4 [K2Cu(CN)3·NH4Cl·KCl] + x5 KCNO + x6 NH3,
are involved seven elements in six molecules, i. e., we
have seven linear equations in six unknown variables.
The chemical reaction has a unique solution x1 = 2, x2 =
7, x3 = 1, x4 = 2, x5 = 1, x6 = 6.
Example 3.10. For instance, the chemical reaction
x1 AgPF6 + x2 Re(CO)5Br + x3 CH3CN
+ x4 K2S2O3  x5 KBr
+ x6 [Re(CO)5(CH3CN)]PF6
+ x7 Ag2S + x8 K2O + x9 SO2,
has eleven elements involved in nine molecules. The
above chemical reaction reduces to a system of eleven
linear equations in nine unknown variables, whose
unique solution is x1 = 4, x2 = 4, x3 = 4, x4 = 3, x5 = 4, x6
= 4, x7 = 2, x8 = 1, x9 = 4.
By the last three counterexamples we showed that the
above Standen’s38 statement is an absurd.
IV. Next, we shall consider the Blakley’s51 paradox.
Among other chemical reactions, he considered hydroly-
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Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522 511
sis of two organic substances: bradykinin and grami-
cidin–S.
Example 3. 11. The following reaction
x1 C2H5NO2 + x2 C3H7NO3
+ x3 C6H14N4O2 + x4 C5H9NO2
+ x5 C9H11NO2
 x6 H2O + x7 C50H73N15O11. (3.20)
was studied in51. Blakley considered balancing of the
above chemical reaction by a matrix approach using a
module basis.
He »proved« that »hydrolysis (3.20) of bradykinin is
unique«. It represents only an empirical discovered rela-
tionship, which is wrong.
We shall prove the absurdity of his statement by us-
ing the well–known algebraic method for balancing
chemical equations.
By the way, we shall show that this method is power-
ful and its usage is not limited as some traditional ori-
ented chemists think.
Let us consider the scheme of the chemical reaction
(3.20).
C
2H
5N
O
2
C
3H
7N
O
3
C
6H
14
N
4O
2
C
5H
9N
O
2
C
9H
11
N
O
2
H
2O
C
50
H
73
N
15
O
11
C 2 3 6 5 9 0 –50
H 5 7 14 9 11 –2 –73
N 1 1 4 1 1 0 –15
O 2 3 2 2 2 –1 –11
From the above scheme imediatelly follows the
stoichiometric matrix
A =
−
− −
−
− −
⎡
⎣
⎢
⎢
⎢
⎤
⎦
⎥
2 3 6 5 9 0 50
5 7 14 9 11 2 73
1 1 4 1 1 0 15
2 3 2 2 2 1 11
⎥
⎥
with r = rankA = 4.
Since nullityA = n – r = 7 – 4 = 3 > 1, where n is the
total number of reaction molecules, then the chemical re-
action is possible and it has an infinite number of solu-
tions. Let us prove it.
One can reduce the chemical reaction (3.20) to the
following system of linear equations
2 x1 + 3 x2 + 6 x3 + 5 x4 + 9 x5 = 50 x7,
5 x1 + 7 x2 + 14 x3 + 9 x4 + 11 x5
= 2 x6 + 73 x7, (3.21)
x1 + x2 + 4 x3 + x4 + x5 = 15 x7,
2 x1 + 3 x2 + 2 x3 + 2 x4 + 2 x5
= x6 + 11 x7.
The general solution of the system (3.21) is given by
the following expressions
x4 = 66 x1/23 + 93 x2/23 – 45 x3/23,
x5 = – 14 x1/23 – 26 x2/23 + 43 x3/23, (3.22)
x6 = 95 x1/23 + 137 x2/23 – 24 x3/23,
x7 = 5 x1/23 + 6 x2/23 + 6 x3/23,
where xi > 0 (1  i  3) are arbitrary real numbers.
After substitution of (3.22) in (3.20), one obtains a
balanced chemical reaction
x1 C2H5NO2 + x2 C3H7NO3 + x3 C6H14N4O2
+ (66 x1/23 + 93 x2/23 – 45 x3/23) C5H9NO2
+ (– 14 x1/23 – 26 x2/23 + 43 x3/23) C9H11NO2
 (95 x1/23 + 137 x2/23 – 24 x3/23) H2O
+ (5 x1/23 + 6 x2/23
+ 6 x3/23) C50H73N15O11. (3.23)
From (3.23) follows this system of inequalities
22x1 + 31x2 – 15x3 > 0,
– 14x1 – 26x2 + 43x3 > 0, (3.24)
95x1 + 137x2 – 24x3 > 0,
5x1 + 6x2 + 6x3 > 0.
From the first and the second inequality of (3.24) im-
mediately follows this expression
14 x1/43 + 26 x2/43 < x3
< 22 x1/15 + 31 x2/15. (3.25)
The expression (3.25), the third and the fourth in-
equality of (3.24) are necessary and sufficient conditions
for (3.23) to hold.
For instance, if we substitute x1 = x2 = 1 and x3 = 2 in
(3.23), then as a particular case appears this reaction
C2H5NO2 + C3H7NO3 + 2 C6H14N4O2
+ 3 C5H9NO2 + 2 C9H11NO2
 8 H2O + C50H73N15O11,
for which Blakley51 stated that it is unique, but it is not
true.
Now, we shall give two more particular cases for
which (3.23) holds.
Let x1 = x2 = x3 = 1, then from (3.23) one obtains the
reaction
23 C2H5NO2 + 23 C3H7NO3
+ 23 C6H14N4O2 + 114 C5H9NO2
+ 3 C9H11NO2
 208 H2O + 17 C50H73N15O11.
Let x1 = x2 = 1 and x3 = 3, then (3.23) becomes
23 C2H5NO2 + 23 C3H7NO3
+ 69 C6H14N4O2 + 24 C5H9NO2
+ 89 C9H11NO2
 160 H2O + 29 C50H73N15O11.
The other particular cases of (3.23) are not consid-
ered, because we took into account the Remark 3.7.
Example 3.12. In51 the hydrolysis of gramicidin–S
was studied, given by the following reaction
x1 C5H9NO2 + x2 C5H11NO2 + x3 C6H13NO2
+ x4 C9H11NO2 + x5 C5H12N2O2
 x6 C60H92N12O10 + x7 H2O. (3.26)
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512 Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522
For the chemical reaction (3.26) if we write its
stoichiometric scheme, then one obtains
C
5H
9N
O
2
C
5H
11
N
O
2
C
6H
13
N
O
2
C
9H
11
N
O
2
C
5H
12
N
2O
2
C
60
H
92
N
12
O
10
H
2O
C 5 5 6 9 5 –60 0
H 9 11 13 11 12 –92 –2
N 1 1 1 1 2 –12 0
O 2 2 2 2 2 –10 –1
from where follows the stoichiometric matrix
A =
−
− −
−
− −
⎡
⎣
⎢
⎢
⎢
⎤5 5 6 9 5 60 0
9 11 13 11 12 92 2
1 1 1 1 2 12 0
2 2 2 2 2 10 1⎦
⎥
⎥
⎥
with r = rankA = 4.
Since nullityA = n – r = 7 – 4 = 3 > 1, then chemical
reaction is possible and it has an infinity number of solu-
tions. Let us prove it.
The system of linear equations obtained from (3. 26)
is
5 x1 + 5 x2 + 6 x3 + 9 x4 + 5 x5 – 60 x6 = 0,
9 x1 + 11 x2 + 13 x3 + 11 x4
+ 12 x5 – 92 x6 – 2 x7 = 0, (3.27)
x1 + x2 + x3 + x4 + 2 x5 – 12 x6 = 0,
2 x1 + 2 x2 + 2 x3 + 2 x4
+ 2 x5 – 10 x6 – 1 x7 = 0.
The general solution of (3.27) is given by the expres-
sions
x1 = – 4 x4 + 20 x6 – x7,
x2 = 7 x4 – 57 x6 + 9 x7/2, (3.28)
x3 = – 4 x4 + 35 x6 – 5 x7/2,
x5 = 7 x6 – x7/2,
where x4, x6 and x7 are arbitrary real numbers.
After substitution of (3.28) in (3.26), the balanced
chemical reaction has this form
(– 4 x4 + 20 x6 – x7) C5H9NO2
+ (7 x4 – 57 x6 + 9 x7/2) C5H11NO2
+ (– 4 x4 + 35 x6 – 5 x7/2) C6H13NO2
+ x4 C9H11NO2 + (7 x6 – x7/2) C5H12N2O2
 x6 C60H92N12O10 + x7 H2O, (3.29)
From (3.29) follows this system of inequalities
– 4 x4 + 20 x6 – x7 > 0,
7 x4 – 57 x6 + 9 x7/2 > 0, (3.30)
– 4 x4 + 35 x6 – 5 x7/2 > 0,
7 x6 – x7/2 > 0.
From the second and the third inequality of (3.30)
immediately follows this expression
– 14 x4/9 + 114 x6/9 < x7
< – 8 x4/5 + 14 x6. (3.31)
The expression (3.31), the first and the fourth in-
equality of (3.30) are necessary and sufficient conditions
for (3.29) to hold.
For instance, if we substitute x4 = 2, x6 = 1 and x7 =
10 in (3.29), then as a particular case appears this reac-
tion
2 C5H9NO2 + 2 C5H11NO2 + 2 C6H13NO2
+ 2 C9H11NO2 + 2 C5H12N2O2
 C60H92N12O10 + 10 H2O.
for which Blakley51 stated that it is unique, but it is not
true.
Now, we shall give more two particular cases for
which holds (3.29).
Let x4 = x6 = 9 and x7 = 102, then from (3.29) one ob-
tains the reaction
42 C5H9NO2 + 9 C5H11NO2
+ 24 C6H13NO2 + 9 C9H11NO2 + 12 C5H12N2O2
 9 C60H92N12O10 + 102 H2O.
Let x4 = 27, x6 = 21 and x7 = 250, then (3.29) be-
comes
62 C5H9NO2 + 117 C5H11NO2
+ 2 C6H13NO2 + 27 C9H11NO2 + 22 C5H12N2O2
 21 C60H92N12O10 + 250 H2O.
The other particular cases of (3.29) are not consid-
ered because we took into account the Remark 3.7.
V. Das in his article52 applied the partial equation
method for balancing chemical equations. There he
wrote: if the number of reactants and products is equal
to or less (or at most one more) than the total number of
elements involved in the chemical equation, then there
will be only one way of balancing a chemical equation.
For the first two cases this statement is a paradox!
The next two counterexamples given below contradict to
the above statement.
Example 3.13. The following chemical reaction
x1 NO2 + x2 HClO  x3 HNO3 + x4 HCl,
has involved four elements in four molecules, but it has
not a unique solution as stated above. It is an unfeasible
reaction, because x1 = x2 = x3 = x4 = 0.
Now, we shall give another counterexample.
Example 3.14. In the reaction
x1 KIO2 + x2 Pb(NO3)2  x3 KNO3 + x4 PbI2,
five elements are involved in four molecules. This reac-
tion has only a trivial solution x1 = x2 = x3 = x4 = 0. It
shows that this reaction is unfeasible.
VI. Next, we shall elaborate another very interesting
paradox. García53 gave a »half–reaction method« for bal-
ancing chemical equations. He described his »method«
on this way. The chemical reaction is divided into two
half–reactions and each one is balanced independently.
These two balanced half–reactions are added together to
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Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522 513
get the correct stoichiometry of the reaction. One
half–reaction is formed with compounds that contain the
same elements other than oxygen and hydrogen. The re-
maining compounds, and others if it is necessary, consti-
tute the other half–reaction.
Example 3.15. This chemical reaction
x1 FeS2 + x2 HNO3  x3 Fe2(SO4)3
+ x4 NO + x5 H2SO4, (3.32)
he »balanced« like this
2 FeS2 + 10 HNO3  Fe2(SO4)3
+ 10 NO + H2SO4 + 4 H2O. (3.33)
Unfortunately reaction (3.33) is quantitatively and
qualitatively different from the reaction (3.32). Actually,
the reaction (3.33) is augmented reaction (3.32) by four
water molecules. Reactions (3.32) and (3.33) belong to
different types of reactions, and according to it, they are
incompatible. The reaction (3.32) belongs to the type of
unfeasible reactions, because its vectors of molecules do
not generate a vector space V over 
.
For our next analysis we shall use the newest method7
for balancing chemical equations founded by virtue of
theory of complex finite dimensional vector spaces. We
chose it, because in this particular Garcia’s case, it was
the most suitable method for comparative analysis of
chemical reactions which belong to different classes. Ap-
plication of this method confirmed its scientific suprem-
acy.
From (3.32) one obtains the scheme given below
v 1
=
F
eS
2
v 2
=
H
N
O
3
v 3
=
F
e 2
(S
O
4)
3
v 4
=
N
O
v 5
=
H
2S
O
4
Fe 1 0 2 0 0
S 2 0 3 0 1
H 0 1 0 0 2
N 0 1 0 1 0
O 0 3 12 1 4
The vector equation of reaction (3.32) is
x1v1 + x2v2 = x3v3 + x4v4 + x5v5,
i. e.,
x1 (1, 2, 0, 0, 0)
T + x2 (0, 0, 1, 1, 3)
T
= x3 (2, 3, 0, 0, 12)
T + x4 (0, 0, 0, 1, 1)
T
+ x5 (0, 1, 2, 0, 4)
T,
or
(x1, 2x1, x2, x2, 3x2)
T
= (2x3, 3x3 + x5, 2x5, x4, 12x3 + x4 + 4x5)
T.
The system of linear equations
x1 = 2x3, 2x1 = 3x3 + x5, x2 = 2x5, x2 = x4,
3x2 = 12x3 + x4 + 4x5,
is inconsistent, because one obtains the contradiction x5
= x1/2 and x5 = – x1. It means that the vectors v1, v2, v3,
v4 and v5 of molecules of the chemical reaction (3.32)
are linearly independent and they do not generate a vec-
tor space V over 
. By this we showed that the chemical
reaction (3.32) is unfeasible.
On the other hand, the rank of the reaction matrix A
of the chemical reaction (3.32) is
r = rankA = rank
1 0 2 0 0
2 0 3 0 1
0 1 0 0 2
0 1 0 1 0
0 3 12 1 4
⎡
⎣
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
= 5
According to the algebraic criterion (2.6) for balanc-
ing chemical reactions, the reaction (3.32), has nullityA
= n – r = 5 – 5 = 0, that means that the reaction (3.32) is
unfeasible. Both proofs, vector and algebraic, confirmed
the same, that the reaction (3.32) is unfeasible. It contra-
dicts the García’s »procedure« named as »half–reaction
method« for balancing chemical reactions.
Now, we shall consider the reaction (3.33) in its un-
balanced form
x1 FeS2 + x2 HNO3  x3 Fe2(SO4)3
+ x4 NO + x5 H2SO4 + x6 H2O (3.34)
Now, we need the stoichiometric scheme for the
above chemical reaction. From this particular reaction
(3.34), we shall derive very easy required stoichiometric
scheme
v 1
=
F
eS
2
v 2
=
H
N
O
3
v 3
=
F
e 2
(S
O
4)
3
v 4
=
N
O
v 5
=
H
2S
O
4
v 6
=
H
2O
Fe 1 0 2 0 0 0
S 2 0 3 0 1 0
H 0 1 0 0 2 2
N 0 1 0 1 0 0
O 0 3 12 1 4 1
From the above scheme one obtains this vector equa-
tion
x1v1 + x2v2 = x3v3 + x4v4 + x5v5 + x6v6,
i. e.,
x1 (1, 2, 0, 0, 0)
T + x2 (0, 0, 1, 1, 3)
T
= x3 (2, 3, 0, 0, 12)
T + x4 (0, 0, 0, 1, 1)
T
+ x5 (0, 1, 2, 0, 4)
T + x6 (0, 0, 2, 0, 1)
T,
or
(x1, 2x1, x2, x2, 3x2)
T
= (2x3, 3x3 + x5, 2x5 + 2x6, x4, 12x3
+ x4 + 4x5 + x6)
T.
The solution of the system of linear equations
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514 Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522
x1 = 2x3, 2x1 = 3x3 + x5, x2 = 2x5 + 2x6,
x2 = x4, 3x2 = 12x3 + x4 + 4x5 + x6,
is
x2 = 5x1, x3 = x1/2, x4 = 5x1,
x5 = x1/2 and x6 = 2x1. (3.35)
If we substitute (3.35) in (3.34), and after that, if we
divide the reaction by x1/2 one obtains (3.33). The vec-
tors v1, v2, v3, v4, v5 and v6 of reaction molecules are lin-
early dependent and they generate a vector space V over
. By this we confirmed that the reaction (3.33) has a
unique solution.
Similarly as in the previous equation analysis, now
we can go to the next step. Now we can calculate the
rank of the reaction matrix A of the reaction (3.33).
Therefore, we can express its rank in this way
r = rankA = rank
1 0 2 0 0 0
2 0 3 0 1 0
0 1 0 0 2 2
0 1 0 1 0 0
0 3 12 1 4 1
⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
⎥
= 5
According to the algebraic criterion (2.6) for balanc-
ing chemical equations, the reaction matrix A has
nullityA = n – r = 6 – 5 = 1. By this, again we showed
that chemical reaction (3.33) has a unique solution. The
analysis of the reaction (3.33), established that it belongs
to the type of solvable equations, which have a unique
solution.
This comparative analysis confirmed that García’s
half–reaction simple »method«53 is completely wrong.
Generally speaking, his so–called »method« cannot rec-
ognize the type of reaction, and much less to decide if
the chemical equation is solvable or not. To support it,
we shall give a dozen of counterexamples, where water
molecules in García’s procedure of reaction extension
(3.33), may be substituted by other molecules of the ele-
ments involved in the reaction (3.32), as it is exposed by
the following reactions
2 FeS2 + 8 HNO3  Fe2(SO4)3
+ 8 NO + H2SO4 + 3 H2,
2 FeS2 + 14 HNO3  Fe2(SO4)3
+ 14 NO + H2SO4 + 6 H2O2,
10 FeS2 + 34 HNO3  5 Fe2(SO4)3
+ 22 NO + 5 H2SO4 + 12 NH2,
14 FeS2 + 50 HNO3  7 Fe2(SO4)3
+ 38 NO + 7 H2SO4 + 12 NH3,
10 FeS2 + 34 HNO3  5 Fe2(SO4)3
+ 22 NO + 5 H2SO4 + 6 N2H4,
10 FeS2 + 34 HNO3  5 Fe2(SO4)3
+ 26 NO + H2SO4 + 4 (NH4)2SO3,
14 FeS2 + 50 HNO3  7 Fe2(SO4)3
+ 38 NO + H2SO4 + 6 (NH4)2SO4,
4 FeS2 + 16 HNO3  2 Fe2(SO4)3
+ 13 NO + 2 H2SO4 + 3 NH4O,
34 FeS2 + 130 HNO3  17 Fe2(SO4)3
+ 106 NO + 17 H2SO4 + 12 (NH4)2O.
Neither one of the above reactions nor chemical reac-
tion (3.33) is equivalent to the chemical reaction (3.32).
Therefore, reactions (3.32) and (3.33) are incompatible.
In a similar way, García considered the following two
reactions53.
Example 3. 16. The following chemical reaction
x1 CrI3 + x2 KOH + x3 Cl2  x4 K2CrO4
+ x5 KIO3 + x6 KCl, (3.36)
he »balanced« like this
2 CrI3 + 52 KOH + 21 Cl2  2 K2CrO4
+ 6 KIO3 + 42 KCl + 26 H2O. (3.37)
On top of all he said: This method is appropriate to
balance any kind of reaction, even those that include
complex ions or reactions of compounds with oxidation
numbers difficult to determine.
By the same analysis which we used in the previous
counterexample, very easy we shall show the absurdity
of his statement.
Reaction (3.36) belongs to the type of unfeasible re-
action. Its stoichiometric matrix A has rank r = rankA =
6 and its nullityA = n – r = 6 – 6 = 0 verifies that it is an
unfeasible reaction. Also, this reaction generates an in-
consistent system of linear equations which has only a
trivial solution xi = 0, (1  i  6). Thus, this algebraic
criterion verifies that the reaction (3.36) is unfeasible
too.
Reaction (3. 37) generates a consistent system of lin-
ear equations which has a unique solution given in
(3.37). Also, the nullityA = n – r = 7 – 6 = 1, shows that
this equation has a unique solution. Therefore, reactions
(3.36) and (3.37) are incompatible, because they are two
completely different types of reactions – the first one is
an unfeasible reaction, while the second one is a unique
reaction.
Now, we shall mention just two chemical reactions,
where water molecules in García’s »procedure« of reac-
tion extension (3.37), are substituted by other molecules
of the elements involved in the reaction (3.36):
2 CrI3 + 26 KOH + 8 Cl2  2 K2CrO4
+ 6 KIO3 + 16 KCl + 13 H2,
2 CrI3 + 26 KOH + 21 Cl2  2 K2CrO4
+ 6 KIO3 + 16 KCl + 26 HCl.
Neither one of the above reactions nor chemical reac-
tion (3.37) is equivalent to the chemical reaction (3.36).
Therefore, reactions (3.36) and (3.37) are incompatible.
In the same paper, García considered two ionic reac-
tions too. Unfortunately, also in these particular cases the
same absurdity appears again.
Example 3.17. For example, he »balanced« an un-
feasible reaction
x1 IO3
– + x2 Br
–  x3 IO2
– + x4 Br2, (3.38)
where xi = 0 (1  i  4) as a unique reaction
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Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522 515
IO3
– + 2 Br– + 2 H+  IO2
–
+ Br2 + H2O. (3.39)
Reactions (3.38) and (3.39) are two completely dif-
ferent reactions and they are incomparable!
A second ionic example that García considered is
given in the next example.
Example 3.18. The reaction
x1 ClO
– + x2 P4  x3 H2PO4
– + x4 Cl
–, (3.40)
is »balanced« like this
10 ClO– + P4 + 2 H2O + 4 OH
–
 4 H2PO4
– + 10 Cl–. (3.41)
Where is the hydrogen atom in the left side of (3.40)?
Reaction (3.40) is an absurd, because it does not contain
hydrogen atom in its left side. With this kind of reac-
tions current chemistry does not work!
Reaction (3.41) does not dependent on (3.40) and it
represents a solvable equation. In other words, reactions
(3.40) and (3.41) are completely different types of reac-
tions and they are incompatible.
VII. In54 the so–called »formal balance numbers«
(FBN) are introduced like this: Formal balance numbers
are an aid that may grossly facilitate the problem of bal-
ancing complex redox equations. They may be chosen as
being equal to the traditional values of oxidation num-
bers, but not necessarily. An inspection of the redox
equation may suggest the optimal values that are to be
assigned to formal balance numbers. In most cases,
these optimal values ensure that only two elements will
šchange their state’ (i. e. the values of the formal balance
numbers), allowing the use of the oxidation number tech-
nique for balancing equations, in its simplest form. Just
as for oxidation numbers, the algebraic sum of the for-
mal balance numbers in a molecule/neutral unit is 0,
while in an ion it is equal to its charge (the sum rule).
It was quickly detected that the »procedure« given
in54 boils down to using of well–known unconventional
oxidation numbers, which previously were advocated by
Tóth55 and Ludwig56.
Consider this sentence from previous definition: They
may be chosen as being equal to the traditional values of
oxidation numbers, but not necessarily. It is a paradox!
If the »formal balance numbers« can be the same as oxi-
dation numbers or not, then the whole definition is illog-
ical. This definition represents only a contradictory
premise, which does not have any correlation with bal-
ancing chemical equations. It is just one thing.
Another thing, the above definition does not speak
anything about balancing chemical reactions in a chemi-
cal sense of the word, or their solution in a mathematical
sense.
Recent research7 confirmed that a chemical equation
can be balanced if and only if it generates a vector
space. That is a necessary and sufficient condition for
balancing a chemical equation!
The so–called called »formal balance numbers«,
which actually are the same as the well–known oxidation
numbers, do not represent any criterion for balancing
chemical equations.
Also the author of54 asserted that his »procedure« is
probably the fastest of all possible methods! Obviously
the author omitted to prove it. Perhaps, his statement is
valid (if he can prove it?) in some metachemistry, but
from a viewpoint of current mathematics and chemis-
try it is not true. Why? The reason is very simple. In
mathematics, as well as in chemistry there is neither a
definition for speed of equation solution nor its unit, and
according to it, it is impossible to compare which
method is faster. It is just one thing; another thing is that
the definition of so–called »formal balance numbers«
(FBN)54 is paradoxical and it produces only inconsistent
procedure for balancing chemical equations. According
to it, the author’s assertion of54 is an absurd.
The »procedure«54 founded by virtue of so–called
»formal balance numbers« (FBN), with several
counterexamples was refuted in7.
VIII. Ten Hoor in57 obtained this result:
C + x O2  2(1 – x) CO + (2x – 1) CO2, (3.42)
(1/2  x  1)
if the coefficient of a product is allowed to be equal to
zero. Taking x equal to its smallest or largest extreme
value, equation (3.42) reduces to
C + 1/2 O2  CO, (3.43)
or
C + O2  CO2, (3.44)
respectively.
The above statement is wrong. The reaction (3.42)
holds if only if 1/2 < x < 1, like it is shown on Fig. 2, but
not as it is given in57.
Figure 2: The interval 1/2 < x < 1
In (3.42) x does not have any extreme value, because
it is presented by the following linear functions: x, 2 – 2x
and 2x – 1. None of these functions have extrema, since
their second derivatives are equal to zero. Then on what
basis ten Hoor57 states that for smallest or largest ex-
treme value of x the reaction (3.42) reduces to (3.43) and
(3.44) respectively?
The chemical reaction (3.42) has two subgenerators 2
– 2x and 2x – 1 which generate the following particular
cases:
1° For x = 1, then (3.42) reduces to
C + O2  CO2.
I. B. RISTESKI: NEW DISCOVERED PARADOXES IN THEORY OF BALANCING CHEMICAL REACTIONS
516 Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522
2° For x = 1/2, then (3.42) becomes
C + 1/2 O2  CO.
3° For x < 1/2, then (3.42) transforms into
C + x O2 + (1 – 2x) CO2  2(1 – x) CO,
4° For x > 1, then from (3.42) one obtains
C + x O2 + (2x – 2) CO  (2x – 1) CO2,
5° For 1/2 < x < 1, holds this reaction
C + x O2  2(1 – x) CO + (2x – 1) CO2.
Next, for the reaction (3.42) we shall determine its
minimal coefficients by using of Moore–Penrose gener-
alized inverse matrix46.
From the chemical reaction (3.42) follows this
scheme
C O C
O
C
O
2
C 1 0 –1 –1
O 0 2 –1 –2
According to the above scheme, the reaction matrix A
of (3.42) has this form
A =
1 0 1 1
0 2 1 2
− −
− −
⎡
⎣⎢
⎤
⎦⎥
The Moore–Penrose generalized inverse matrix A+ is
A+ = AT(A AT)–1 =
1 2 1 6
1 3 1 3
1 3 0 0
1 6 1 6
/ /
/ /
/ /
/ /
−
−
−
− −
⎡
⎣
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
The reaction (3.42) reduces to this matrix form
Ax = 0, (3. 45)
where x = (x1, x2, x3, x4)T is the vector of the coefficients
of (3.42), 0 = (0, 0)T is the zero vector and T denoting
transpose.
The general solution of the matrix equation (3.45) is
x = (I – A+A)a, (3.46)
where a is an arbitrary vector and I is a unite matrix. For
a = (1, 1, 1, 1)T, one obtains
xmin = (4/3, 1, 2/3, 2/3)
T.
Then the reaction (3.42) with its minimal coefficients
attains this form
4/3 C + O2  2/3 CO + 2/3 CO2,
but not as ten Hoor asserted in57. By this proof we have
shown that his statement is paradoxical.
Also, the assumptions 1 and 2 which ten Hoor used
in57 are completely wrong, because carbon burns accord-
ing to the Boudouard’s reaction58. Wrong suppositions
can not lead to correct results.
IX. Authors of the article59 studied several chemical
reactions, but unfortunately there are given lots of erro-
neous results.
Let us mention them. In their article59 they provided
the following wrong definition: a chemical equation is a
written representation of a chemical reaction, showing
the reactants and products, they physical states, and the
direction in which the reaction proceeds.
According to the above definition a chemical equa-
tion will show like this
a A(s) + b B(g)  c C(s) + d D(g). (3.47)
For instace, if r = s = 2 in (2.1), then as a particular
case appears the reaction (3.47). Actually, it is not a defi-
nition for a chemical equation, just opposit it is a defini-
tion for a chemical reaction. Obviously the authors can
not distinguish what is a chemical reaction and what is
its chemical equation. These two things are different en-
tities given by the Definitions 2.2 and 2.1 (in a descrip-
tive form), i.e., 2.11 (in an analitical form), respectively.
It is just one reason what the above definition is wrong.
In order to be balanced certain chemical reaction is
not necessary to be known its reactants and products as it
is described in the above definition! It holds if only if re-
action is given in a conventional form, but in an opposite
case it does not hold. For instace, the chemical reaction
(3.47) given in a convntional form can be presented in an
algebrical free form too
a A(s) + b B(g) + c C(s) + d D(g) = 0. (3.48)
After determination of its coefficients, one obtains
that some of them have a negative sign and others have a
positive sign. Positive coefficients stay in front of reac-
tants and negative coefficients stay in front of products
of reaction, that means that chemical reaction is
self–adaptive. For example, in48 are balanced chemical
reactions given in an algebrical free form.
Next, we shall give an another reason why the above
definition is wrong.
Chemical equation does not have arrow mark as a re-
action, just sign for equality. It is a main difference be-
tween chemical reaction and chemical equation.
More accurately speaking, any chemical reaction
has a chemical equation, but the opposite does not
hold. Why?
Next, we shall give an explanation about it by a new
example. Let us balance the following chemical reaction
x1 Pb2O3 + x2 C  x3 Pb0.987O
+ x4 Pb3O4 + x5 CO + x6 CO2. (3.49)
From the above reaction follows this system of linear
equations
2 x1 = 0.987 x3 + 3 x4,
3 x1 = x3 + 4 x4 + x5 + 2 x6, (3.50)
x2 = x5 + x6.
In order to avoid fractional coefficients of the system
(3.50), we shall multiply its first equation by 1000, such
that one obtains this system of linear equations
I. B. RISTESKI: NEW DISCOVERED PARADOXES IN THEORY OF BALANCING CHEMICAL REACTIONS
Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522 517
2000 x1 = 987 x3 + 3000 x4,
3 x1 = x3 + 4 x4 + x5 + 2 x6, (3.51)
x2 = x5 + x6.
The systems (3.50) and (3.51) are equivalent and they
have same solution.
Now, according to the system (3.51) the chemical re-
action (3.49) will transform into this particular form
x1 Pb2000O3 + x2 C  x3 Pb987O
+ x4 Pb3000O4 + x5 CO + x6 CO2. (3.52)
Do the expression (3.52) is a correct chemical reac-
tion? No! It represents only an ordinary chemical absurd,
because the molecules Pb2000O3, Pb987O and Pb3000O4 do
not exist in chemistry. This is the cause why from chemi-
cal equation does not follow chemical reaction.
Chemical reactions (3.49) and (3.52) in a mathemati-
cal sense both are equivalent reactions, since they reduce
to the same system of linear equations, but in a chemical
sense they are not equivalent reactions. That means, that
math. equivalency  chem. equivalency.
In other words, from a mathematical point of view,
the systems (3.50) and (3.51) both are equivalent, but
from a chemical view point they are not, since they gen-
erate different chemical reactions.
The above explanation, we can articulate roughly on
this way: performing of reaction is a chemical subject,
and its balancing is a mathematical topic. This is the rea-
son why balancing of chemical reactions is pure mathe-
matical matter, but not a chemical issue.
Next, we shall continue with the balancing of the re-
action (3.49), because its general solution is necessary
for a comparative analysis of other particular chemical
reactions.
The general solution of the system (3.50) is
x4 = 2 x1/3 – 0.329 x3,
x5 = – x1/3 + 2 x2 – 0.316 x3, (3.53)
x6 = x1/3 – x2 + 0.316 x3,
where xi, (1  i  3) are arbitrary real numbers.
Balanced reaction has this form
x1 Pb2O3 + x2 C  x3 Pb0.987O
+ (2 x1/3 – 0.329 x3) Pb3O4
+ (– x1/3 + 2 x2 – 0.316 x3) CO
+ (x1/3 – x2 + 0.316 x3) CO2, (3.54)
where xi, (1  i  3) are arbitrary real numbers.
Since x4, x5 and x6 are > 0, then from (3.54) one ob-
tains this system of inequalities
2 x1/3 – 0.329 x3 > 0,
– x1/3 + 2 x2 – 0.316 x3 > 0, (3.55)
x1/3 – x2 + 0.316 x3 > 0.
From (3.55), we obtain this relation
3 x2 – 0.948 x3 < x1 < 6 x2 – 0.948 x3. (3.56)
The inequality (3.56) is necessary and sufficient con-
dition to hold the general reaction (3.54).
Now, we can analyze the general reaction (3.54) for
all possible values of x1, x2 and x3. As particular reactions
of (3.54) we shall derive the following cases.
1° For x1 = 3, x2 = 2.4 and x3 = 4.5, from (3.53) fol-
lows x4 = 0.5195, x5 = 2.378 and x6 = 0.022, i.e., one ob-
tains this particular reaction
3 Pb2O3 + 2.4 C  4.5 Pb0.987O
+ 0.5195 Pb3O4 + 2.378 CO
+ 0.022 CO2. (3.57)
2° For x1 = 0, the reaction (3.54) transforms into this
particular reaction
0.329 x3 Pb3O4 + x2 C  x3 Pb0.987O
+ (2 x2 – 0.316 x3) CO
+ (– x2 + 0.316 x3) CO2, (3.58)
(0.158 x3 < x2 < 0.316 x3).
3° For x2 = 0.158 x3, from (3.58) one obtains
0.329 Pb3O4 + 0.158 C  Pb0.987O
+ 0.158 CO2. (3.59)
4° For x2 = 0.316 x3, the reaction (3.58) becomes
0.329 Pb3O4 + 0.316 C  Pb0.987O
+ 0.316 CO. (3.60)
5° For x2 < 0.158 x3, from (3.58) follows
0.329 x3 Pb3O4 + x2 C + (0.316 x3 – 2 x2) CO
 x3 Pb0.987O + (– x2 + 0.316 x3) CO2, (3.61)
where x2 and x3 are arbitrary real numbers.
6° For x2 > 0.316 x3, the reaction (3.58) becomes
0.329 x3 Pb3O4 + x2 C + (x2 – 0.316 x3) CO2
 x3 Pb0.987O + (2 x2 – 0.316 x3) CO, (3.62)
where x2 and x3 are arbitrary real numbers.
7° For x2 = 0, from (3.54) one obtains this particular
reaction
x1 Pb2O3 + (x1/3 + 0.316 x3) CO
 x3 Pb0.987O + (2 x1/3 – 0.329 x3) Pb3O4
+ (x1/3 + 0.316 x3) CO2, (3.63)
where x1 and x3 are arbitrary real numbers.
From (3.63) follows these inequalities
x1/3 + 0.316 x3 > 0 and 2 x1/3 – 0.329 x3 > 0.
From this system follows this inequality
– x1/0.048 < x3 < 2 x1/0.987. (3.64)
The reaction (3.63) holds if only if the inequality
(3.64) is satisfied.
8° For x1 < 0.4935 x3, from (3.63) follows
x1 Pb2O3 + (0.329 x3 – 2 x1/3) Pb3O4 + (x1/3 +
0.316 x3) CO  x3 Pb0.987O
+ (x1/3 + 0.316 x3) CO2, (3.65)
where x1 and x3 are arbitrary real numbers.
9° For x1 = 0.4935 x3, from (3.63) follows
0.4935 Pb2O3 + 0.4805 CO
 Pb0.987O + 0.4805 CO2. (3.66)
I. B. RISTESKI: NEW DISCOVERED PARADOXES IN THEORY OF BALANCING CHEMICAL REACTIONS
518 Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522
10° For x3 = 0, from (3.54) one obtains this particular
reaction
x1 Pb2O3 + x2 C  (2 x1/3) Pb3O4
+ (– x1/3 + 2 x2) CO + (x1/3 – x2) CO2, (3.67)
(3 x2 < x1 < 6 x2).
11° For x1 = 3 x2, from (3.67) follows
3 Pb2O3 + C  2 Pb3O4 + CO. (3.68)
12° For x1 = 6 x2, from (3.67) one obtains
6 Pb2O3 + C  4 Pb3O4 + CO2. (3.69)
13° For x1 < 3 x2, from (3.67) follows
x1 Pb2O3 + x2 C + (x2 – x1/3) CO2
 (2 x1/3) Pb3O4 + (– x1/3 + 2 x2) CO, (3.70)
where x1 and x2 are arbitrary real numbers.
14° For x1 > 6 x2, the reaction (3.67) transforms into
x1 Pb2O3 + x2 C + (x1/3 – 2 x2) CO
 (2 x1/3) Pb3O4 + (x1/3 – x2) CO2, (3.71)
where x1 and x2 are arbitrary real numbers.
The authors of the article59 gave this statement: Bal-
ancing the chemical equation (with three molecules)
means finding the smallest whole numbers x1, x2 and x3
(as its coefficients).
The above statement does not hold for every reaction.
It has just particular meaning and holds if only if a
chemical reaction has atoms with integers, in an opposite
case, when the reaction contains atoms with fractional
oxidation numbers, it does not hold. For example, see the
previous reactions (3.57), (3.58) and so on.
Also, in the article59 is »balanced« this reaction
x1 Cu + x2 HNO3  x3 Cu
+2
+ x4 NO2 + x5 NO3
– + x6 H2O, (3.72)
like this
3 Cu + 4 HNO3  3 Cu
+2
+ 2 NO + 2 NO3
– + 2 H2O. (3.73)
The last reaction (3.73) is wrong.
The correct form of balanced reaction (3.72) is
x1 Cu + x2 HNO3  x1 Cu
+2 + (x2/2) NO2
+ (x2/2) NO3
– + (x2/2) H2O, (3.74)
where x1, x2 > 0 are arbitrary real numbers.
Next reaction was »balanced« in59 too
x1 OH
– + x2 SnO2
–2 + x3 Bi(OH)3
 x4 Bi + x5 SnO3
–2 + x6 H2O, (3.75)
in this form
0 OH– + 3 SnO2
–2 + 2 Bi(OH)3
 2 Bi + 3 SnO3
–2 + 3 H2O. (3.76)
The general form of the reaction (3.75) is
x1 OH
– + x2 SnO2
–2 + (2x2/3 – x1/3) Bi(OH)3
 (2x2/3 – x1/3) Bi
+ x2 SnO3
–2 + x2 H2O, (x1 < 2x2). (3.77)
For x1 = 0 and x2 = 3, from (3.77) immediately fol-
lows (3.76), like a particular reaction.
Another reaction is »balanced« in59
x1 CH3CH2OH + x2 Cr2O7
–2 + x3 H
+
 x4 CH3CO2H + x5 Cr
+3 + x6 H2O, (3.78)
in this form
3 CH3CH2OH + 2 Cr2O7
–2 + 16 H+
 3 CH3CO2H + 4 Cr
+3 + 11 H2O. (3.79)
The general form of the reaction (3.78) is
x1 CH3CH2OH + x2 Cr2O7
–2
+ (– 4x1 + 14 x2) H
+  x1 CH3CO2H
+ 2x2 Cr
+3 + (– x1 + 7 x2) H2O, (3.80)
(0 < x1 < 7x2/2).
For x1 = 3 and x2 = 2, from (3.80) immediately fol-
lows (3.79), like a particular reaction.
In59 the authors only determined the particular reac-
tions (3.76) and (3.79), but considered reactions (3.75)
and (3.78) have general forms with two arbitrary param-
eters, given by (3.77) and (3.80), respectively.
X. As a last paradox we discovered in the theory of
balancing chemical equations is the case considered be-
low. The authors of the article60 studied this chemical re-
action
CH0.686O0.32 + w H2O + m O2 + 3.76 m N2
 x1 H2 + x2 CO + x3 CO2 + x4 H2O
+ x5 CH4 + 3.76 m N2. (3.81)
Immediately from (3.81) we have seen its absurdity.
In this reaction H2O appears as a reactant and as a prod-
uct at the same time. Is it logical? No! It is just one
thing.
Another illogical thing is not reacted N2. Nitrogen
appears in (3.81) as reactant 3.76m N2 and 3.76m N2 as a
product at the same time.
Also, the treatment of hydrogen as a product is illogi-
cal. Burning of CH0.686O0.32 presented by (3.81) does not
give H2 as a product!
We think that the reaction
x1 CH0.686O0.32 + x2 O2 + x3 N2  x4 CO
+ x5 CO2 + x6 H2O + x7 CH4 + x8 N2O3, (3.82)
which represents a correct version of (3.81) is very in-
teresting for chemistry.
I. B. RISTESKI: NEW DISCOVERED PARADOXES IN THEORY OF BALANCING CHEMICAL REACTIONS
Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522 519
From (3.82) one obtains this scheme
C
H
0.
68
6O
0.
32
O
2
N
2
C
O
C
O
2
H
2O
C
H
4
N
2O
3
C 1 0 0 –1 –1 0 –1 0
H 0.686 0 0 0 0 –2 –4 0
O 0.32 2 0 –1 –2 –1 0 –3
N 0 0 2 0 0 0 0 –2
Reaction matrix A of (3.82) has this form
A =
1000 0 0 1 1 0 1 0
0686 0 0 0 0 2 4 0
0320 2 0 1 2 1 0 3
0 000
.
.
.
.
− − −
− −
− − − −
0 2 0 0 0 0 2−
⎡
⎣
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
The reaction (3.82) reduces to the following system
of linear equations
x1 = x4 + x5 + x7,
0.686 x1 = 2 x6 + 4 x7, (3.83)
0.32 x1 + 2 x2 = x4 + 2 x5 + x6 + 3 x8,
2 x3 = 2 x8.
The general solution of the system (3.83) is
x5 = 1.977 x1/4 + x2/2 – 3 x3/4 – 3 x4/4,
x6 = – 1.337 x1/2 + x2 – 3 x3/2 + x4/2, (3.84)
x7 = 2.023 x1/4 – x2/2 + 3 x3/4 – x4/4,
x8 = x3,
where xi > 0 (1  i  4) are arbitrary real numbers.
Since xi > 0 (5  i  8), then from (3.84) one obtains
this system of inequalities
1.977 x1 + 2 x2 – 3 x3 – 3 x4 > 0,
– 1.337 x1 + 2 x2 – 3 x3 + x4 > 0, (3.85)
2.023 x1 – 2 x2 + 3 x3 – x4 > 0.
From (3.85) immediately follows this inequality
1.337 x1 – x4 < 2 x2 – 3 x3 < 2.023 x1 – x4. (3.86)
Reaction (3.82) holds if and only if the condition
(3.86) is satisfied.
The reaction (3.82) contains three subgenerators
which induce a topology of its solutions, but we omitted
it, since it will be a subject of the author’s next research.
The other particular cases of (3.82) are not consid-
ered because we took into account the Remark 3.7.
For instance, for x1 = 2, x2 = 4, x3 = 2 and x4 = 1, from
(3.84) one obtains this particular solution x5 = 0.7385, x6
= 0.163, x7 = 0.2615 and x8 = 2. Then (3.82) becomes
2 CH0.686O0.32 + 4 O2 + 2 N2  CO
+ 0.7385 CO2 + 0.163 H2O
+ 0.2615 CH4 + 2 N2O3.
Now, we shall determine the minimal coefficients of
the reaction (3.82). The Moore–Penrose generalized ma-
trix A+ will have this form
A+ =
0334436538275454 0 037800281937753
0164755899838
. .
.
−
− 609 0 019118215665491
0123566924878957 0 0143386617
.
. .− 49118
0338363333974391 0 077822527216911
0 2559853
−
−
. .
. 84055086 0 068263419384166
0 257141220018618 014519
.
. .− 3039626989
0 071214743695069 0183886228538830
012
− −. .
. 3566924878957 0 014338661749118−
⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎢
⎢
⎢
.
−0 048377720464309 0 036283290348232
01715167592608
. .
. 28 0128637569445621
0128637569445621 03464781770
−
−
.
. . 84216
0 003380429711109 0 002535322283332
0 0891388
−
−
. .
. 09341523 0 066854107006142
0104876595295905 0 07865
.
. .− 7446471929
0 044141518588323 0 033106138941242
0128
. .
.− 637569445621 0153521822915784−
⎤
⎦
⎥
⎥
⎥
⎥
⎥
⎥
⎥
⎥
.
The reaction (3.82) reduces to the matrix form (3.45),
whose general solution is given by the expression (3.46).
For a = (1, 1, 1, 1, 1, 1, 1, 1)T, one obtains
xmin = (x1, x2, x3, x4, x5, x6, x7, x8)
T,
where
x1 = 1.241594646560723466,
x2 = 1.574780831709873263,
x3 = 0.568914376217595052,
x4 = 0.721001830633894027,
x5 = 0.433611414778957395,
x6 = 0.251904161474584061,
x7 = 0.086981401147872044,
x8 = 0.568914376217595052.
5 DISCUSSION
We cannot always trust chemical experiments! We
cannot always trust mathematics either, for it can mis-
lead us unless we define away the problem area. How-
ever, we surely can trust pure logic – no questionable ex-
periments or unusual mathematical operations.
Are they methods when somebody can find counter-
examples on every step? Obviously, the answer is nega-
tive! It is merely a pale picture of the old chemically irra-
tional traditionalism of the past chemistry.
As it is showed by all counterexamples given in this
work the traditional procedures for balancing chemical
reactions are inconsistent. These particular procedures,
we hope that they are the last traditional unsuccessful ap-
proach of balancing chemical reactions. Long time ago
chemistry lost the battle with mathematics in sense of
balancing chemical reactions.
Before we finish this discussion, we would like to
stress here, that our facts for arguing are founded by vir-
tue of scientific results.
I. B. RISTESKI: NEW DISCOVERED PARADOXES IN THEORY OF BALANCING CHEMICAL REACTIONS
520 Materiali in tehnologije / Materials and technology 45 (2011) 6, 503–522
From current view point of balancing chemical equa-
tions, we feel free to tell that traditional approach of bal-
ancing chemical reactions is only a history, and Jones’
problem44, which was as one of the hardest problems of
balancing chemical reactions is completely solved45.
There exists a completely satisfactory ways of avoid-
ing paradoxes7,34,45,46,48–50. The theory used is based on the
idea of formal approach of balancing chemical reactions.
In these works completely new general highly sophisti-
cated methods are developed for balancing chemical re-
actions and their stability by virtue of the theory of gen-
eralized matrix inverse using Moore–Penrose, Drazin
and von Neumann matrices. By these methods chemistry
is cleaned from old traditional inconsistent procedures
for balancing chemical reactions, such that is open a
brand new direction for development of this topic in
chemistry and its foundation on genuine principles. That
is the newest trend in chemistry about this issue, which
showed that traditionalism in chemistry is over.
6 CONCLUSION
By this work the consideration of paradoxes in chem-
istry will begin very seriously as a special object and in
any way it will increase researchers’ carefulness to avoid
the appearance of paradoxes. Sure, no perfect science!
Appearance of paradoxes is always possible.
It is more than certain, that this work opened doors
for the next research in chemistry for its diagnostic of
paradoxes and their resolution. It will accelerate the new-
est contemporary research in chemistry and it will de-
stroy all barriers which hamper the development of
chemistry and lay its foundation on genuine scientific
principles.
This work affirms:
• that all formally provable mathematical methods are
true if chemical reactions are considered as a consis-
tent formal system,
• that all mathematical truths can be formally provable,
and
• that the new branch Foundation of Chemistry
proves the consistency and completeness of the for-
mal approach of balancing chemical reactions and
that it will be a special kind of chemistry, i.e., it will
be a finite theory which contains only perfectly well
known concepts with true axioms and positive con-
clusions. It affirms that the principles used in the
mathematical approach of balancing chemical reac-
tions, will not lead to contradictions.
After that, what is proven about the paradoxes in this
work is there a general chemistry? Or it should be
refounded on genuine principles as elementary chemis-
try?
The replies to these two questions are looking for
deep reform in chemistry in a formalistic way, because in
the opposite case paradoxes will be a stumbling block for
a long time. To avoid this awkward position, reforms in
chemistry are needed as soon as possible.
Sure, that the above topic is considered in a rough
form, but it happens with every pioneer’s job.
ACKNOWLEDGEMENT
The author would like to thank Prof. Dr Valery C.
Covachev, from Bulgarian Academy of Sciences for
reading the work and for his remarks which improved
the quality of this work.
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N. A. KATANAHA, L. B. GETSOV: FEATURES OF CREEP IN CONDITIONS OF LONG OPERATION
CHARACTERISTICS OF CREEP IN CONDITIONS OF
LONG OPERATION
ZNA^ILNOSTI LEZENJA PRI DOLGOTRAJNI UPORABI
Nikolay A. Katanaha1, Leonid B. Getsov2
1Saint-Petersburg State Polytechnical University, Saint-Petersburg 194 021 Hlopina 13, Russia
NPO CKTI, 195213 St. Petersburg, 195213 Zanevski pr. 43 apt. 89, Russia
katanaha@mail.ru, guetsov@yahoo.com
Prejem rokopisa – received: 2011-07-28; sprejem za objavo – accepted for publication: 2011-10-20
The creep of materials is considered in various temperature ranges: at low, at elevated and at high temperatures. Different
approaches of extrapolation of experimental data for creep curves are shown. For the available isochronous curves of steels
R2M and EP291, the coefficients for the upgraded formula Soderberg were obtained. The isochronous creep curves and the
curves with a partition into first and second stages were constructed by using the coefficients obtained. Calculation of the
coefficients was performed using the computational mathematics package MathCad and OriginPro. The Arrhenius equation for
description of the temperature dependence of creep rate in a narrow temperature range is presented.
The results of this work are a new method of determining the extrapolated values of material creep resistance relating to
long-life, a verification of this method for experimental data of the steels R2M and EP291.
Key words: steel, creep, isochronous curves, creep curve, Norton formula, Soderberg formula, Arrhenius equation for a creep
rate, steels R2M, EP291, 20Cr12WNiMoV
Lezenje materialov je obravnavano za razli~na obmo~ja temperatur: pri nizki, povi{ani in visoki temperaturi. Prikazani so
razli~ni na~ini ekstrapolacije eksperimentalnih podatkov s krivulj lezenja. Za razpolo`ljive izohrone krivulje za jekli R2LM in
EP291 so bili dolo~eni koeficienti za izbolj{avo Soderbergove ena~be. Izohrone krivulje in krivulje z delitvijo na prvo in drugo
stopnjo so oblikovane z upo{tevanjem prve in druge faze in z uporabo dolo~enih koeficientov. Koeficienti so izra~unani z
uporabo matemati~nih izra~unov na podlagi matemati~nih paketov MathCad in OriginPro. Arheniusova ena~ba za odvisnost
hitrosti lezenja od temperature je prikazana za ozko obmo~je temperature.
Rezultati tega dela sta nova metoda za dolo~anje ekstrapoliranih vrednosti odpornosti materiala proti lezenju pri dolgi uporabi in
verifikacija metode na podlagi eksperimentalnih podatkov za jekli R2M in EP291.
Klju~ne besede: jeklo, lezenje, izohrone krivulje, krivulje lezenja, Nortonova ena~ba, Soderbergova ena~ba, Arheniusova
odvisnost za hitrost lezenja, jekla R2M, EP291, 20Cr12WNiMoV
1 INTRODUCTION
The use of isochronous curves of creep in the
strength calculations of details for power plants and nec-
essary demands of increasing resource of these details up
to 300 000 h or more, requires the need of resolving of
several problems connected with the processing of ex-
perimental data of creep.
Considering the features of the steel creep in condi-
tions of insignificant residual strain, the stress change
over time can be neglected.
In these conditions creep should be considered in var-
ious temperature ranges:
a) At low temperatures. Creep is characterized by the
first stage of transient creep. For example, further
course of the creep process is almost inhibited with
increasing time, leading to a lack of long damage at
stresses less than the yield stress (Figure 1a);
b) At elevated temperatures. As test results of various
materials show the rate of steady creep depends on
the duration of tests: more time, lower slope of the
creep curve. In this connection it should be recog-
nized, that the adequacy of isochronous creep curves,
obtained by extrapolation, is very low (Figure 1b).
The corresponding curves are lower than those,
which were obtained by direct experiment. It gives a
conservative estimate of the deformation process de-
tails by soft loading. However, when the processes of
stress relaxation are defined (hard load – uneasiness
strain) and when the durability assessment is deter-
mined in accordance with the equivalent stress, we
find higher values of durability and safety margins.
These conditions are the material behavior at stress
raisers, the stress relaxation at the high temperature
fasteners, the redistribution of stresses in the steam
pipeline bending, in the blades and turbine disks;
c) At high temperatures, as it is well known, the pro-
cesses of diffusion creep take place without the first
stage of creep (Figure 1c).
The second stage of creep is followed by the third
(the stage of accelerated creep) depending on the defor-
mation ability of the material and the stress level at the
elevated and high temperatures.
2 EXPERIMENTAL
Consequently, using the experimental data of creep at
the terminal duration, different approaches for the ex-
trapolation should be taken into account.
Materiali in tehnologije / Materials and technology 45 (2011) 6, 523–527 523
UDK 669.14.018:620.172.251 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)523(2011)
An adequate description of creep at low temperatures
can be obtained by using the expression:
 exp( )p a c c t= ⋅ ⋅ − ⋅ (1)
where: a A= ⋅ k , c C= ⋅1 ( – stress, t – time, A, C, k, l
– constants, depending on temperature). For high tem-
peratures, the Norton formula gives the consistent re-
sults:
p b= (2)
where: b B= ⋅m ( – stress, B, m – constants, depending
on temperature).
3 THE RESEARCH RESULTS
At elevated temperatures, at which details power
plants operate usually, the use of sum (1) and (2) for de-
termining the rate of creep:
 exp( )p a c c t b= ⋅ ⋅ − ⋅ + (3)
is possible for time not exceeding the experimental. To
solve the problem of extrapolation of the creep data of
long duration, the experimental data of long duration
have been used. These data have been received by A. A.
Chizhik for the steel R2M pearlitic at 50 °C, 525 °C,
550 °C (Figure 2), the steels of martensitic class
18Cr11MoVNbNi (EP291) (Figure 3) and
20Cr12WNiMoV (Figure 4) at 550 °C 1.
Dependence of the duration of the first and second
stages of creep has been determined in accordance with
the test base for steel R2M (Figure 5) as well as for steel
EP291.
Taking this in consideration, the use of the Norton
formula with coefficients derived from a relatively short
N. A. KATANAHA, L. B. GETSOV: FEATURES OF CREEP IN CONDITIONS OF LONG OPERATION
524 Materiali in tehnologije / Materials and technology 45 (2011) 6, 523–527
Figure 3: Approximation of the experimental creep data of the steel
EP291 at 550 °C, presented in 1
Figure 2: Approximation of the experimental creep data of the steel
R2M at 500 °C (a), 525 °C (b) and 550 °C (c), presented in 1
Slika 2: Aproksimacija eksperimentalnih podatkov za jeklo R2M pri
500 °C (a), 525 °C (b) in 550 °C, ref. 1
Figure 1: Schematic diagram of creepa) at low b) at elevated c) at
high temperature
Slika 1: Shemati~na krivulja lezenja a) pri nizki, b) pri povi{ani in c)
pri visoki temperature
Figure 4: Creep curve of the steel 20Cr12WNiMoV at 550 °C
Slika 4: Krivulja lezenja za jeklo 20Cr12WNiMoV pri 550 °C
a)
b)
c)
duration of data, can give very significant errors in the
calculation regarding the operation of equipment with
very long lifetime (more than 100 000 h) both due to the
neglect of the first stage of creep and the overrated val-
ues of the exponent m. The various attempts of descrip-
tion of this phenomenon led to the need to modify the
Soderberg formula as:
p a c t b t ta c b b= ⋅ ⋅ − − ⋅ ⋅ + ⋅ ⋅ ⋅1 1 12 2 2 21  ( exp( )) (4)
where: ( – stress, t – time, a1, a2, b1, b2, b3, c1, c2 –
constants depending on temperature).
The expression for the creep rate is:
 ( exp( ))p a c c t b ta c c b b= ⋅ ⋅ ⋅ ⋅ − − ⋅ ⋅ + ⋅ ⋅1 1 1 12 2 2 2 21    (5)
N. A. KATANAHA, L. B. GETSOV: FEATURES OF CREEP IN CONDITIONS OF LONG OPERATION
Materiali in tehnologije / Materials and technology 45 (2011) 6, 523–527 525
Table 1: Values of the coefficients of equation (4)
Tabela 1: Vrednosti koeficientov ena~be (4)
R2M 18Cr11MoVNbNi (EP291)
500 °C 525 °C 550 °C 550 °C
I II I II I II I II
a1 6.322E-8 1.402E-7 6.942E-7 1.314E-6 1.717E-5 1.043E-5 7.383E-6 6.626E-6
a2 2.028 1.854 1.596 1.448 1.031 1.126 1.202 1.203
b1 1.435E-10 8.803E-11 5.510E-11 9.066E-11 3.217E-12 4.293E-12 1.768E-7 2.059E-7
b2 –3.752E-1 –3.915E-1 –2.784E-1 –2.914E-1 –1.843E-1 –1.554E-1 –5.973E-1 –6.215E-1
b3 1.988 2.121 2.091 2.017 2.643 2.515 1.070 1.087
c1 9.341E-5 6.374E-5 1.181E+8 6.662E-7 1.016E-3 1.175E-3 1.294E-5 2.718E-5
c2 7.793E-1 8.868E-1 -5.270 -5.148 0 0 1.214 1.197
SD 1.937E-2 1.776E-2 2.796E-2 2.718E-2 1.290E-2 1.088E-2 6.917E-2 7.044E-2
Figure 5: Duration of the first and second stages of creep for the steel
R2M at 500 °C (a – 100 MPa, b – 120 MPa, c – 140 MPa, d – 160
MPa), 525 °C (e – 80 MPa, f – 100 MPa, g – 120 MPa), 550 °C (h – 60
MPa, i – 100 MPa, j – 120 MPa)
Slika 5: Trajanje prve in druge faze lezanja za jeklo R2M pri 500 °C
(a- 100 MPa, b- 120 MPa, c- 140 MPa, d- 160 MPa), 525 °C (e- 80
MPa, f- 100 MPa, g- 120 MPa), 550 °C (h- 6a MPa, i- 100 MPa, j- 120
MPa)
Figure 6: Isochronous curves of steel R2M at 500 °C (a), 525 °C (b)
and 550 °C(c) 1 – 100 h; 2 – 500 h; 3 – 1 000 h; 4 – 5 000 h; 5 – 10
000 h; 6 – 20 000 h; 7 – 30 000 h; 8 – 40 000 h; 9 – 50 000 h; 10 – 60
000 h; 11 – 70 000 h; 12 – 80 000 h; 13 – 90 000 h; 14 – 100 000 h; 15
– 300 000 h; 16 – 500 000 h.
Slika 6:. Izohrone krivulje za jeklo R2M pri 500 °C (a), 525 °C (b), in
550 °C (c). 1 – 100 h; 2 – 500 h; 3 – 1 000 h; 4 – 5 000 h; 5 – 10000 h;
6 – 20 000 h; 7 – 30 000 h; 8 – 40 000 h; 9 – 50 000 h; 10 – 60 000 h;
11 – 70 000 h; 12 – 80 000 h; 13 – 90 000 h; 14 – 100 000 h; 15 – 300
000 h; 16 – 500 000 h.
The values of the coefficients of equation (4) for the
investigated materials are given in Table 1. The coeffi-
cients are calculated in accordance with the minimum
standard deviation for the experimental points and the
proposed approximation. Calculation of the coefficients
has been performed by means of the computational
mathematics package MathCad and OriginPro.
The standard deviation has been used for comparing
the results obtained during the calculation of the creep
deformation with the original data calculated from:
SD
n
y y
y
i
ii
n
= ⋅
−⎛
⎝
⎜⎜
⎞
⎠
⎟⎟
=
∑1
1
2
(6)
In some cases, the dependence (4) for certain values
of the coefficients may give an inadequate representation
of the isochronous curves of creep that has double bends.
For example, this situation occurred during processing
the experimental data for the steel R2M at 550 °C. In
case of these incorrect results on the basis of the analy-
sis, it is necessary to use of c2 = 0 in equation (4).
Data in Table 1 show, that SD values are at low level.
The isochronous creep curves of steel have been built ac-
cording to the values of the coefficients listed in Table 1
(Figure 6).
Comparing Figures 2 and 3 with Figures 6 and 7,
we can see significant differences in the type of curves.
It is related to the effectiveness of the proposed approxi-
mation by means of the dependence (4) and the use di-
rectly the experimental points for finding the values of
the coefficients, rather than curves.
Unfortunately, methods of calculation that have been
used for the calculated construction of the isochronous
creep curves, were based on the unique experimental
data of long duration obtained for a metal melting and
these methods didn’t make provision for the scatter of
N. A. KATANAHA, L. B. GETSOV: FEATURES OF CREEP IN CONDITIONS OF LONG OPERATION
526 Materiali in tehnologije / Materials and technology 45 (2011) 6, 523–527
Figure 8: Comparison of experimental values of creep rate for the
steel R2M with the extrapolated and interpolated values
Slika 8: Primerjava eksperimentalnih hitrosti lezenja za jeklo R2M z
ekstrapoliranimi in interpoliranimi vrednostmi
Figure 7: The isochronous curves of steel EP291 at 550 °C 1 – 100 h;
2 – 500 h; 3 – 1 000 h; 4 – 5 000 h; 5 – 10 000 h; 6 – 20 000 h; 7 – 30 000
h; 8 – 40 000 h; 9 – 50 000 h; 10 – 60 000 h; 11 – 70 000 h; 12 – 80 000
h; 13 – 90 000 h; 14 – 100 000 h; 15 – 300 000 h; 16 – 500 000 h
Slika 7: Izohromne krivulje za jeklo EP291 pri 550 °C 1 – 100 h; 2 –
500 h; 3 – 1 000 h; 4 – 5 000 h; 5 – 10 000 h; 6 – 20 000 h; 7 – 30 000
h; 8 – 40 000 h; 9 – 50 000 h; 10 – 60 000 h; 11 – 70 000 h; 12 –
80 000 h; 13 – 90 000 h; 14 – 100 000 h; 15 – 300 000 h; 16 – 500 000 h
Table 2: Creep rate
Tabela 2: Hitrost lezenja
t/ h 100 MPa
T = 500 °C T = 500 °C ex-trapolation T = 525 °C
T = 525 °C in-
terpolation T = 550 °C
T = 550 °C ex-
trapolation
100 2,876E-07 1,538E-07 5,725E-07 7,713E-07 1,891E-06 1,070E-06
500 1,448E-07 2,392E-08 1,923E-07 4,535E-07 1,280E-06 2,490E-07
1 000 1,073E-07 1,590E-08 1,258E-07 3,124E-07 8,251E-07 1,454E-07
5 000 5,821E-08 4,107E-08 7,823E-08 9,236E-08 1,405E-07 1,023E-07
10 000 4,488E-08 3,451E-08 6,450E-08 7,310E-08 1,139E-07 8,969E-08
20 000 3,460E-08 2,650E-08 5,318E-08 6,039E-08 1,002E-07 7,860E-08
30 000 2,972E-08 2,270E-08 4,750E-08 5,401E-08 9,296E-08 7,276E-08
40 000 2,668E-08 2,034E-08 4,385E-08 4,990E-08 8,816E-08 6,888E-08
50 000 2,454E-08 1,867E-08 4,121E-08 4,693E-08 8,461E-08 6,602E-08
60 000 2,291E-08 1,742E-08 3,917E-08 4,463E-08 8,182E-08 6,376E-08
70 000 2,163E-08 1,642E-08 3,752E-08 4,278E-08 7,952E-08 6,192E-08
80 000 2,057E-08 1,561E-08 3,615E-08 4,123E-08 7,759E-08 6,036E-08
90 000 1,968E-08 1,492E-08 3,499E-08 3,992E-08 7,592E-08 5,903E-08
100 000 1,892E-08 1,433E-08 3,397E-08 3,878E-08 7,446E-08 5,785E-08
the data depending on the characteristics of melting and
the variations in the mode of heat treatment. Thus the use
of these methods in practice requires the application of
certain stocks in the deformations (or the stresses).
Next step was getting a universal dependence of
isochronous curves on temperature for any values (in a
relatively narrow temperature range). The Arrhenius
equation was used to describe the temperature depend-
ence of creep rate in a narrow temperature range.
 p p ec
U
RT= −0 (7)
where
U – activation energy of creep,
R = 8,32 · 10–3 kJ/mol·K – gas constant,
p0 – constant.
Table 2 and Figure 8 show the values of creep rate
obtained in the experiment by means of extrapolation
and interpolation on temperature.
The comparison of the experimental values of the
creep rate with the extrapolated and interpolated values
on temperature in dependence on time, showed that their
differences were 3–30 % and quite acceptable.
4 CONCLUSION
The new method of determining the extrapolated val-
ues of material creep resistance related to long-life was
proposed.
The verification techniques relating the experimental
data obtained previously for the steels R2M and EP291
by A. A. Chizhik were made.
5 REFERENCES
1 A. A. Lanin, V. S. Balina. Heat-resistant metals and alloys.
EnergoTeh. Saint-Petersberg, 2006, 221 p.
2 L. B. Getsov, H. A. Katanaha, I. P. Popova. Methods of calculation of
the creep characteristics at the first and second stages on the basis of
the test results on relaxation using a limited number of isochronous
creep curves. Problems of mechanical engineering, 13 (2010) 6,
35–41
N. A. KATANAHA, L. B. GETSOV: FEATURES OF CREEP IN CONDITIONS OF LONG OPERATION
Materiali in tehnologije / Materials and technology 45 (2011) 6, 523–527 527

M. PIRNAT et al.: A THERMODYNAMIC AND KINETIC STUDY OF THE SOLIDIFICATION AND DECARBURIZATION ...
A THERMODYNAMIC AND KINETIC STUDY OF THE
SOLIDIFICATION AND DECARBURIZATION OF
MALLEABLE CAST IRON
TERMODINAMI^NA IN KINETI^NA ANALIZA STRJEVANJA IN
RAZOGLJI^ENJA BELEGA LITEGA @ELEZA
Miran Pirnat1, Primo` Mrvar2, Jo`e Medved2
1SIJ, Acroni, d. o. o, Jesenice, Slovenia,
2University of Ljubljana, Faculty of Natural Sciences and Engineering, Department of Materials and Metallurgy, A{ker~eva 6,
1000 Ljubljana, Slovenia
miran.pirnat@acroni.si
Prejem rokopisa – received: 2011-04-05; sprejem za objavo – accepted for publication: 2011-10-05
An analysis of the solidification and decarburization of white-heart malleable cast iron (MCI) is presented. The solidification
and decarburization courses were examined with simple and differential scanning calorimetry. The microstructure
characteristics and the physical properties of the white-heart malleable cast iron changed during the decarburization process.
Also, the electrical resistivity changed with the change of carbon contents and the macro-and microstructures. Based on this
hypothesis, a measuring method for simultaneous measurements of the electrical resistivity and dimensional variations during
the decarburization process of white-heart malleable cast iron was developed. In addition, a physico-mathematical model was
developed to follow the carbon concentration and to determine the depths of the decarburization zone during the decarburization
process. The decarburization process was presented as a function of the specific electrical conductivity, the carbon concentration
and the decarburization time.
Keywords: white-heart malleable cast iron (MCI), thermal analysis,electrical resistivity, specific electrical resistivity,
decarburization time, depth of decarburization zone
^lanek opisuje spremljanje strjevanja in razoglji~enja belega litega `eleza. Potek strjevanja in razoglji~enja sta bila preiskana z
enostavno in diferen~no vrsti~no kalorimetrijo. Med procesom razoglji~enja belega litega `eleza se spreminjajo zna~ilnosti
zgradbe in fizikalne lastnosti. Prav tako se zaradi spremembe koncentracije ogljika ter makro- in mikrostrukture spreminja tudi
elektri~na upornost. Na tej hipotezi je bila razvita merilna metoda isto~asnega merjenja elektri~ne upornosti in dimenzijskih
sprememb med procesom razoglji~enja belega litega `eleza. Razvit je fizikalno-matemati~ni model, s katerim je mo`no med
procesom razoglji~enja spremljati koncentracijo ogljika in dolo~iti globino razoglji~enja. Potek procesa razoglji~enja je
prikazan kot funkcija specifi~ne elektri~ne upornosti, koncentracije ogljika in ~asa razoglji~enja.
Klju~ne besede: belo lito `elezo, termi~na analiza, elektri~na upornost, specifi~na elektri~na upornost, ~as razoglji~enja, globina
razoglji~enja
1 INTRODUCTION
White-heart malleable cast iron (MCI) was prepared
from a chilled hypoeutectic iron alloy. Afterwards, it was
decarburized to achieve adequate mechanical properties.
The morphology of the solidified phases, the temperature
regions of the corresponding reactions and the formed
phases were determined with a thermodynamic analysis
of the MCI solidification and decarburization process.
The fraction of pearlite and the heat treatment1 are essen-
tial to obtain the desired properties of the MCI. The fol-
lowing methods were used to examine the solidification
process and solid-state transformations: simple thermal
analyses (TA), dilatometric analyses, and simultaneous
thermal analyses (DSC). An "in situ" measuring appara-
tus, as a part of the laboratory equipment, was also de-
veloped to follow the electrical resistivity during the
decarburization process. The goal of the examination
was to design a model for the "in situ" monitoring of the
decarburization process by determining the carbon con-
centrations and the depths of decarburization zone. The
thermal analyses could be used for the quality control of
the MCI, since it made it possible to determine the met-
allurgical quality of the cast iron in the shortest possible
time. The chemical composition and the nucleation con-
ditions determined the obtained microstructure.2–7 The
simple thermal analysis made it possible to determine the
reference liquidus temperature and the temperatures of
the transformations and to forecast the latter properties
of the castings.8 In the solidification of the MCI it was
important that all the remnant melt solidified entirely as
a chill in the form of a cementite eutectic. The simple
thermal analysis and dilatometric curves helped to exam-
ine the solidification process and the cooling of spheroi-
dal-graphite cast iron (Figure 1).9
The as-cast MCI was decarburized in approximately
40 h in an oxidizing atmosphere at 1050 oC, when the
cast microstructure changed with the reactions:
1. Formation of austenite:
(Fe + Fe3C)  Fe
2. Decomposition of cementite:
Fe3C  3Fe + C(malleablizedgraphite )
3. Decarburization process:
Materiali in tehnologije / Materials and technology 45 (2011) 6, 529–535 529
UDK 669.131.2:536.7 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)529(2011)
C(malleablized graphite) and [C] + H2O(g) = CO(g) + H2(g)
4. Precipitation of ferrite:
Fe  Fe
5. Formation of pearlite:
Fe  (Fe + Fe3C)
6. Precipitation of graphite:
Fe  Graphite(tertiary)
The decarburization proceeded predominantly by the
reactions:
[C]–Fe + H2O(g) = CO(g) + H2(g) (1)
and
Cmalleablized graphite + H2O(g) = CO(g) + H2(g) (2)
The decarburization process of steel, such as electri-
cal steel,10,11 proceeded by the decarburization reaction
(1). The decarburization process of MCI started with a
carbon loss in austenite, [C]–Fe (reaction 1), and then it
was continued by Cmalleablized graphite loss according to reac-
tion (2) after the cementite decomposed according to re-
action:
Fe3C = 3Fe + Cmalleablized graphite (3)
The decarburization process depends on the chemical
composition of the decarburized material (steel or MCI),
on the applied atmosphere,10,11 the temperature and first
of all on the wall thickness of the casting. For MCI it
was also essential to know how the microstructure was
influenced by the wall thickness.12 The relationship be-
tween the mass fraction of carbon and the decarburi-
zation time is presented in Figure 2. The decarburization
was achieved by annealing in the temperature range from
1070 °C to 1075 oC in a gas atmosphere with water vapor
(H2O(g)), and the process consisted of carbon diffusion
from the interior towards the surface of the MCI casting,
of the water vapor transport to the surface of the MCI
casting, the oxidation of MCI carbon on the surface of
the MCI casting where proceeded and also the oxidation
of iron and of the other elements. The relations between
the decarburization time and the remaining carbon con-
centration in the MCI are represented by the following
equations:12
lg C = k – m·t (4)
C = 10k – m·t (5)
t
k C
m
=
− lg
(6)
with:
C – remaining carbon concentration,
t – time,
k, m – constants
The microstructural characteristics and the physical
properties of the malleable cast iron changed during the
decarburization process. Due to changed carbon concen-
trations and changed macro- and microstructures the
electrical resistivity also varied. The relationship be-
tween the electrical resistivity and the specific electrical
resistivity applied for the calculations of the specific
electrical resistivity during the MCI decarburization pro-
cess are described with the equation:
= R
A
l
( m) (7)
where:
 = specific electrical resistivity of the specimen ( m)
R = electrical resistivity of the specimen ()
A = cross-section of the specimen (mm2)
l = length (mm)
Matthiessen’s rule13 describes the relation between
the specific electrical resistivity and the temperature, as
follows:
  ( ) ( )T T= +0 G (8)
where 0 is a term that is independent of the temperature
and takes into account the influence of the alloying ele-
ments and G(T) is a temperature-dependent term.
M. PIRNAT et al.: A THERMODYNAMIC AND KINETIC STUDY OF THE SOLIDIFICATION AND DECARBURIZATION ...
530 Materiali in tehnologije / Materials and technology 45 (2011) 6, 529–535
Figure 2: Relationship between the mass fraction of carbon and the
decarburization time12
Slika 2: Odvisnost masnega dele`a ogljika od ~asa razoglji~enja12
Figure 1: Cooling curve and dilatometric curves of slightly hypo-
eutectic spheroidal graphite cast iron9
Slika 1: Ohlajevalna krivulja in dilatometrske krivulje podevtektske
sive litine s kroglastim grafitom9
2 EXPERIMENTAL
"In-situ" simple thermal and dilatometric analyses of
the same alloy were made in industrial conditions. The
"in-situ" measuring equipment is presented in Figure 3.14
A sample for chemical analysis was taken after each
measurement and for laboratory decarburization. For an
easier comparison, some of those specimens were
decarburized together with industrial castings in indus-
trial conditions, while the others were prepared only for
laboratory decarburization tests. The chemical analyses
of the as-cast cast irons and of the decarburized speci-
mens were made after "in-situ"dilatometrical measure-
ments, thermal analyses and differential scanning calo-
rimetry to follow how the electrical resistivity varied
during the laboratory decarburization process.The chem-
ical compositions of the as-cast cast iron samples were
evaluated and are presented in Tables 1 and 2.
The measurements of the electrical resistivity and of
the dimensional changes were performed with laboratory
equipment to follow the variations of the electrical resis-
tivity of the MCI specimens during the decarburization
process at 1050 °C, 1075 °C and 1100 °C for periods of
(12, 24 and 48) h in atmospheres of argon, CO2, N2 and
N2 + H2O. Figure 4 shows the used equipment.
The Olympus BX 1 optical microscope with the DP
70 video camera and the analySIS 5.0 software for ana-
lyzing micrographs was used in the metallographic ex-
aminations. Multiple image alignment (MIA), a method
of stitched overview, was applied to determine the
microstructural constituents and the variation of their
fractions with the distance from the surface of the speci-
mens. The microstructural components were determined
across the specimen’s cross-section from the center (x =
0) towards the edge (x = 2 500 μm). The distance from
the center to the edge of the specimens decarburized in
laboratory conditions was divided into five sections, each
was 500 μm long. The microstructural changes and the
variations in the microstructural constituents were deter-
mined in each section. Afterwards, the single pictures
were stitched into a joint picture.
3 RESULTS AND DISCUSSION
The simple thermal analysis was applied to record the
cooling curves and the curves of the dimensional
changes of specimens Nos. 1, 2, 3 and 4. In addition,
characteristic temperatures of the solidification and of
the transformations were determined, too. The cooling
curve with the marked characteristic temperatures for the
M. PIRNAT et al.: A THERMODYNAMIC AND KINETIC STUDY OF THE SOLIDIFICATION AND DECARBURIZATION ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 529–535 531
Figure 4: Laboratory equipment for measuring electrical resistivity,
and performing dilatometric analyses during the decarburization pro-
cess
Slika 4: Posnetek naprave za izvedbo meritev elektri~ne upornosti in
dilatometrijske analize pri razoglji~enju
Table 1: Chemical compositions of as-cast cast iron samples, including the C+Si sum and the Mn/S ratio w/%
Tabela 1: Kemijska sestava vzorcev 1, 2, 3, 4, vsota C + Si in Mn/S v masnih dele`ih
Chemical composition in mass fractions, w/%
Sample C Si Mn S Al P Cr C+Si Mn/S
1. 2.8405 0.9174 0.5026 0.1952 0.0010 0.0321 0.0395 3.7579 2.5748
2. 3.0633 0.9549 0.5005 0.1879 0.0024 0.0321 0.0396 4.0182 2.6637
3. 3.0032 0.9305 0.4991 0.2132 0.0009 0.0323 0.0398 3.9337 2.3410
4. 3.0259 0.9368 0.5045 0.1818 0.0004 0.0312 0.0040 3.9627 2.7750
Table 2: Chemical compositions of as-cast cast iron samples prepared for laboratory decarburization, including the C+Si sum and the Mn/S ratio
Tabela 2: Kemijska sestava vzorcev belega litega `eleza za razoglji~enje z vsoto C + Si in razmerjem Mn/S v masnih dele`ih, w%
Chemical composition in mass fractions, w/%
C Si Mn S Al P Cr C+Si Mn/S
2.99 0.85 0.49 0.17 0.004 0.0321 0.04 3.84 2.88
Figure 3: Equipment for "in situ" simple thermal and dilatometric
analyses with the measuring cell14
Slika 3: Naprava za enostavno termi~no in dilatometrijsko analizo z
merilno celico14 "in situ"
specimen No. 3 is presented in Figure 5. It shows that
the solidification started at 1280 °C with the precipita-
tion of primary austenite (L  Fe) and continued until
the temperature of the eutectic reaction (L  (Fe +
Fe3C)) at 1154 °C. After the eutectic reaction was com-
pleted, the cooling continued in the solid state down to
the eutectoid reaction (Fe  (Fe + Fe3C)) at 740 °C. Af-
terwards, the cooling continued and no further changes
were detected on the cooling curve. The liquidus and sol-
idus temperatures of all the specimens, i.e., of Nos. 1, 2,
3 and 4, were collected and are presented in Table 3 with
the results of the differential scanning calorimetry. Next
to characteristic temperatures, the reaction enthalpies
were determined also.
Figure 6a presents the heating curves of industrially
cast specimens, i.e., of the initial specimen, and of the
decarburized specimen with 0.18 % C. Both curves ex-
hibit the exothermic and endothermic peaks of the
eutectoid transformation (Fe  Fe + Fe3C); (Fe + Fe3C
 Fe) at 757 °C and 752 °C. Both peaks were much
more pronounced with the initial specimen since the
enthalpy of transformation was –16.48 J/g, while the
enthalpy of transformation in the decarburized specimen
was much smaller, only –3.632 J/g. The melting of the
initial specimen (blue line) commenced at 1139 °C with
the solidification of the (Fe + Fe3C  L) eutectic, and it
continued with the melting of the primary austenite (Fe
 L) at 1188 °C.
The melting enthalpy was at –1 34.2 J/g. In the
decarburized specimen (red line) no melting of the
eutectic (Fe + Fe3C  L) was detected. After the
eutectoid transformation (Fe + Fe3C  Fe) at 752 °C,
the heating was continued until the primary austenite
melted (Fe  L) at 1 331 °C, then the peritectic reaction
(Fe + L  Fe) proceeded at 1 452 °C and the -ferrite
melted (Fe  L) at 1 479 °C.
Figure 6b presents the cooling of the initial (blue
line) and of the decarburized specimen (red line). The
solidification of the decarburized specimen started at
1500 °C with precipitation of the -ferrite (L  Fe). The
peritectic reaction (Fe + L Fe) at 1467 °C, at 1366 °C
the peak of formation of primary austenite (L  Fe), at
942 °C the peak of formation of hypoeutectoid ferrite
from austenite (Fe  Fe) and at 769 °C a smaller peak
of eutectoid transformation (Fe  Fe + Fe3C) were de-
tected. The melting enthalpy was 19.68 J/g, and of
eutectoid transformation 7.771 J/g.
The solidification of the basic specimen started at
1290 °C with the precipitation of the primary austenite
(L  Fe) and it was completed at 1140 °C with eutectic
reaction or the solidification of the eutectic (L  Fe +
Fe3C), respectively. At 723 °C a big exothermic peak of
eutectoid transformation (Fe  Fe + Fe3C) appeared.
The enthalpy of solidification was 125.9 J/g, and of
eutectoid transformation was 76.09 J/g.
Figure 7 shows the results of measurements of the
electrical resistivity and of the dimensional changes dur-
ing decarburization of specimen No. 8. The decarburi-
zation proceeded at 1100 °C for 12 hours in the N2+H2O
atmosphere. The obtained value of the electrical resistiv-
ity after heating to the decarburization temperature of
1100 °C was 0.2725  and it was constantly dropping as
the decarburization continued. After 40 000 s of decar-
burization, the electrical resistivity dropped to 0.2525 .
M. PIRNAT et al.: A THERMODYNAMIC AND KINETIC STUDY OF THE SOLIDIFICATION AND DECARBURIZATION ...
532 Materiali in tehnologije / Materials and technology 45 (2011) 6, 529–535
Figure 6: Heating (a) and cooling curves (b) of the as-cast specimen
(blue), and specimen with mass fraction of C 0.18 %, decarburized in
industrial conditions (red), obtained by differential scanning calorime-
try
Slika 6: a) Segrevalni in b) ohlajevalni krivulji diferen~ne vrsti~ne
kalorimetrije; modro: izhodno lito stanje; rde~e: industrijsko razoglji-
~eni vzorec z masnim dele`em C 0,18 %
Figure 5: Dilatometric and cooling curve and derivative of the cooling
curve with marked temperatures of the solidification start, of the
eutectic solidification and of the eutectoid transformation for sample
No. 3
Slika 5: Dilatometrijska in ohlajevalna krivulja z odvodom ohlaje-
valne krivulje s temperaturami za~etka in evtekti~nega strjevanja ter
evtektoidne premene za vzorec 3
With a lowering of the electrical resistivity dimensional
changes occurred. The dimensional changes dropped
from 400 μm at the beginning of the decarburization pro-
cess to 120 μm at the end of the process.
Figure 8 presents the changes of the microstructure
and of the fractions of the microstructural constituents as
a function of the distance x from the center of the speci-
men to its edge. The stitched metallographic image pres-
ents the microstructures and the fractions of the
microstructural constituents, i.e., pearlite, ferrite, and
graphite. The fractions of microstructural constituents
are also presented graphically as a function of the dis-
tance x from the specimen center to its edge. The frac-
tion of microstructural constituents changed from the
center to the edge and greater changes of the
microstructural constituents were detected at adistance of
1 000 μm to 1 500 μm from the center. The fraction of
graphite was reduced to 1 % and the fraction of pearlite
to 70 %. In contrast, the fraction of ferrite was constantly
increasing and the share of ferrite was 25 % at adistance
of 1 500 μm, while its share at the edge of the specimen
reached as high as 90 %.
Based on measurements of the electrical resistivity
and the changed lengths of specimens during the
decarburization of white-heart cast iron and applying a
physico-mathematical model of white-heart cast iron
decarburization, variations of the carbon concentrations
and the specific electrical resistivity during the decarbu-
rization process were evaluated as a function of the
decarburization time. The variations of the carbon con-
centrations were calculated for (12, 24, 36, 48 and 60) h
of decarburization at a temperature T = 1 000 °C for
specimens that were decarburized in various atmo-
spheres. These relations are presented in Figure 9. The
relations between the decarburization time and the
depths of the decarburized zone are shown in Figure 10.
It is evident from the plot in Figure 9 that the greatest
variations of the specific electrical resistivity and of the
carbon concentration occurred between 12 h and 24 h of
decarburization. A similar behavior was also found with
measurements of the electrical resistivity of laboratory
specimens, where the corresponding time interval was
between 11.1 h and 22.2 h. The course of the decarbu-
rization process in Figure 9, i.e., the relation between
the carbon concentrations and the time of decarburi-
M. PIRNAT et al.: A THERMODYNAMIC AND KINETIC STUDY OF THE SOLIDIFICATION AND DECARBURIZATION ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 529–535 533
Figure 8: Micrograph with microstructural constituents in the speci-
men (T3 t1);T = 1 100 °C; 12 h; N2 + H2O mixture
Slika 8: Posnetek mikrostrukture in dele`i mikrostrukturnih sestavin
vzorca (T3 t1) ; T = 1 100 °C; 12 h; N2 + H2O
Figure 7: Variations of electrical resistivity and dimensional changes
with time during decarburization of the specimen, decarburized in lab-
oratory conditions (T3 t1); T = 1 100 °C, 12 h, N2 + H2O mixture,
Slika 7: Elektri~na upornost in dimenzijske spremembe v odvisnosti
od ~asa razoglji~enja za vzorec (T3 t1); T = 1 100 °C;12 h; N2 + H2O
Table 3: Characteristic temperatures obtained with simple thermal analysis (TA) and differential scanning calorimetry (DSC) in °C
Tabela 3: Zna~ilne temperature TA in DSC v stopinjah Celzija
Temperatures of solidification and phase transformations in solid state
TA DSC
Solidification and phase
transformations in solid state Specimen
Decarburized speci-
men in industrial con-
ditions, containing
mass fraction of C
0.18 %
As-cast specimen
1 2 3 4
L  Fe 1 500
Fe + L   1 467
L  Fe 1 280 1 280 1 281 1 366 1 290
L  (Fe + Fe3C) 1 158 1 154 1 154 1 154 1 140
Fe  Fe 942
Fe  Fe + Fe3C) 740 738 740 733 769 723
zation, was similar to that described where the
decarburization of the electrical sheet was investigated15.
Furthermore, Figure 10 shows that an increased
decarburization rate in the time interval between 12 h
and 24 h of decarburization resulted in greater depths of
the decarburized zone that varied between 1 mm and 6
mm, depending on the decarburization time and the
decarburization conditions. The relationships between
the wall thickness of the casting and the decarburization
time at T = 1 000 °C are presented, showing that a depth
of the decarburization zone of 5 mm was reached after
50 h of decarburization.12
The temperature of the solidification and of the phase
transformations were determined by analyzing the course
of the MCI solidification and decarburization by phy-
sico-metallurgical means. Among those temperatures,
the essential in the MCI decarburization process is the
temperature of eutectoid transformation (Fe  Fe +
Fe3C) at 752 °C.
The mechanism of the MCI decarburization process
was determined. This process started at T  752 °C by
carbon loss in the austenite, [C]-Fe, and was continued by
Cmalleablized graphite loss, described with the reactions:
[C]-Fe + H2O(g) = CO(g) + H2(g), and
Cmalleablized graphite + H2O(g) = CO(g) + H2(g)
4 CONCLUSIONS
In the first part of the research, the MCI decarburi-
zation was investigated with the thermal analyses, chem-
ical analyses, and differential scanning calorimetric anal-
yses of as-cast malleable samples. The thermal analyses
revealed the entire course of the solidification process,
the course of the cooling, and reactions relating to how
the austenite, cementite eutectic and pearlite were
formed. Differential scanning calorimetry of the as-cast
specimens confirmed the course of the reactions that
were determined by the thermal analysis.
Further examinations were focused on a laboratory
examination of the malleablizing process from the ther-
modynamic and kinetic points of view and the dilato-
metric analyses of specimens during the malleablizing
process were performed. The measurements of the elec-
trical resistivity were added to follow the decarburization
process more precisely. Electrical resistivity changes
during the decarburization process of the specimens
decarburized in laboratory conditions were confirmed by
an assessment of the changes in the microstructures.
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534 Materiali in tehnologije / Materials and technology 45 (2011) 6, 529–535
Figure 10: Variation of the depth of the decarburization zone during
the decarburization process at T = 1 100 °C
Slika 10: Sprememba globine razoglji~enja v odvisnosti od ~asa
razoglji~enja pri temperaturi T = 1 100 °C
Figure 9: Variation of the carbon concentration and of the specific
electrical resistivity during the decarburization process at T = 1 100 °C
Slika 9: Sprememba koncentracije ogljika in specifi~ne elektri~ne
upornosti v odvisnosti od ~asa razoglji~enja pri temperaturi T = 1 100
°C
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M. PIRNAT et al.: A THERMODYNAMIC AND KINETIC STUDY OF THE SOLIDIFICATION AND DECARBURIZATION ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 529–535 535

V. KEVORKIJAN et al.: MODELLING AND PREPARATION OF CORE FOAMED Al PANELS ...
MODELLING AND PREPARATION OF CORE FOAMED
Al PANELS WITH ACCUMULATIVE HOT-ROLL
BONDED PRECURSORS
NA^RTOVANJE IN IZDELAVA Al-PANELOV S SREDICO IZ
Al-PEN NA OSNOVI VE^STOPENJSKO TOPLO VALJANIH
PREKURZORJEV
Varu`an Kevorkijan1, Uro{ Kova~ec2, Irena Paulin3, Sre~o Davor [kapin4,
Monika Jenko3
1Independent Researcher, Betnavska cesta 6, 200 Maribor, Slovenia
2Impol Aluminium Industry, Partizanska 38, 2310 Slovenska Bistrica, Slovenia
3Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
4Institut "Jo`ef Stefan", Jamova 39, 1000 Ljubljana, Slovenia
varuzan.kevorkijan@impol.si
Prejem rokopisa – received: 2011-08-29; sprejem za objavo – accepted for publication: 2011-10-25
In this paper, laboratory and semi-industrial processes for the preparation of aluminium foam samples and core-foamed panels
with closed porosity were investigated. The samples were prepared starting from the accumulative hot-roll bonded precursors,
with titanium hydride (TiH2) or dolomite (Ca0.5Mg0.5CO3) powder added as the foaming agent. The formation of the precursors
was performed in three steps. In the initial stage, titanium dihydride or dolomite particles were deposited on a single side of a
selected number of aluminium strip samples made from the alloy AA 1050. In the second step, by putting together in pairs,
single-sided coated strips, precursors with a two-layered structure were prepared. The samples were hot-rolled to a final
thickness of 1.9–3.8 mm, introducing a total deformation of about 45–49 % by a process well-known as accumulative hot-roll
bonding. In the third stage of the precursor’s formation, the desired multilayered precursor’s structure was achieved by hot-roll
multi-passing, i.e., by repeating (with 2–16 passes) the accumulative hot-roll bonding procedure. The obtained precursors were
foamed in an electrical furnace, under different foaming conditions, based on the initial temperature of the thermal
decomposition of the foaming agent. The microstructure of the obtained foam samples was investigated with optical and
scanning electron microscopy. According to the accumulated experimental results, one can conclude that the usage of dolomite
powder as a foaming agent with a higher temperature of thermal decomposition (>750 °C) compared to TiH2, which thermally
decomposed even at the temperature of hot-rolling (>350 °C), enabling the formation of multilayered precursors at higher
temperatures of hot-rolling without any intermediate annealing. This consequently increases the productivity of the foamed core
panel production without influencing their final quality.
Key words: Al foams, core foamed Al panels, precursors preparation, accumulative hot-roll bonding, comparison of different
foaming agents, characterisation
V delu opisujemo razvoj laboratorijskih in polindustrijskih postopkov priprave vzorcev aluminijskih pen in panelov iz Al-pen z
zaprto poroznostjo. Vzorce panelov smo izdelovali na osnovi ve~stopenjsko toplo valjanih prekurzorjev, ki so kot sredstvo za
penjenje vsebovali delce titanovega dihidrida (TiH2) ali dolomitnega prahu (Ca0,5Mg0,5CO3). Postopek priprave ve~plastnih
prekurzorjev je potekal v treh fazah. V za~etni fazi smo delce titanovega dihidrida ali dolomita nana{ali na izbrano stran
aluminijevega traku zlitine AA 1050. Sledila je priprava
dvoplastnega prekurzorja. Po dva in dva premazana trakova smo zlo`ili tako, da sta se premazani stranici stikali ter dvoj~ek
vro~e zvaljali na 1,9–3,8 mm s skupno deformacijo med 45–49 %. Postopek smo v sklepni fazi izdelave prekurzorjev ponavljali
do `elene ve~plastnosti (od 2- do 16-krat ). S postopkom ve~kratnega valjanja in podvajanja plasti penilnega sredstva smo
dosegli enakomerno porazdelitev penilnega sredstva skozi celoten prerez izdelanih prekurzorjev. Dobljene prekurzorje smo nato
penili v elektri~ni pe~i pri razli~nih pogojih glede na temperaturo termi~nega razkroja sredstva za penjenje. Mikrostrukturo
dobljenih vzorcev pen smo analizirali z opti~no in vrsti~no elektronsko mikroskopijo. Z raziskavami smo potrdili, da uporaba
dolomitnega prahu kot penilnega sredstva z vi{jo temperaturo termi~nega razkroja (>750 °C) v primerjavi s TiH2, ki se termi~no
razkraja `e pri temperaturi toplega valjanja (>350 °C), omogo~a izdelavo ve~plastnih prekurzorjev pri vi{jih temperaturah
toplega valjanja brez vmesnega `arjenja in posledi~no povi{uje produktivnosti brez vpliva na kakovost kon~nega izdelka.
Klju~ne beside: Al-pene, Al-paneli s penasto sredico, priprava prekurzorjev, ve~stopenjsko toplo valjanje, primerjava razli~nih
sredstev za penjenje, karakterizacija
Materiali in tehnologije / Materials and technology 45 (2011) 6, 537–544 537
UDK 669.715.017:621.771 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)537(2011)
1 INTRODUCTION
Accumulative Roll Bonding (ARB) is a new and
promising manufacturing process for closed-cell alu-
minium foams and core-foamed panels 1–3. The ARB
process, as one of the severe plastic deformation pro-
cesses, has been applied to many metals and alloys to
produce an ultra-fine crystal grains structure. The pro-
cessing procedure is presented in Figure 1.
Instead of aluminium powder, Al sheet is applied as a
starting material, making this processing route cost-ef-
fective and promising for further industrialization. Dur-
ing the first stage, two aluminium strips are stacked with
the appropriate amount of selected blowing-agent pow-
der of the proper morphology. Second, the strip is
roll-bonded by a 50 % reduction and cut into two. After
repeatedly covering the surface with blowing agent, the
two strips are stacked to be the initial dimension, and
then roll-bonded again. Since these procedures can be re-
peated indefinitely, the desired multilayered precursor’s
structure could be achieved. By multiplying the foam-
ing-agent layers, its uniform distribution across the entire
cross-section of the prepared precursor can be achieved.
Finally, the obtained precursors are foamed to the end
products: foamed aluminium sheet or core-foamed pan-
els.
Investigations performed in the past few years have
confirmed the great potential of the ARB processing
route in the cost-effective and highly productive fabrica-
tion of foamed aluminium sheets and various
core-foamed panels by TiH2 as a foaming agent 3, 4. An
additional reduction in cost can be achieved by replacing
the expensive TiH2 with alternative, inexpensive foaming
agents, particularly carbonates, among which natural car-
bonates such as CaCO3 marble powder or CaMg(CO3)2
dolomite powder are particularly attractive.
Successfully replacing the TiH2 with carbonates in
the foaming precursor fabricated by the powder metallur-
gical or stir-casting route has been reported by several
authors 5–9. However, similar investigations of replacing
the TiH2 foaming agent with dolomite particles in accu-
mulative hot-roll bonded precursors are still not com-
pleted. Therefore, the purpose of this work was to inves-
tigate the possible benefits of replacing TiH2 with
dolomite particles in multilayered precursors made by
ARB.
2 EXPERIMENTAL PROCEDURE
The samples were prepared by starting from the accu-
mulative hot-roll bonded precursors, with titanium hy-
dride (TiH2) or dolomite (Ca0,5Mg0,5CO3) powder added
as the foaming agent.
The formation of the precursors was performed in
three steps. In the initial stage, titanium dihydride or do-
lomite particles were deposited on a single side of a se-
lected number of aluminium strip samples (100 mm in
width, 200 mm in height and with a thickness of approx.
2–4 mm) made from the alloy EN AW 1050. The suspen-
sion of the powdered foaming agent (TiH2 or MgAl3O4)
in acetone was spread out on the strip with a painter’s
roll. Titanium hydride (supplier: AG Materials Inc.) and
dolomite powders (supplier: Granit, d. o. o., Slovenska
Bistrica, Slovenia) of five different average particle sizes
were applied as foaming agents. The average particle
size of the powders used in the experiments is listed in
Table 1. The particle size distribution of the a powdered
foaming agents was measured using laser particle ana-
lyzer (Malvern Mastersizer 2000). The relative error of
the measurement was within ±1 %.
The surface concentration of the foaming agent
achieved after removing the acetone was between ap-
proximately 2.5 × 10–3 and 3.75 × 10–3 mg/mm2. In the
second step, by putting together in pairs single-sided
coated strips, precursors with a two-layered structure
were prepared. The samples were hot-rolled to a final
thickness of 1.9–3.8 mm, introducing a total deformation
of about 45–49 % by a process that is well known as ac-
cumulative hot-roll bonding. It was found that the con-
sistency of the obtained precursors (the adhesion be-
tween two layers) was strongly affected by the
parameters of the hot-rolling (the strip deformation and
the temperature of the hot-rolling). These were signifi-
V. KEVORKIJAN et al.: MODELLING AND PREPARATION OF CORE FOAMED Al PANELS ...
538 Materiali in tehnologije / Materials and technology 45 (2011) 6, 537–544
Figure 1: (a) Schematic illustration of the manufacturing process of a
precursor sheet using the ARB process. (b) Prediction of gradual dis-
tribution of added foaming-agent particles 3.
Slika 1: (a) Shematski prikaz postopka ve~stopenjskega toplega valja-
nja prekurzorja; (b) prikaz porazdelitve delcev penila v ve~plastnem
prekurzorju 3.
cantly limited by the technical possibilities of the exist-
ing mini hot-rolling mill. The diameter of the working
rolls was only 200–450 mm. However, by carefully se-
lecting the optimal power of the hot-rolling mill unit
(about 18–50 kW) for the proper deformation of the strip
and the temperature of the hot-rolling (between 300–350
°C), the complete recrystallization of the strip was
achieved, avoiding in that way any possible hardening of
the wrought aluminium alloy. In the third stage of the
precursor’s formation, the desired multilayered precur-
sor’s structure was achieved by hot-roll multi-passing,
i.e., by repeating (with 2–16 passes) the accumulative
hot-roll bonding procedure. By the multiple accumula-
tive hot-roll bonding and by doubling the layers of the
foaming agent, its uniform distribution was achieved
across the entire cross-section of the prepared precursors.
The precursors were foamed in a conventional batch
electrical furnace with air-atmosphere circulation under
various experimental conditions (time, temperature) and
by applying the same cooling method. Before foaming,
the individual precursors were placed on a ceramic plate
covered by a boron nitride layer. The plate dimensions
and the precursor size 100 mm × 100 mm were selected
to allow the complete expansion of the precursor to
foam. The arrangement was placed inside a pre-heated
batch furnace at a selected temperature and held for the
selected holding time. In samples with TiH2, thermal de-
composition was observed, starting from approximately
350 °C. These samples were foamed at 700–750 °C, for
3–18 s. Samples with the dolomite foaming agent, for
which the thermal decomposition is initiated above 750
°C, were foamed at 780 °C for 2–5 min. After that, the
sample was removed from the furnace and the foaming
process was stopped by rapid cooling with pressurised
air to room temperature. The thermal history of the foam
sample was recorded, using a thermocouple located di-
rectly in the precursor material. The density of the foam
was calculated using Archimedes’ method. The porosity
of the manufactured foam was calculated using the rate:
1-(foam density/aluminium density). Macro and micro-
structural examinations were performed on sections ob-
tained by wire precision cutting across the samples and
on samples mounted in epoxy resin, using optical and
scanning electron microscopy (SEM/EDS). The average
particle size of the pores in the foams was estimated by
analysing optical and scanning electron micrographs of
the as-polished foam bars using the point-counting
method and image analysis and processing software.
3 MODELLING OF THE SURFACE
CONCENTRATION AND MORPHOLOGY OF
THE FOAMING AGENT IN ARB PRECURSORS
In precursors fabricated by the ARB processing
route, the proper surface concentration and morphology
of the foaming agent are crucial for the development of
high-quality foams with closed cells and a uniform
microstructure. Because of that, it was very important to
correlate by the model developed in this work the surface
concentration and morphology of the foaming agent with
an average cell size and microstructure of the foam.
The particles of foaming agent are, as much as possi-
ble, homogeneously dispersed on the surface of the alu-
minium sheet, as is evident in Figure 2.
V. KEVORKIJAN et al.: MODELLING AND PREPARATION OF CORE FOAMED Al PANELS ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 537–544 539
Figure 2: SEM micrograph of the surface of an aluminium sheet after
the first cycle of ARB, with homogeneously distributed dolomite par-
ticles
Slika 2: SEM-posnetek povr{ine aluminijskega traku po prvem
prehodu ve~stopenjskega toplega valjanja s homogeno porazdeljenimi
delci dolomita
Table 1: The average particle size and the cumulative particle size dis-
tribution of the TiH2 and dolomite powder applied as foaming agents
Tabela 1: Povpre~na velikost delcev in porazdelitev delcev po
velikosti TiH2 in dolomitnega prahu, uporabljenih kot sredstvo za
penjenje
TiH2 powder TIH-0420
Average particle size (μm) 20.4
Cumulative particle size distribution (μm
D10 13.1
D25 17.4
D50 20.4
D75 23.7
D90 41.4
Uniformity of particle size distribution (μm)
D90–D10 28.3
Dolomite powder D-4
Average particle size (μm) 20.8
Particle size distribution
D10 11.2
D25 15.9
D50 20.8
D75 23.1
D90 27.2
Uniformity of particle size distribution (μm)
D90–D10 16.0
The neighbouring particles of foaming agent are mu-
tually apart for an average minimum distance Xmin. As
will be demonstrated in the model, this distance is deter-
mined by the surface concentration of the foaming agent
C and its average particle size d50.
The surface concentration of the foaming agent is de-
fined as follows:
C = m/S (1)
Here, m is the total mass of the foaming agent and S
is the surface of the aluminium sheet. The molar surface
concentration C’ is defined as:
C’= m/(MS) (2)
where M represents the molar mass of the selected
foaming agent. Finally, the particle surface concentration
C’’ is introduced as:
C’’ = N/S = 1/Xmin
2 (3)
N is the total number of particles of foaming agent:
N = mp/M = d50
3/6M (4)
In Eq. 4, mp represents the mass of an average particle
of foaming agent, d50 is the average particle size and  is
the theoretical density of the foaming agent.
Therefore, by combing Eqs.1–4 we can derive, for
the particle surface concentration, the following expres-
sion:
C’’ = 6C/(d50
3) (5)
and, consequently, for Xmin:
Xmin = [(d50
3 )/6C]1/2 (6)
Applying, for the bubbles’ growth, a simple,
stoichiometric model, in which the complete thermal de-
composition of an individual foaming agent particle pro-
vides a gas phase for the bubble’s nucleation and growth,
the maximum bubble diameter Dmax can be determined
with the ideal gas equation:
pmaxV = nRT (7)
where n corresponds to the number of moles of gas
phase inside the bubble, V is the bubble volume, R is the
universal gas constant and T is the temperature.
pmax can be calculated by applying Laplace’s equa-
tion:
pmax = (2lg/r) + gh + p0 (8)
The maximum bubble pressure pmax is the sum of the
capillary (2lg/r), hydrostatic (gh) and atmospheric (p0)
pressures. The capillary pressure depends on the surface
tension lg at the gas-liquid interface and the bubble ra-
dius (r); the hydrostatic pressure is determined by the
immersion depth (h) and the density of the molten alu-
minium alloy ().
By considering a spherical bubble and by combining
Eqs. (7) and (8), we can calculate:
[(2lg/r) + gh + p0] (4/3)rmax
3 = nRT (9)
During an early stage of the bubble’s growth, only
the capillary pressure (2lg/r) should be considered,
while in the final stage of the bubble’s growth the only
important pressure is the atmospheric (p0). Note that for
laboratory conditions, the hydrostatic pressure (gh) is
always negligible.
Based on this, in the case considered by the model:
pmax = p0 (10)
Therefore, the maximum bubble diameter Dmax is fi-
nally determined by the formula:
Dmax = d50[(kRT)/p0M]
1/3 (11)
Here, k represents the stoichiometric constant (k = 1
for TiH2 and k = 2 for dolomite), R is the universal gas
constant, T is temperature,  is the density of the foam-
ing agent, p0 is the atmospheric pressure and M is the
molar mass of the foaming agent. In Tables 2 and 3, the
calculated values for the maximum bubble radius for
cells grown from the TiH2 and dolomite particles of dif-
ferent initial particle size are listed.
Table 2: Maximum bubble radius for a cell grown from TiH2 particles
with different initial particle sizes (d50).
Tabela 2: Maksimalni premer pore iz TiH2 delcev razli~ne za~etne
velikosti (d50)
The average particle
size of TiH2 (μm)
3 20 40 75 140
Maximum bubble diam-
eter, Dmax /μm
54 360 808 1.352 2.526
Table 3: Maximum bubble radius for bubbles created by dolomite par-
ticles of different initial particle sizes (d50)
Tabela 3: Maksimalni premer pore iz delcev dolomita razli~ne
za~etne velikosti (d50)
The average particle
size of dolomite (μm) 3 5 10 20 35
Maximum bubble diam-
eter, Dmax/μm
40 68 132 268 468
According to the theoretical prediction based on Eq.
11 and the calculated values reported in Tables 2 and 3,
at the same initial particle size of the applied foaming
agents (3 μm or 20 μm) and the same foaming conditions
(temperature, time), the bubbles created by TiH2 should
be coarser.
In order to obtain a stable foam microstructure, with
isolated closed cells, it is necessary that:
Xmin > D50 (12)
By combining Eqs. 6, 11 and 12, the condition (12)
leads to the required correlation between the surface con-
centration and the foaming-agent morphology:
C < const. (d50/T)
2/3 (13)
Based on the model, it is evident that for the selected
morphology of the foaming agent d50 and foaming tem-
perature T the surface concentration of the foaming agent
C cannot be selected arbitrarily, but, in order to achieve a
V. KEVORKIJAN et al.: MODELLING AND PREPARATION OF CORE FOAMED Al PANELS ...
540 Materiali in tehnologije / Materials and technology 45 (2011) 6, 537–544
foam microstructure with stable individual closed cells,
it should fulfil the condition expressed by Eq.13.
4 RESULTS AND DISCUSSION
4.1 Foam properties as a function of the composition
of accumulative roll-bonded precursors
The results of the investigation of the foam properties
(density and cell size distribution) achieved by applying
various types (TiH2 or dolomite) and surface concentra-
tions of foaming agents are reported in Tables 4 and 5.
Table 4: The experimentally measured density and cell-size distribu-
tion of the aluminium foam samples at various concentrations of TiH2
foaming agent. The foaming conditions: 700 °C, 120 s.
Tabela 4: Eksperimentalno izmerjene vrednosti gostote in poraz-
delitve velikosti por v vzorcih aluminijske pene, izdelanih pri razli~ni
koncentraciji delcev TiH2, uporabljenega kot penila. Pogoji penjenja:
700 °C, 120 s.
Surface concentrations
of TiH2 (mg/mm2)
2.5 × 10–3 3.0 × 10–3 3.5 × 10–3
Density of Al foam
(% T. D.) 25.9 ± 1.3 23.8 ± 1.2 21.2 ± 1.1
Cell size distribution (mm)
D10 2.6 ± 0.3 3.2 ± 0.3 3.3 ± 0.3
D25 2.7 ± 0.3 3.5 ± 0.3 3.8 ± 0.3
D50 2.8 ± 0.3 3.7 ± 0.4 5.1 ± 0.5
D75 3.0 ± 0.3 4.2 ± 0.3 5.5 ± 0.6
D90 4.6 ± 0.5 8.4 ± 0.3 10.4 ± 1.0
Uniformity of cell size distribution (μm)
D90–D10 2.0 ± 0.2 5.2 ± 0.6 7.1 ± 0.7
Table 5: The experimentally measured density and cell-size distribu-
tion of aluminium foam samples at various concentrations of dolomite
foaming agent. The foaming conditions: 700 °C, 120 s.
Tabela 5: Eksperimentalno izmerjene vrednosti gostote in poraz-
delitve velikosti por v vzorcih aluminjske pene, izdelanih pri razli~ni
koncentraciji delcev dolomite, uporabljenega kot penila. Pogoji
penjenja: 700 °C, 120 s.
Surface concentrations
of dolomite (w/%) 2.5 × 10
–3 3.0 × 10–3 3.5 × 10–3
Density of Al foam
(% T.D.) 15.4 ± 0.8 12.8 ± 0.6 11.1 ± 0.6
Cell size distribution (mm)
D10 2.0 ± 0.2 2.4 ± 0.2 2.9 ± 0.3
D25 2.3 ± 0.2 2.9 ± 0.3 3.3 ± 0.3
D50 2.5 ± 0.3 3.2 ± 0.3 4.7 ± 0.5
D75 2.7 ± 0.3 4.8 ± 0.6 5.1 ± 0.5
D90 5.8 ± 0.6 7.6 ± 0.8 9.7 ± 1.0
Uniformity of cell size distribution (μm)
D90-D10 3.8 ± 0.4 5.2 ± 0.5 6.8 ± 0.7
Generally, the panels foamed with the dolomite
foaming agent were with a more uniform cell-size distri-
bution and lower average bubble size. The most uniform
cell-size distribution was achieved in foam samples
foamed with the minimum surface concentration (2.5 ×
10–3 mg/mm2) of dolomite powder, Figure 3.
The experimentally determined values of the average
bubble radius in the foamed panels were at least for one
order of magnitude higher that those predicted by the
model. The reason for that difference is in effects limit-
ing the stability of the individual bubbles, which are not
considered by the model. These effects are as follows:
bubble flow, drainage, rupture or coalescence, and coars-
ening.
From the difference between the theoretically pre-
dicted and experimentally determined values of the bub-
ble radius, it is possible to estimate the stability of the
real foam systems considered in this work. The experi-
mental findings clearly confirm that coarser bubbles are
more stable that finer ones. In addition, it is also evident
that the stability of the bubbles is much higher in foams
created by dolomite particles than in the counterparts
foamed by TiH2. However, in both cases the average bub-
ble sizes are proportional to the average initial size of the
foaming particles – finer foaming particles lead to finer
bubbles, while coarser ones lead to larger bubbles, as
was predicted by the model.
On the other hand, the density of the aluminium foam
samples was inversely proportioned to the bubble radius:
the foam samples with finer bubbles had the higher den-
sity and, in contrast, the foam samples with larger bub-
bles were specifically lighter. At the same time and un-
der the same foaming conditions (temperature, time), the
foams made with the dolomite were with a significantly
lower density than the samples with similar cell size
foamed by the TiH2.
As is evident from the cell-size distribution data
listed in Tables 3 and 4, an increase of the foam-
ing-agent surface concentration (either TiH2 or dolomite)
leads to the formation of foams with larger bubbles and a
lower density. However, also in that case, the samples
foamed with dolomite were with smaller bubbles and
lower densities.
V. KEVORKIJAN et al.: MODELLING AND PREPARATION OF CORE FOAMED Al PANELS ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 537–544 541
Figure 3: The homogeneous microstructure of the Al foam made from
the ARB precursor with dolomite particles as the foaming agent. The
ARB precursor was made by four roll-bonding cycles.
Slika 3: Homogena mikrostruktura vzorca Al-pene iz prekurzorja,
izdelanega po postopku ve~stopenjskega toplega valjanja
The experimentally developed foam microstructures
were mostly influenced by the degree of foam movement
slowing down (i.e., the foam stability) attained in partic-
ular trials. In addition, the microstructure of foams was
also influenced in some cases by the layered structure of
the ARB precursors.
The slowing down of the movement of the foam in-
cludes the prevention of flow (the movement of bubbles
with respect to each other caused either by external
forces or changes in the internal gas pressure during
foaming), drainage (flow of liquid metal through the
foam), coalescence (sudden instability in a bubble wall
leading to its disappearance) and coarsening (slow diffu-
sion of gas from smaller bubbles to bigger ones).
The layered structure of the ARB precursors caused,
in some samples, the appearance of so-called "line segre-
gation", i.e., the flow of liquid metal across the line, Fig-
ure 4. Such segregation is most probably involved in the
non-sufficient and/or different plastic deformation of in-
dividual layers during various roll-bonding cycles.
Generally, the foam stability in the liquid and semi-
solid state of an aluminium alloy is improved by increas-
ing the thickness of the bubble wall, decreasing the sur-
face tension of the molten metal and increasing its vis-
cosity. Simultaneously decreasing the surface tension
and increasing the viscosity of molten aluminium can be
achieved by introducing to the molten metal some
amount of ceramic particles, e.g., formed in situ, by the
thermal decomposition of the foaming agent at the inter-
face between molten aluminium and the gas phase.
Quantitatively, the foam stability in the liquid and
semi-solid state might be expressed by a dimensionless
foam-stability factor (FSF), which we defined as:
FSF = X0/(t) (14)
where X0 is the average distance between neighbouring
bubbles (which is proportional to the wall thickness), 
is the dynamic viscosity of the slurry,  is the surface
tension at the interface molten or semi-solid alu-
minium-gas and t is the foaming time.
Evidently, a higher FSF means better foam stability.
Unfortunately, in practice, it is not easy to determine the
viscosity and the surface tension of the slurry of the mol-
ten aluminium alloy and the foaming agent particulates
as well as the wall thickness of bubbles during their
growth. Because of that, the usage of FSF is mostly lim-
ited to theoretical and, to some level, qualitative consid-
erations.
The main differences between the TiH2 and the dolo-
mite foaming agent, which influences the foam micro-
structure development, are the following: (i) in the nature
and reactivity of products of its thermal decomposition
and (ii) the temperature interval of thermal decomposi-
tion.
V. KEVORKIJAN et al.: MODELLING AND PREPARATION OF CORE FOAMED Al PANELS ...
542 Materiali in tehnologije / Materials and technology 45 (2011) 6, 537–544
Figure 5: SEM micrograph of the internal surface of the bubble wall:
a) bubble wall fully covered by Al2O3, CaO and MgO particles in bub-
bles foamed by dolomite and b) clean internal surface in bubbles
foamed by TiH2
Slika 5: SEM-posnetek notranje povr{ine por: a) nastale z razkrojem
dolomita ter povsem prevle~ene z delci Al2O3, CaO in MgO in b)
neprevle~ene, nastale z razkrojem TiH2
Figure 4: The characteristic "line segregation" in the samples of alu-
minium foams made from ARB precursors using dolomite as a foam-
ing agent
Slika 4: Zna~ilno "linijsko izcejanje" v vzorcih aluminijske pene z
dolomitom kot sredstvom za penjenje. Vzorci pene so iz prekurzorjev,
izdelanih po postopku ve~kratnega toplega valjanja
In ARB precursors with TiH2, the thermal decompo-
sition of the foaming agent was observed starting from
approximately 350 °C. The TiH2 decomposes to Ti(s) and
H2(g), Eq. 15, which do not tends to react with molten
aluminium in the way of creating secondary ceramic
phases making it possible to slow down the movement of
the foam. Because of that, the internal surface of the
bubble wall foamed by the TiH2 is clean (Figure 5b).
TiH2(s) = Ti(s) + H2(g) (15)
In contrast to this, the thermal decomposition of the
dolomite foaming agent, Eq. (16), proceeded above 700
°C, resulting in highly reactive products: MgO(s), CaO(s)
and CO2(g) 7.
CaMg(CO3)2(s) = CaO(s) + MgO(s) + 2CO2(g) (16)
CaO and MgO are in the form of solid particulate ag-
gregates, appearing at the molten metal-CO2(g) interface,
while CO2(g) reacts with the molten aluminium in a bub-
ble covering bubble internal surface with a thin, mostly
continuous, film of Al2O3 (Figure 5a). Therefore, all the
products of the dolomite thermal decomposition (MgO,
CaO and Al2O3) are very effective in slowing down the
foam movement, decreasing the surface tension and in-
creasing the local viscosity of the slurry. Due to the lim-
ited foam movement, bubbles in solidified samples re-
main significantly finer, resulting in foams with a higher
density (a lower volume fraction of bubbles).
4.2 Fabrication of prototype core-foamed aluminium
panels
Accumulative hot-roll bonded precursors with dolo-
mite particles as a foaming agent were successfully
foamed to core-foamed panels with a wall thickness of
about 2 mm and a thickness of the foamed core of about
10 mm. The flat panel samples (approx. 150 mm long
and 80 mm wide) were routinely foamed from the ARB
precursors using the existing laboratory capabilities, Fig-
ure 6.
5 CONCLUSION
Our experimental findings confirmed that the ARB
procedure is very promising (semi)-industrial route for
the production of precursors for core-foamed aluminium
panels. By combining a pair of AA1050 aluminium alloy
strips and various foaming agents (TiH2 or dolomite
powders), several foaming precursors were prepared
through two to four cycles of ARB and successfully
foamed into core-foamed aluminium panels.
The use of dolomite powder as a foaming agent with
a higher temperature of thermal decomposition (>700
°C) compared to TiH2, which thermally decomposed
even at the temperature of hot-rolling (>350 °C), enabled
the formation of multilayered precursors at higher tem-
peratures of the hot-rolling without any intermediate an-
nealing. This significantly improved the productivity of
the production of core-foamed aluminium panels without
influencing their final quality.
The microstructure uniformity and stability of the
foams made from ARB precursors depend on the surface
concentration and morphology of the foaming agent as
well as on the foaming temperature.
Acknowledgement
This work was supported by funding from the Public
Agency for Research and Development of the Republic
of Slovenia, as well as the Impol Aluminium Company
and Bistral, d. o. o., from Slovenska Bistrica under con-
tract No. 2410-0206-09.
6 REFERENCES
1 N. Tsuji, Y. Saito, S. H. Lee, Y. Minamino, Adv. Eng. Mater., 5
(2003) 5, 338
2 Y. Saito, H. Utsumonia, N. Tsuji, T. Sakai, Acta Mater., 47 (1999),
579
3 K. Kitazono, E. Sato, K. Kuribayashi, Scripta Mater., 50 (2004), 495
V. KEVORKIJAN et al.: MODELLING AND PREPARATION OF CORE FOAMED Al PANELS ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 537–544 543
Figure 6: The cross-section of the prototype core-foamed panel made
from ARB precursor with dolomite particles as the foaming agent.
Slika 6: Pre~ni prerez prototipnega panela s sredico iz aluminijske
pene. Panel je izdelan iz ve~stopenjsko toplo valjanega prekurzorja z
dolomitom kot sredstvom za penjenje
4 K. Kitazono, E. Sato, Materials Science Forum, 475–479 (2005), 433
5 L. E. G. Cambonero, J. M. R. Roman, F. A. Corpas, J. M. R. Prieto,
Journal of Materials Processing Technology, 209 (2009), 1803
6 D. P. Papadopoulos, H. Omar, F. Stergioudi, S. A. Tsipas, N.
Michailidis, Colloidal and Surfaces A: Physiochem. Eng. Aspects,
382 (2011), 118
7 V. Gergely, D. C. Curran, T. W. Clyne, Composites Science and
Technology, 63 (2003), 2301
8 V. Kevorkijan, S. D. [kapin, I. Paulin, B. [u{tar{i~, M. Jenko, Mater.
Tehnol., 44 (2010) 6, 363
9 US Pat. No. 7.452.402 issued on November 18, 2008
V. KEVORKIJAN et al.: MODELLING AND PREPARATION OF CORE FOAMED Al PANELS ...
544 Materiali in tehnologije / Materials and technology 45 (2011) 6, 537–544
U. HANOGLU et al.: NUMERICAL SOLUTION OF HOT SHAPE ROLLING OF STEEL
NUMERICAL SOLUTION OF HOT SHAPE ROLLING OF
STEEL
NUMERI^NA RE[ITEV VRO^EGA VALJANJA JEKLA
Umut Hanoglu, Siraj-ul-Islam, Bo`idar [arler
Laboratory for Multiphase Processes, University of Nova Gorica, Vipavska 13, SI-5000 Nova Gorica, Slovenia
umut.hanoglu@ung.si
Prejem rokopisa – received: 2011-02-02; sprejem za objavo – accepted for publication: 2011-07-29
The modeling of hot shape rolling of steel is represented by using a meshless method. The physical model consists of coupled
thermal and mechanical models. Both models are numerically solved by using a strong formulation. The material is assumed to
behave ideally plastic. The model decomposes the 3D geometry of the steel billet into a traveling 2D cross section which lets us
analyze the large shape reductions by a sequence of small steps. A uniform velocity over each of the cross-sections is assumed.
The meshless method, based on collocation with radial basis functions is used to solve the thermo-mechanical problem. The
node distribution is calculatedby elliptic node generation at each deformation step to the new form of the billet. The solution is
calculated in terms of temperatures and displacements at each node. Preliminary numerical examples for the new rolling mill in
[tore Steel are shown.
Keywords: steel, hot rolling, radial basis functions, meshless numerical method
Modeliranje vro~ega valjanja je predstavljeno z uporabo brezmre`ne numeri~ne metode. Fizikalni model je sestavljen iz
sklopljenega termi~nega in mehanskega modela. Oba sta numer~no re{ena z uporabo mo~ne formulacije. Predpostavljamo, da se
material vede idealno plasti~no. V modelu razstavimo 3D-geometrijo jeklene gredice v premikajo~ 2D-prerez, ki omogo~a
analizo velikih sprememb oblike v majhnih korakih. Predpostavimo uniformno hitrost preko vsakega prereza. Za re{itev
termo-mehanskega problema je uporabljena brezmre`na numeri~na metoda, ki temelji na kolokaciji z radialnimi baznimi
funkcijami. Distribucijo diskretizacijskih to~k smo za vsako novo obliko prereza gredice izra~unali na podlagi elipti~nega
generatorja diskretizacijskih to~k. Re{itev je podana kot temperatura in premik v vsaki to~ki. Prikazani so preliminarni
numeri~ni primeri za novo valjarsko progo v podjetju [tore Steel.
Klju~ne besede: jeklo, vro~e valjanje, radialne bazne funkcije, brezmre`na numeri~na metoda
1 INTRODUCTION
The main aim of this paper is elaboration of the
coupled thermo-mechanical computational model deve-
loped for hot shape rolling of steel. The output of the
thermal model is the temperature field and mechanical
model the displacement (deformation). Shape rolling is a
3D process, however it is analyzed with 2D imaginary
slices which is denoted as a slice model. The coordinate
system of a 2D slice is based on Langrangian description
where the slice travels across the rolling contact. The
third axis, the rolling direction, is based on the Eularian
description where there is a constant inflow and outflow
of steel through the rolling direction. This is considered
as a mixed Eularian-Langrangian model. It was dis-
cussed previously by many authors 1,2.
In many publications of rolling Finite Element
Method (FEM) was used which is based on a mesh. A
novel numerical method used in this paper to solve the
involved partial differential equations is the Local Radial
Basis Function Collocation Method (LRBFCM). This is
a completely meshless procedure. LRBFCM has been re-
cently used in highly sophisticated simulations like
multi-scale solidification modeling 3, convection driven
melting of anisotropic metals 4, continuous casting of
steel 5. This paper is organized in a way that, first the
thermal model and afterwards the mechanical model are
developed. Overall it becomes a coupled thermo mecha-
nical model. The flow chart of the process is shown in
Figure 1.
2 THERMAL MODEL
The thermal model of the shape rolling process is
aimed to calculate the temperature field of the steel slab
during the rolling process. The three dimensional domain
3D with boundary 3D is considered. The solution
procedure is based on Cartesian coordinate system with
axes x, y, z. Slices coincide with coordinates and the
rolling direction is z. The steady state temperature
distribution in the rolled product is defined through the
following equation,
∇⋅ = ∇⋅ ∇ +( ) ( )c T k T Sp v ; p  3D (x,y,z) (1)
Since we analyze the process with 2D slices perpen-
dicular to the rolling direction and assume thata uniform
velocity over the slices (homogenous compression) takes
place. The Equation (1) can be transferred into
c
T
t
k T Sp
∂
∂
= ∇⋅ ∇ +( ) ; p  2D (x,y)
p z t vdt v A
A z
tz ( , ) ( )
= =∫ ∫entry entry d
1
(2)
Materiali in tehnologije / Materials and technology 45 (2011) 6, 545–547 545
UDK 621.771:519.68 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)545(2011)
with p, , t, cp, T, k, ventry, Aentry, (A)z and S standing for
position vector, density, time, specific heat, temperature,
thermal conductivity, entry speed of billet, entry cross
sectional area, instant cross sectional area and internal
heat generation due to plastic deformation. It is assumed
in the slice model that the heat transport takes place
only in the direction perpendicular to the rolling
direction and that the homogenous deformation takes
place. The Neumann boundary condition on the part of
the boundary denoted as N, Robin boundary condition
on the part of the boundary denoted as R are taken into
consideration ( = N  R) which are described below,
− ∇ ⋅ = − =k T k
T
q( )
( )
p n
p
n 
∂
∂
; p  N (3)
[ ]− ∇ ⋅ = − = −k T k T h T T R( ) ( ) ( ) ( )p n p
n
p p


∂
∂
; p  R (4)
The N nodes at the domain and N nodes at the
boundary are used to discretize the temperature in
LRBFCM where for each node pn = {px, py}T. For each
node there is a defined influence domain with N
neighboring nodes. For each influence domain a radial
basis function in terms of multiquadric is written
y p p /x p p /y ci x xn y yn= − + − +( ) ( )max max
2 2 2
The temperature can now be interpolated as
T n
n
N
n=
=
∑  	

1
with the collocation coefficients to be determined. The
main equation in 2D can be rewritten by using the
explicit time stepping,
 
  	

  	


c
T T
t
k
k
p
i i
i n
n
N
n
i n
n
N
n
+
=
=
−
= ∇ ⋅ ∇ +
+ ⋅ ∇
⎛
⎝
∑
∑
1
1
1
Δ
)⎜
⎞
⎠
⎟ + S
; p  2D (5)
3 MECHANICAL MODEL
A strong form is chosen here for analysis due to its
compatibility with LRBFCM. Adomain  with
boundary ,  = U  T is considered where U is the
essential and T represents the natural boundary
conditions. The strong formulation of the static metal
deformation problems is:
LT + b = 0 (6)
In the calculations, in order to avoid complications of
a 3D solution, the slab is analyzed, compatible with the
thermal model, with imaginary traveling 2D slices that
are perpendicular to the rolling direction. For a 2D slice
method, L is the 3×2 derivative matrix with elements
L px11 = ∂ ∂/ , L12 0= , L21 0= , L py22 = ∂ ∂/ , L py31 = ∂ ∂/
and L py32 = ∂ ∂/ , s = [ ]
T  x y xy, , is the stress vector,
and b = [ ]Tb bx y, is the body force. At the essential
boundary U
u(p) = u(p) ; p  U (7)
where u(p) is displacement vector and u(p) is the
prescribed displacement vector. At the naturalboundary
condition T
NTs = t ; p  T (8)
is valid, where t is the prescribed surface traction
t= x y[ ]
 
,
T , N is the 2×3 matrix of direction cosines of
the normal direction at the boundary which can be
defined as N11 = N32 = nx, N12 = N21 = 0, N31 = N22 = ny
(nx,ny) represent correlation of the normal at the boun-
dary). In a 2D system the equations for mechanical
model can be written as,
∂
∂
∂
∂
 x
x
xy
y
xp p
b+ + = 0,
∂
∂
∂
∂
 y
y
xy
x
yp p
b+ + = 0 (9, 10)
The discretization is made in terms of displacement
on x and y axes for each slice,
ux n
n
N
xn( ) ( )p p=
=
∑  	

1
, uy n
n
N
yn( ) ( )p p=
=
∑  	

1
(11, 12)
Since the strain vector  = [x,y,xy]T can be written in
terms of displacement as e = Lu, the strain vector  can
be expressed as a stress vector by using 6×6 stiffness
matrix C which depends on the material characteristic
assumption such as elastic, elastic-plastic or ideally
plastic.
	
  
xn
n
N
n
x
n
y
n
y
C
p
C
p
C C
p=
∑ + + +
1
11
2
2 33
2
2 13 31
2∂
∂
∂
∂
∂
∂ ∂
( )
p
C C
p p
C
p
x
yn
n
N
n
y x
n
x
⎡
⎣⎢
⎤
⎦⎥
+
+ + +
=
∑ 	  

1
12 33
2
13
2
2( )
∂
∂ ∂
∂
∂
+
⎡
⎣⎢
⎤
⎦⎥
+ =C
p
bn
y
x32
2
2 0
∂
∂

(13)
U. HANOGLU et al.: NUMERICAL SOLUTION OF HOT SHAPE ROLLING OF STEEL
546 Materiali in tehnologije / Materials and technology 45 (2011) 6, 545–547
Figure 1: Flow chat of the coupled thermo mechanical model.
Slika 1: Blo~ni diagram sklopljenega termo-mehanskega modela

  
xn
n
N
n
y x
n
x
nC C
p p
C
p
C
p=
∑ + + +
1
21 33
2
31
2
2 23
2
( )
∂
∂ ∂
∂
∂
∂
∂ y
yn
n
N
n
y
n
y x
C
p
C C
p p
2
1
22
2
2 23 32
2
⎡
⎣⎢
⎤
⎦⎥
+
+ + +
=
∑   
 ∂
∂
∂
∂ ∂
( ) +
⎡
⎣⎢
⎤
⎦⎥
+ =C
p
bn
x
y33
2
2 0
∂
∂

(14)
4 TRANSFINITE INTERPOLATION (TFI)
This technique is used to generate initial grid which
is confirming to the geometry encountered in different
stages of plate and shape rolling. Suppose that there
exists a transformation r(p, p) = {px(p, p), py(p, p)}T
which maps the unit square, 0 < p < 1, 0 < p < 1 in the
computational domain onto the interior of the region
ABCD in the physical domain such that the edges p =
0,1 map to the boundaries AB, CD and the edges p = 0,1
are mapped to the boundaries AC, BD. The transfor-
mation is used for this purpose is defined as
r(p,p) = (1 – p)rl(p) + rr(p)rb(p) +
prt(p) – (1 – p)(1 – p)rb(0) – (1 – p)prt(0) – (15)
(1 – p)prb(1) – pprt(1)
Where rb, rt, rl, rr represent the values at the bottom,
top, left and right edges respectively. The initial grid is
refined through ENG 6. Figure 2 shows initial node
generation through TFI and its correlation with ENG.
5 CONCLUSION
In this paper the thermal and mechanical formula-
tions are given for hot shape rolling. The numerical
method for the solution of the problem is based on
meshfree LRBFCM. The preliminary result of mechani-
cal model for elastic case is presented in Figure 3. The
future work will include plastic deformation in a se-
quence of 10 rolling stands as recently installed in [tore
–Steel Company.
6 REFERENCES
1 J. Synka and A. Kainz, International Journal of Mechanical Sciences,
45 (2003), 2043–2060
2 M. Glowacki, Journal of Materials Processing Technology,168
(2005), 336–343
3 B. [arler, G. Kosec, A. Lorbiecka and R. Vertnik, Mater. Sci. Forum,
649 (2010), 211–216
4 G. Kosec and B. [arler, Int. J. Cast. Met. Res., 22 (2009), 279–282
5 R. Vertnik and B. [arler, Int. J. Cast. Met. Res., 22 (2009), 311–313
6 J. F. Thompson, B. K. Soni and N. P. Weatherill, Handbook of grid
generation, 1st ed., CRC Press, USA 1999
7 W. F. Chen and D. J. Han, Plasticity for Structural Engineers, 1st ed.,
Springer-Verlag, USA 1988, p. 606
U. HANOGLU et al.: NUMERICAL SOLUTION OF HOT SHAPE ROLLING OF STEEL
Materiali in tehnologije / Materials and technology 45 (2011) 6, 545–547 547
Figure 3: Simulation of flat rolled (180 × 180) mm cross sectioned
16MnCrS5 steel at 1100 °C with Young’s modulus E = 97362.21 MPa
and Poisson’s ratio v = 0.35678. The total reduction is 16.66 % and
preliminary analyzed with 5 slices by using elastic stiffness matrix 7.
Arrows represents the displacement vector for each slice. The exit
speed is equal to 1.14389 times the entry speed of the billet. Due to
symmetry only the top right part of the billet is considered.
Slika 3: Simulacija prereza (180 × 180) mm plo{~atega valjanja za
jeklo16MnCrS5 pri 1100 °C z Youngovim modulom E = 97362.21
MPa in Poissonovim razmerjem v = 0.35678. Skupno zmanj{anje je
16,66 % in predhodno analizirano s 5 rezinami z uporabo elasti~ne
togostne matrike 7. Pu{~ice pomenijo vektor premika za vsako rezino.
Izhodna hitrost je enaka 1.14389-kratniku vhodne hitrosti gredice.
Zaradi simetrije je upo{tevana samo zgornja polovica gredice.
Figure 2: Transformation from computational domain to physical
domain (left), TFI and nodes displacement through ENG (right). The
collocation points are put on the intersection of grid lines.
Slika 2: Transformacija izra~unskega obmo~ja v fizi~no obmo~je
(levo) TFI in premik to~k preko ENG (desno). Kolokacijske to~ke so
postavljene v prerez mre`nih linij.

M. VON^INA et al.: SOLIDIFICATION AND PRECIPITATION BEHAVIOUR IN THE AlSi9Cu3 ALLOY ...
SOLIDIFICATION AND PRECIPITATION BEHAVIOUR IN
THE AlSi9Cu3 ALLOY WITH VARIOUS Ce ADDITIONS
STRJEVANJE IN IZLO^ANJE V ZLITINI ALSI9CU3 PRI
RAZLI^NIH DODATKIH Ce
Maja Von~ina, Stanislav Kores, Primo` Mrvar, Jo`ef Medved
University of Ljubljana, Faculty of Natural Sciences and Engineering, Department of Materials and Metallurgy, A{ker~eva 12,
1000 Ljubljana, Slovenia
maja.voncina@omm.ntf.uni-lj.si
Prejem rokopisa – received: 2011-03-07; sprejem za objavo – accepted for publication: 2011-08-12
The effect of Ce additions on the AlSi9Cu3 alloy was investigated using an equilibrium thermodynamic calculation, thermal
analysis, differential scanning calorimetry (DSC) and scanning electron microscopy (SEM). The purpose was to study the
variations that occur during solidification and precipitation with different Ce additions, as well as theireffect on the mechanical
properties.
The results show that Ce additions shift the temperature of the eutectic solidification (Al + Al2Cu) and the solidus temperature
to higher values. It was found that the precipitation reaction is more intense when the specimen is previously cooled with a
higher cooling rate. Moreover, when the fraction of the precipitates regarding the temperature at different cooling rates was
taken into account, it was found that the precipitation is faster when Ce is added and also when the specimen was cooled faster.
Ce also changed the morphology of the eutectic Al2Cu phase. Furthermore, the Ce phase was detected, indicating the
Al–Ce–Cu–Si (Al9Ce2Cu5Si3) phase. The mechanical properties, such as hardness and tensile strength, increase with larger Ce
additions.
Keywords: AlSi9Cu3 alloy, Ce addition, reaction kinetics, solidification and precipitation, mechanical properties
Vpliv dodatka Ce v zlitini AlSi9Cu3 je bil preiskan z uporabo izra~una ravnote`nega strjevanja, termi~ne analize, diferen~ne
vrsti~ne kalorimetrije (DSC) ter vrsti~ne elektronske mikroskopije (SEM). Namen je bil preiskati spremembe, ki nastopijo pri
strjevanju in izlo~anju v zlitini z razli~nimi dodatki Ce ter njegov vpliv na mehanske lastnosti.
Rezultati so pokazali, da dodatek Ce zvi{a temperaturo strjevanja evtektika (Al+ Al2Cu) in solidus temperature k vi{jim
vrednostim. Ugotovljeno je bilo, da reakcija izlo~anja pote~e intenzivnej{e, ko je vzorec predhodno ohlajen z ve~jo hitrostjo
ohlajanja. Pri preiskavi dele`a izlo~kov glede na razli~ne temperature se poka`e, da pote~e izlo~anje hitreje, ko zlitini dodamo
Ce, ter prav tako, ko je vzorec predhodno hitreje ohlajen. Ce v zlitini AlSi9Cu3 spreminja tudi morfologijo evtektske faze
Al2Cu. Analizirali smo Ce fazo AlCeCuSi (Al9Ce2Cu5Si3). Mehanske lastnosti, kot sta trdota ter natezna trdnost, se pove~ujejo z
dodatkom Ce v zlitini AlSi9Cu3.
Klju~ne besede: zlitina AlSi9Cu3, dodatek Ce, reakcijska kinetika, strjevanje ter izlo~anje, mehanske lastnosti
1 INTRODUCTION
Al–Si–Cu alloys are widely used for thin-wall
castings1 in automobile, aircraft and in the chemical in-
dustry. The AlSi9Cu3 alloy is a heat-treatable alloy with
good castability. These alloys usually contain copper and
often magnesium as the main alloying element, together
with various other alloying elements or impurities, such
as Fe, Mn or Cr.2 Cu in the AlSi9Cu3 alloy reduces the
corrosion resistance and improves the mechanical prop-
erties. A controlled cooling process and/or optimal alloy-
ing with Ce makes it possible to achieve suitable me-
chanical properties, like tensile strength and hardness. A
small amount of Mg causes the formation of the Mg2Si
phase3,4 and additionally it increases the mechanical
properties in Al–Si–Cu alloys.5,6,7,8,9 The microstructure
in Al–Si alloys dictates the mechanical and technological
properties of the castings. For this reason a specific
microstructure and the mechanical properties must be
achieved. This can be established with a smaller grain
size and with a modification of the (Al +  Si) eutectic
and/or with high cooling rates.
Rare-earth metals, such as cerium (Ce), have been
found to improve the mechanical properties of Al–Si
castings by modifying their microstructure and enhanc-
ing their tensile strength10 and ductility11, heat resistance
and extrusion behaviour.12 It was reported that Ce-phases
may act as nucleation sites for (Al) or (Si) crystals in
both hypo- and hypereutectic Al-Si alloys.13 Cerium has
a high activity in an aluminium melt because of its spe-
cific electron structure. It forms a quaternary
intermetallic compound with aluminium, silicon and
copper and this leads to the formation of an
Al–Ce–Cu–Si phase between the dendrite structure. The
cerium phase acts as a barrier for dislocation movement
that increases the mechanical properties of the mate-
rial,14.
This paper treats the influence of Ce addition on the
course of the solidification, precipitation and cooling,
and also of the cooling rate, on the solidification and pre-
cipitation and on the mechanical properties of the
AlSi9Cu3 alloy.
Materiali in tehnologije / Materials and technology 45 (2011) 6, 549–554 549
UDK 669.715:620.17 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)549(2011)
2 EXPERIMENTAL
A commercial AlSi9Cu3 alloy was melted in an elec-
tric induction furnace.Various concentrations (w(Ce) = 0,
0.01, 0.02, 0.05 and 0.1) of pure (99.9 %) Ce were
added. After 10 min the melt was poured into a measur-
ing cell with a controlled cooling system (simple thermal
analysis-STA) with the purpose of recording the cooling
curves at different cooling rates. A new measuring cell
for the controlled cooling of specimens from the melt to
low temperatures was designed in order to obtain various
cooling rates. Simultaneously, the specimens for the ten-
sile tests were also cast into a mould made according to
the DIN50125 standard. The characteristic solidification
temperatures were determined from the cooling curves,
and the influence of Ce was defined.
Differential scanning calorimetry (DSC) using a Jupi-
ter 449c, NETZSCH, was applied to analyse the solidifi-
cation process and to determine the characteristic tem-
peratures of single reactions and the produced or
consumed enthalpies. The measurements were carried
out under a protective Ar atmosphere according to the
temperature program: heating rate 10 °C/min up to 710
°C  holding at 710 °C for 10 min  cooling rate 10
°C/min. Moreover, the DSC curves were plotted, the
temperatures of the precipitation were marked and the
formation enthalpies of the precipitates were determined.
The precipitation kinetics connected to the Ce addition
and the cooling rate was also determined.
Light and electron microscopy were applied to ana-
lyse the microstructures. Single microstructural phases
were determined quantitatively with the system for ana-
lysing images. A quantitative analysis for the identifica-
tion of the phases was performed by energy-dispersive
and wave length-dispersive X-ray spectroscopy. A ce-
rium phase was identified. The hardness was measured
using a universal Brinell hardness tester and the tensile
strength was defined on as-cast specimens made accord-
ing to the EN 10002-1 standard using a GLEEBLE
1500D simulator of thermomechanical states.
3 RESULTS AND DISCUSSION
The chemical composition of the investigated sam-
ples is presented in Table 1.
From the chemical composition, equilibrium solidifi-
cation and calculated equilibrium the vertical cross-sec-
tion diagrams were simulated using the Thermo-Calc
program TCW5 and database COST507 (Figure 1a).
The course of the equilibrium solidification was deter-
mined (Figure 1b).
The equilibrium solidification of the AlSi9Cu3 alloy
proceeds as follows (Figure 1): Si2Ti, AlFeSi- , primary
crystals of Al, AlMnSi-, eutectic (Al +  Si) and just be-
low the solidus temperature the Ce phase Al8Ce. Under
the solidus, the Mg2Si and Al2Cu-! phase also precipi-
tated.
M. VON^INA et al.: SOLIDIFICATION AND PRECIPITATION BEHAVIOUR IN THE AlSi9Cu3 ALLOY ...
550 Materiali in tehnologije / Materials and technology 45 (2011) 6, 549–554
Figure 1: Equilibrium phase diagram (a) and schematic representation of equilibrium solidification of AlSi9Cu3 alloy (b) with w = 0.02 % Ce
Slika 1: Ravnote`ni fazni diagram (a) ter shematski prikaz ravnote`nega strjevanja zlitine AlSi9Cu3 (b) z w = 0,02 % Ce
Table 1: Chemical composition of AlSi9Cu3 alloy, w/%
Tabela 1: Kemijska sestava zlitine AlSi9Cu3, w/%
Specimen Mg Mn Cu Ti Fe Si Ce
(nominal)
Ce
(actual)
Al
AlSi9Cu3 0.35 0.242 2.61 0.04 0.694 10.72 0 rest
AlSi9Cu3 + 0.01 % Ce 0.34 0.27 2.55 0.04 0.75 10.60 0.01 rest
AlSi9Cu3 + 0.02 % Ce 0.35 0.29 2.685 0.04 0.80 10.66 0.015 rest
AlSi9Cu3 + 0.05 % Ce 0.32 0.29 2.565 0.04 0.81 10.59 0.043 rest
Figure 2 shows a typical cooling curve together with
a differential cooling curve of the investigated AlSi9Cu3
alloy with 0.02 % Ce. The characteristic solidification
temperatures were determined in all the specimens with
various Ce additions (Table 2). The striped line indicates
the theoretical, with the Thermo-Calc calculated,
liquidus temperature calculated from the chemical com-
position:
TL teor. = 656.38468 – 6.78571 · w(Si) – 1.42857 ·
w(Cu) + 1.34798 · 10–10 · w(Fe) – 1.04224 · 10–10 ·
w(Mn) – 3.15848 · w(Mg) – 2.24953 · w(Zn)
The dotted line in Figure 2 indicates the theoretically
calculated eutectic temperature calculated from the
chemical composition:
TEteor. = 574.2834 – 0.57134 · w(Si) – 2.57143 · w(Cu)
– 3 · w(Fe) – 1.14639·10–10 · w(Mn) – 5.73489 · w(Mg)
– 1.38954 · w(Zn)
Table 2 and Figure 3 represent the characteristic so-
lidification temperatures with respect to the Ce addition.
The temperature of the eutectic solidification (Al +
Al2Cu) and the solidus temperature shift to higher tem-
peratures. The temperature of the eutectic solidification
(Al + AlCuMgSi) could not be detected when larger
M. VON^INA et al.: SOLIDIFICATION AND PRECIPITATION BEHAVIOUR IN THE AlSi9Cu3 ALLOY ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 549–554 551
Figure 4: Heating and cooling DCS curves of AlSi9Cu3 alloy without
Ce (a) and with w = 0.02 % Ce (b)
Slika 4: Segrevalne in ohlajevalne DSC-krivulje zlitine AlSi9Cu3 brez
Ce (a) in z w = 0,02 % Ce (b)
Figure 2: Cooling curve and differential cooling curve of AlSi9Cu3
alloy with w = 0.02 % Ce and with the theoretical calculated equilib-
rium liquidus and eutectic temperature (horizontal line)
Slika 2: Ohlajevalna in diferencirana ohlajevalna krivulja zlitine
AlSi9Cu3 z w = 0,02 % Ce ter z vrisanima teoreti~no izra~unanima
ravnote`no likvidus in evtektsko temperaturo (vodoravne ~rte)
Figure 3: Comparison of some characteristic solidification tempera-
tures for AlSi9Cu3 alloy with respect to Ce addition
Slika 3: Primerjava nekaterih karakteristi~nih temperature pri
strjevanju zlitine AlSi9Cu3 glede na dodatek Ce
Table 2: Characteristic solidification temperatures for AlSi9Cu3 after STA
Tabela 2: Zna~ilne temperature strjevanja zlitine AlSi9Cu3 po STA
w(Ce)/% TLteor/°C TL/min/°C TL/max/°C TL p./°C TL r./°C
0 578.8 561 564 17.81 3
0.01 579.7 562 566.8 17.71 4.8
0.02 579.1 560.5 563 18.6 2.5
0.05 579.9 563.7 565.9 16.1 2.2
w (Ce)/% TEteor./°C
TE/min
/°C
TE/max
/°C
TE p.
/°C
TE r.
/°C
TE2(Mg2Si)
/°C
TE3(Al2Cu)
/°C
TE4(AlCuMgSi)
/°C
TS
/°C
0 557.4 556.5 559 0.9 2.5 512 494 478.5 463
0.01 557.5 563.9 564.7 –6.4 0.8 520.1 501.8 483.2 475
0.02 556.9 556 559 0.9 3 507 497.5 476.5
0.05 557.4 563.4 564.4 –6.0 1 523.2 503.3 480.5
concentrations of Ce were added, presumably because
these elements combined with Ce.
The DSC analysis was made on all the specimens af-
ter the STA. From the heating (Figure 4a) and cooling
(Figure 4b), all the characteristic temperatures during
heating/cooling were determined, including the melt-
ing/solidification and precipitation enthalpies with vari-
ous Ce additions.
When the precipitation kinetics was investigated, it
was found that the precipitation reaction is more intense
when the specimen is previously cooled faster (Figure
5). This is a consequence of a more supersaturated solid
solution. Moreover, when the fraction of the precipitates
regarding the temperature at different cooling rates was
taken into account, it was found that the precipitation ki-
M. VON^INA et al.: SOLIDIFICATION AND PRECIPITATION BEHAVIOUR IN THE AlSi9Cu3 ALLOY ...
552 Materiali in tehnologije / Materials and technology 45 (2011) 6, 549–554
Figure 8: Microstructure (SEM) of AlSi9Cu3 + 0.02 % Ce with EDS
analysis
Slika 8: Mikrostruktura (SEM) zlitine AlSi9Cu3 + 0,02 % Ce z
EDS-analizo
Figure 6: Fraction of Al2Cu precipitates with respect to the tempera-
ture in the AlSi9Cu3 alloy at various cooling rates (a) and at various
concentrations of Ce (b)
Slika 6: Dele` izlo~kov Al2Cu v odvisnosti od temperature v zlitini
AlSi9Cu3 po razli~nih hitrostih ohlajanja (a) ter pri razli~nih koncen-
tracijah Ce (b)
Figure 7: Microstructure of Al2Cu phase without Ce (a) and with w =
0.05 % Ce (b)
Slika 7: Mikrostrukturni posnetek faze Al2Cu brez Ce (a) in z w =
0,05 % Ce (b)
Figure 5: DSC curves of the precipitation region of AlSi9Cu3 alloy at
different cooling rates
Slika 5: DSC-krivulje s podro~ja izlo~anja zlitine AlSi9Cu3 pri
razli~nih hitrostih ohlajanja
netics is faster when Ce is added (Figure 6a) and also
when the specimen was cooled faster (Figure 6b).
When the microstructure was investigated, no influ-
ence on the size and distribution of the microstructure
components was detected, only the morphology of the
Al2Cu phase changed. In the alloy without Ce, the
eutectic phase Al2Cu appears to be "crumbled" (Figure
7a), but when Ce was added the Al2Cu phase was fully
formed (Figure 7b).
Besides the usual phases that occur in these types of
alloys, the Ce phase was also detected with the EDS
analyser, indicating the Al–Ce–Cu–Si (calculated
stoichiometry was Al9Ce2Cu5Si3, Figure 8) phase. This
phase forms in a needle shape. To establish what hap-
pens with the Al2Cu, mapping (Figure 9a) and line anal-
yses (Figure 9b) through the Al2Cu were made. It was
proved that the Ce is bound to the Al2Cu eutectic phase.
The tensile strength and Brinell hardness of the
AlSi9Cu3 alloy with respect to Ce are presented in Fig-
ure 10a and 10b. The tensile strength for a small
amount of Ce is slightly reduced and at a higher concen-
tration of Ce it is increased, probably because of the
modified Al2Cu eutectic phase. The hardness due to the
Ce addition was investigated for different cooling rates.
For the smaller cooling rate (10 K/min) the hardness
slightly decreased when 0.01 % Ce was added to the al-
loy, but it increased as well as the tensile strength for
higher Ce additions. With higher cooling rates the hard-
ness increased because the influence of the Ce was re-
duced.
M. VON^INA et al.: SOLIDIFICATION AND PRECIPITATION BEHAVIOUR IN THE AlSi9Cu3 ALLOY ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 549–554 553
Figure 10: Tensile strength (a) and hardness (b) of AlSi9Cu3 alloy for
various Ce additions
Slika 10: Natezna trdnost (a) ter trdota (b) zlitine AlSi9Cu3 pri raz-
li~nih koncentracijah Ce
Figure 9: Mapping (a) and line analyses (b) of (Al + Al2Cu) eutectic in AlSi9Cu3 alloy with w = 0.02 % Ce
Slika 9: Povr{inska porazdelitev elementov ter linijska analiza (b) evtektske faze (Al + Al2Cu) v zlitini AlSi9Cu3 z w = 0,02 % Ce
4 CONCLUSION
The effects of Ce content on the solidification se-
quence, microstructure and mechanical properties of the
AlSi9Cu3 alloy were investigated. Moreover, the reac-
tion kinetics of the precipitation in the AlSi9Cu3 alloy
was studied also. The results can be summarized as fol-
lows:
The equilibrium solidification of the AlSi9Cu3 alloy
proceeds as follows: Si2Ti, AlFeSi-, primary crystals of
Al, AlMnSi-, eutectic (Al + Si) and just below the sol-
idus temperature the Ce phase Al8Ce is precipitated. Be-
low the solidus the Mg2Si and Al2Cu- phases are pre-
cipitated also. The data base should be complemented
with multicomponent phases with Ce.
The temperature of the eutectic solidification (Al +
Al2Cu) and the solidus temperature are shifted to higher
temperatures with the addition of Ce. The temperature of
the eutectic solidification (Al + AlCuMgSi) could not be
detected when greater concentrations of Ce were added.
When the precipitation kinetics was investigated it
was found that the precipitation reaction is more intense
for the specimen that was previously cooled faster.
Moreover, when the fraction of the precipitates depend-
ing on the temperature at different cooling rates was
taken into account, it was found that the precipitation is
faster when Ce is added and also when the specimen was
cooled faster.
Ce changed the morphology of the eutectic Al2Cu
phase with building into the Al2Cu phase. Furthermore,
the Ce phase was detected, indicating the Al–Ce–Cu–Si
(Al9Ce2Cu5Si3) phase. This phase forms in a needle
shape.
The tensile strength and the hardness vs. slower cool-
ing for a small amount of Ce were slightly reduced.
However, they were increased when a greater concentra-
tion of Ce is added, probably because of the modification
of the Al2Cu eutectic phase. With higher cooling rates
the hardness increased and the influence of Ce is re-
duced.
Acknowledgements
The authors would like to thank dr. Franc Zupani~,
University of Maribor, Faculty of Mechanical Engi-
neering, and dr. Ale{ Nagode, University of Ljubljana,
Faculty of Natural Sciences and Engineering, for work
on the SEM.
5 REFERENCES
1 ASM Metals HandBook Volume 15 – Casting, 1988
2 L. Lasa, J. M. Rodriguez-Ibabe, Characterization of the dissolution
of the Al2Cu phase intwo Al–Si–Cu–Mg casting alloys using calo-
rimetry, Materials Characterization 48 (2002), 371–378
3 G. Terjesen, Effect of elevated temperature on tensile properties and
fracture toughness of an AlSi10Mg alloy. Aluminium, 79 (2003) 9,
748–754
4 B. K. Shah, S. D. Kumar, D. K. Dwivedi, Aging temperature and
abrasive wear behaviour of cast Al-(4 %, 12 %, 20 %)Si-0.3 % Mg
alloys, Materials and Design, 28 (2007), 1968–1974
5 L. Bäckerud, G. Chai, J. Tamminen,Solidification Characteristics of
Aluminum Alloys. Volume 2, Foundry Alloys. AFS/SKANALU-
MINIUM.Department of Structural Chemistry – Arrhenius Labora-
tory, University of Stockholm, 1990
6 S. Esmaeili, D. J. Lloyd, W. J. Poole, Modelling of precipitation
hardening for the naturally aged Al-Mg-Si-Cu alloy AA6111,
ActaMaterialia, 51 (2003), 3467–3481
7 S. Esmaeili, D. J. Lloyd, Modelling of precipitation hardening in
pre-aged AlMgSi(Cu) alloys, ActaMaterialia, 53 (2005), 5257–5271
8 Q. G. Wang, C. J. Davidson, Solidification and precipitation behav-
iour of Al-Si-Mg casting alloys, Journal of material science, 36
(2001), 739–750
9 C. Vorge, M. Jucgues, M. P. Schmidt., Corrosion of Aluminium,
2001
10 M. Von~ina, S. Kores, J. Medved, Influence of Ce addition on the so-
lidification and mechanical properties of AlSi10Mg alloy, Tofa 2010
discussion meeting on Thermodynamics OF Af, Book of abstracts
and Programme, Faculdade de Engenharia da Universidade do Porto,
Portugal, 2010, 75
11 H. R. Ammar, C. Moreau, A. M. Samuel, F. H. Samuel, H. W. Doty,
Influences of alloying elements, solution treatment time and quench-
ing media on quality indices of 413-type Al–Si casting alloys, Mate-
rials Science and Engineering A, 20 (2008), 426–438
12 Y. Wu, J. Xiong, R. Lai, X. Zhang, Z. Guo, The microstructure evo-
lution of an Al–Mg–Si–Mn–Cu–Ce alloy during homogenization,
Journal of Alloys and Compounds, 475 (2009) 1–2, 332–338
13 J. Gröbner, D. Mirkovi}, R. Schmid - Fetzer, Thermodynamic As-
pects of the Constitution, Grain Refining, and Solidification
Enthalpies of Al-Ce-Si Alloys, metallurgical and materials transac-
tions A, 35A (2004), 3349–3362
14 Y. G. Zhao, Q. D. Qin, W. Zhou, Y. H. Liang, Microstructure of the
Ce-modified in situ Mg2Si/Al–Si–Cu composite, Journal of Alloys
and Compounds, 389 (2005), L1–L4
M. VON^INA et al.: SOLIDIFICATION AND PRECIPITATION BEHAVIOUR IN THE AlSi9Cu3 ALLOY ...
554 Materiali in tehnologije / Materials and technology 45 (2011) 6, 549–554
D. A. SKOBIR BALANTI^ et al.: EFFECT OF CHANGE OF CARBIDE PARTICLES SPACING ...
EFFECT OF CHANGE OF CARBIDE PARTICLES
SPACING AND DISTRIBUTION ON CREEP RATE OF
MARTENSITE CREEP RESISTANT STEELS
VPLIV SPREMEMBE RAZDALJE MED KARBIDNIMI IZLO^KI IN
NJIHOVE PORAZDELITVE NA HITROST LEZENJA
MARTENZITNIH JEKEL, ODPORNIH PROTI LEZENJU
Danijela A. Skobir Balanti~, Monika Jenko, Franc Vodopivec, Roman Celin
Institute of Metals and Technology, Lepi pot 11, SI-1000 Ljubljana, Slovenia
danijela.skobir@imt.si
Prejem rokopisa – received: 2011-10-04; sprejem za objavo – accepted for publication: 2011-10-25
The creep rate dependence of particles coarsening and spacing as well as distribution in analysed considering quoted equations.
A simple method for assessment of particle spacing is proposed. Accelerated creep rates at 580 °C for CrV and CrVNb steel
after different tempering times at 800 °C and 650 °C are calculated and determined experimentally. The rate of microstructural
processes increases the creep rate at 800 °C in the CrV steel by 36 times and in the CrVNb by 57 times greater.
Key words: creep resistant steel, creep rate, carbide particles, particles spacing, distribution of particles, tempered martensite
Matemati~no smo analizirali odvisnost hitrosti lezenja od rasti izlo~kov in njihove medsebojne razdalje ter porazdelitve.
Predlagali smo preprosto metodo za oceno medsebojne razdalje med izlo~ki. Izra~unali in eksperimentalno smo dolo~ili
pospe{ene hitrosti lezenja pri 580 °C za jekli CrV in CrVNb po razli~nih ~asih `arjenja pri 800 °C in 650 °C. Mikrostrukturni
procesi, ki vplivajo na zvi{ano hitrost lezenja, potekajo v jeklu CrV 36-krat hitreje in v jeklu CrVNb 57-krat hitreje pri 800 °C.
Klju~ne besede: jekla odporna proti lezenju, hitrost lezenja, karbidni izlo~ki, razdalja med karbidnimi izlo~ki, porazdelitev
izlo~kov, popu{~eni martenzit
1 INTRODUCTION
Modern creep resistant steels have a microstructure
consisting of a distribution of carbide particles in ferrite
with a significant content of chromium and molybdenum
in solid solution.1,2 In these steels particles of inter-
metallic (Lawes) phases are also found,3,4 depending on
steel composition and tempering temperature. These par-
ticles are generally much coarser than carbide parti-
cles,3,4 have a minor effect on creep rate and will be
omitted in further discussion. The great majority of parti-
cles consists of carbides with different stability at in-
creased temperatures, mostly of M23C6 and MC particles.
The composition of M23C6 particles depends on anneal-
ing temperature, however, the content of chromium is al-
ways much higher than the contents of iron and molyb-
denum.5,6 MC particles are mostly vanadium and
niobium carbides that may also have a minor content of
nitrogen.
The effect of temperature on the solubility of carbide
phases is given by the general relation:
lg [M] [C] = A + (B/T) (1)
with [M] and [C] mass fractions of elements in solid so-
lution in ferrite, A and B – constants and T/K – tempera-
ture.
At low solubility, particles are more stable and coarse
slower because, according to the LSW equation (2), the
coarsening rate depends of volume diffusion transport,
which is smaller by low content of carbide forming ele-
ments in substitutional solution in ferrite. The solubility
products for carbides in ferrite were established for VC
and NbC in structural steels with a much lower content
of chromium.7,8 These products can not be used reliably
for creep resisting steels with much higher content of
chromium which is a strong carbide forming element and
may affect the solubility of MC carbides. Based on a
thermodynamic analysis and on solubility product of VC
in structural steel, it was calculated that at 873 K (600
°C) the solubility of NbC particles in ferrite was for one
order of magnitude lower than for VC carbide.9 The sol-
ubility can be deduced also by Thermocalc analysis.
The coarsening rate of carbide particles is calculated
applying the Lifshitz-Slyozov-Wagner10 (LSW) equation:
d3 = dt
3 – d0
3 = 8 SDt / 9kBT (2)
With dt – particles size at tempering time t, d0 – initial
particles size, S – content of carbide constituting metal in
solution in the matrix,  – carbide particle matrix interfa-
cial energy,  – volume of diffusing atoms, kB –
Boltzmann constant, D – diffusion coefficient and T –
temperature in K. In ref.11 the exponent 2 was proposed
for coarsening of grain boundary particles.
For the 0.18C-11.7Cr-1Mo-0.29V steel, the experi-
mental coarsening rate based on the assessment of parti-
cles size after tempering of steel specimens up to 1356 h
Materiali in tehnologije / Materials and technology 45 (2011) 6, 555
UDK 669.14.018.44:539.37 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)555(2011)
at 800 °C was kce = 1.48 × 10–26 m3 s–1 and in acceptable
agreement with the coarsening rate d3cc = 7.25 × 10–27
m3 s–1 deduced applying the LSW equation.12 The in-
crease of size of particles at a determined temperature
can be calculated using the simple relation:11,12,13,14
d3 = kcit (3)
with kci – isothermal coarsening rate.
Based on the acceptable agreement of calculated and
experimental coarsening rate at 800 °C, assuming that by
change of temperature in equation (1) only the diffusion
constant changed and using as base the coarsening rate
calculated for the temperature 800 °C, for the calculation
of coarsening rate of M23C6 particles in range 550 °C to
800 oC the relation was obtained:12
d3cc,T = kcc,Cr,1073 (DCr,T/DCr,1073) (T/1073) t (4)
with kcc – calculated coarsening rate at 800 °C, T/K –
temperature, D – diffusion constant, t/s – time.
The coarsening rate of 2.89 × 10–27 m3 s–1 at 750 °C
was calculated for M23C6 particles in the steel 0.18C-
11.7Cr-1Mo-0.29V using this equation and S = 0.277
J m–2 very similar to that reported in ref.15 of 2.82 × 10–27
m3 s–1 calculated using a model based on equations (1)
and (2). At low temperature the differences obtained us-
ing equation (4) and experimental data on assessment
particle size may be affected by the simultaneous coars-
ening of particles of different carbides.
For the calculation of the coarsening rate for VC par-
ticles at the temperature of 650 °C a modification of
equation (4) was proposed with introducing of proper ra-
tio of parameters for chromium and vanadium12 in equa-
tion (1), thus:
kcc,VC, = kcc,Cr (SV DV,K M23C6 V/ SCr DCr,K MC Cr )
kcc,v,923 K = 4.77 × 10
–29 × 0.01623 =
= 1.28 × 10–31 m3 s–1 (5)
For the 0.18C-11.7Cr-1Mo-0.29V steel and the tem-
perature of 650 °C the calculated coarsening rate for VC
particles was for about two orders of magnitude lower
than for M23C6 particles.12
Creep increases the number of mobile dislocations in
steel and for the calculation of the density of these dislo-
cations the relation was proposed12,16:
 = 	/ MGb (6)
with 	/MPa – stress,  = 0.4 – constant M = 3 (Taylor
factor), G – shear modul at creep temperature and b –
Burgers vector.
By creep tests at 923 K by 80 MPa stress of a
0.14C-12Cr-1.5Mo-0.2V-0.05Nb-0.05N steel the size of
MX particles was assessed in grip and in gauge of tested
specimens.17 Using the simplified relation d3 = kci t it
was deduced that the coarsening rate was 3.73 times
greater in the gauge than in the grip part of specimens
and, according to equation (6), the density of mobile
dislocations12 in the gauge part of specimen was of 2.40
× 1013 m–2. The increased coarsening of particles in
gauge part of specimens in ascribed to higher density of
mobile dislocations.
For the dependence of creep rate on different physi-
cal parameters related to the tested steel the equation was
proposed:18, 19

 	=
⎛
⎝
⎜
⎞
⎠
⎟
b
k TG
D
2
2
B
(7)
with: 
 – creep rate, b – Burgers vector, kB – Boltzmann
constant, T/K – temperature, G – shear modulus, 	 –
acting stress, D – diffusion coefficient and  – density
of mobile dislocations.
A better fit to experimental values of creep rate was
found by modification of equation (7) with introduction
of a constant (A) accounting also for particles spacing, a
stress exponent n > 2, the rationalisation of the stress 	
with the shear modulus and yield stress at creep tempera-
ture and the threshold stress 	th, below which, theoreti-
cally, no creep could occur20:


	 	
=
⎛
⎝
⎜
⎞
⎠
⎟
−⎛
⎝
⎜ ⎞
⎠
⎟A
DGb
k T G
n
B
th
(8)
In the detachment concept of interaction of particles
and mobile dislocations the particle size is included, as
parameter of a real microstructure:21, 22, 23, 24
E Gb r k= − −
⎛
⎝
⎜
⎞
⎠
⎟
⎡
⎣⎢
⎤
⎦⎥
2
3 2
1 1( )
/
	
	 d
(9)
 exp


=
⎛
⎝
⎜
⎞
⎠
⎟ ⋅ −
⎛
⎝
⎜
⎞
⎠
⎟
6
k T
E
k TB B
(10)
with E – creep activation energy, 	d – detachment stress,
r – average particles size and the relaxation parameter k
= Tp/Tm, with Tm – dislocation line energy in the matrix
and T2 – the dislocation line energy decreased by the at-
traction force precipitate – dislocation. The value of the
parameter k is (0 < k < 1).
Theoretically, creep and self diffusion activation en-
ergies for pure  iron are about 300 kJ/mol.25 On the
base of lifetime of specimens tested at different tempera-
tures,26 creep activation energies of about 600 kJ/mol
were calculated.27,28 In ref.29 it is suggested that the dif-
ference may be related partially to the increase of creep
rate due to the change of distribution of particles during
creep tests.
The differences in creep rate for 0.18C-11.7Cr-
1Mo-0.29V tempered for 672 h and 1356 h at 800 °C
calculated according to equation (7) and experimental
creep rate were of 1.53 and 2.26 times, while it was of
several sizes of magnitudes greater if calculated accord-
ing to equation (9). This indicates as unreliable the de-
tachment model of interaction of a mobile dislocation
and carbide particles in creep resistant steels with the
microstructure of tempered martensite.29 In all quoted
equations it is assumed that the distribution of particles
in ferrite is uniform. However, in steels the particles dis-
D. A. SKOBIR BALANTI^ et al.: EFFECT OF CHANGE OF CARBIDE PARTICLES SPACING ...
556 Materiali in tehnologije / Materials and technology 45 (2011) 6,
tribution is not uniform, since by tempering of
martensite carbide particles are precipitated at grain
boundaries and subboundaries and in the interior of
grains and the coarsening rate is greater for particles sit-
uated at grain boundaries, where the diffusion rate is
greater.
In investigations on 0.18C-11.7Cr-1Mo-0.29V steel
tempered for 2 h to 1356 h at 800 °C 30 it was found that
the average M23C6 particles size (d) increased with tem-
pering time (t) as d3 = 1.48 × 10–26 × t m3 s–1, which is
much slower than the rate of decrease of the number of
carbide stringers of ns = 3.15 × 10–17 m2 s–1. The differ-
ence is related to the difference in volume and boundary
diffusion rate. Accordingly, the delaying effect of string-
ers on creep rate, which is very strong by sufficient
stringers density, is ended relatively fast and by a critical
stringers density of about 0.5 × 102 mm–2 the creep rate
is increased for about one order of magnitude (Figure 1).
The particle spacing () can be calculated from the
relation18:
 = (4d/ · f1/3) (11)
with f – volume share of carbides in the investigated
steel.
Assuming that carbide particles are uniformly distrib-
uted in the micrograph and situated in the centre of a
square with the side equal to particles spacing, the spac-
ing could be deduced without prior assessment of parti-
cles size and knowledge of the volume fraction of the
carbide phases with the relation:
 = (F/N)1/2 (12)
with F – as surface of a micrograph with N carbide par-
ticles.
As mentioned earlier, the creep rate for the 0.18C-
11.7Cr-1Mo-0.29V steel was calculated using equation
(7) for two points in Figure 1 below the critical stringers
density using the particles spacing deduced according to
relation (11), the calculated volume of M23C6 carbide
content and experimentally assessed average particles
size.
Specimens of a 0.18C-11Cr-0.94Mo-0.31V (steel a)
and 0.10C-7.9Cr-0.98Mo-0.23V-0.11Nb (steel b) steels
with microstructure in Figures 2, 3 and 4 were cut out
D. A. SKOBIR BALANTI^ et al.: EFFECT OF CHANGE OF CARBIDE PARTICLES SPACING ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 557
Figure 4: Microstructure of steel (b) after 1 year of tempering at 650
°C
Slika 4: Mikrostruktura jekla (b) po enem letu `arjenja na 650 °C
Figure 1: 0.18C-11.7Cr-1Mo-0.29V steel. Dependence of accelerated
creep rate on the number of grain boundary stringers of carbide parti-
cles in tempered martensite29
Slika 1: Jeklo 0.18C-11.7Cr-1Mo-0.29V. Odvisnost pospe{ene hitrosti
lezenja od {tevila nizov karbidnih izlo~kov po mejah zrn v popu{~e-
nem martenzitu29
Figure 2: Initial microstructure of steel (a)
Slika 2: Izhodna mikrostruktura jekla (a)
Figure 3: Initial microstructure of steel (b)
Slika 3: Izhodna mikrostruktura jekla (b)
from industrial tubes and tempered for up to 8760 h (1
year) at 650 °C. 31 For both steels the accelerated creep
rate was determined with 100 h static test at 580 °C and
the load of 170 MPa, as for the earlier mentioned steel.30
For the steel (a), the particles spacing was calculated us-
ing equation (11) and particles size assessed on SE mi-
crographs at magnification 104 and 2 × 104 times, while
the particles spacing for steels (b) and (c) was calculated
using equation (12) and particles counting on SE micro-
graphs at magnification of 5 × 103 times, where only par-
ticles with size of about 0.1 μm (about 20 nm) and more
were discernible, sufficiently clearly.
In Table 1 particles spacings, calculated as well as
experimental creep rates are given.
By steel (a) and (b) the ratio of experimental and cal-
culated creep rate is above 1 for both tempering tempera-
ture, while it is below 1 for the steel (c) tempered at 650
°C. The different ratio of calculated and experimental
rate for the steel (c) may be explained assuming that af-
ter both tempering temperatures by steels (a) and (b) on
used SE micrographs the great majority of both kinds of
carbide particles (M23C6 and VC) was visible, while the
resolution of micrographs of steel (c) was too small for
NbC particles. This explanation is supported by the dif-
ferent stability of carbides, since after tempering at 800
°C in steel (a) consisted only of M23C6 particles, in steel
(b) at 650 °C of M23C6 and VC and in steel (c) of M23C6,
VC and NbC particles. Data in Table 1 indicate that the
difference between calculated and experimental creep
rate is greater after longer tempering time at both tem-
peratures and independent of particles spacing. As the
creep rate is related to the density of particles stringers, it
is concluded, that also by longer tempering at 650 °C
creep rate was changed by decrease of the effect of
stringers of particles.
3 CONCLUSIONS
The following conclusions on the effect of changes in
microstructure, i.e. changes in spacings and distribution
of carbide particles, are based on data from quoted refer-
ences and on results of investigations of three creep re-
sistant steels.
• a simple method for assessment of particles spacing
was devised based on the assumption that all particles
in a micrographs are situated in the centre of a square
with the side equal to particles spacing;
• the creep rate delaying effect of stringers of carbide
particles at grain and subgrain boundaries is signifi-
cantly stronger than the effect of particles distributed
uniformly in the grains interior;
• when by tempering of the steel the density of parti-
cles stringers (number of stringers per unity of the
examined surface) is diminished rapidly below a crit-
ical level, the creep rate is increased strongly, for
about one order of magnitude, in case of tempering
the 0.18C-11.7Cr-1Mo-0.29V steel at 800 °C;
• the creep rates determined experimentally and calcu-
lated using the Ashby equation with the stress expo-
nent n = 2 agree acceptably and differences of both
rates depend also on the distribution of particles. The
effect of distribution of particles was confirmed for
steels with the microstructure with only M23C6 as
well as microstructures with M23C6, VC and NbC
particles.
• at 650 °C the effect of tempering time on change of
distribution of particles is much lower than at 800 °C.
In the first case the creep rate was increased 12.1
times after 1356 h of tempering, while at 650 °C, the
creep rate was increased for 2.14 times for the CrV
steel and 1.38 times for the CrVNb steel after 8760 h
of tempering. Thus the processes occurring in
microstructure by tempering at 800 °C decrease the
CrV creep resistance for about 36 times faster for the
CrV steel and 57 times for the CrVNb steel than pro-
cesses at the for 150 °C lower temperature of 650 °C.
4 REFERENCES
1 K. H. Mayer, F. Masuyama: The development of creep resistant
steels; Ed. F. Abe, T-U. Kern, R. Viswanathan: Creep resistant steels,
Woodhead Publ. LTD., Cambridge, England, (2008), 15–77
2 F. Abe: Strengthening mechanisms in steel for creep and creep rup-
ture; Ed. F. Abe, T-U. Kern, R. Viswanathan: Creep resistant steels,
Woodhead Publ. LTD., Cambridge, England, (2008), 279–304
3 K. Yamamoto, Y. Kimura, Y Mishima: ISIJ Intern. 42 (2003),
1253–1259
D. A. SKOBIR BALANTI^ et al.: EFFECT OF CHANGE OF CARBIDE PARTICLES SPACING ...
558 Materiali in tehnologije / Materials and technology 45 (2011) 6,
Table 1: Experimental and calculated creep rate for specimens of steels a) 0.18C-11.7Cr-1Mo-0.29V, b) 0.2C-11Cr-0.94Mo-0.31V and c)
0.10C-7.9Cr-0.98Mo-0.23V-0.11Nb tempered for different times at 800 °C and 650 °C. Test temperature 580 °C, stress 170 MPa and time 100 h.
Tabela 1: Eksperimentalne in izra~unane vrednosti hitrosti lezenja za vzorce jekel a) 0.18C-11.7Cr-1Mo-0.29V, b) 0.2C-11Cr-0.94Mo-0.31V in c)
0.10C-7.9Cr-0.98Mo-0.23V-0.11Nb, `arjene razli~no dolgo pri 800 °C in 650 °C. Temperatura presku{anja 580 °C, napetost 170 MPa in ~as 100
h.
Tempering Particles spacing (10–6 m)
Creep rate (s–1)
experimental calculated Exp./cal.
800 °C, 672 h, a 1.24 12.2 × 10–8 7.93 × 10–8 1.53
800 °C, 1356 h, a 1.50 21.8 × 10–8 9.62 × 10–8 2.26
650 °C, 2h, b 0.53 4.76 × 10–8 3.39 × 10–8 1.41
650 °C, 8760 h, b 0.61 10.2 × 10–8 3.90 × 10–8 2.61
650 °C, 2 h, c 0.507 1.47 × 10–8 3.24 × 10–8 0.45
650 °C, 8760 h, c 0.508 2.03 × 10–8 3.25 × 10–8 0.62
4 A. Aghajani, F. Richter, C. Somsen, S. G. Fries, I. Steinbach, G.,
Eggeler: Scripta Mater., 61 (2009), 1068–1071
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(2006), 147–153
10 J. W. Martin, R. D. Doherty, B. Cantor: Stability of microstructure in
metallic systems; Ed. Cambridge Un. press, U.K., 1997. p. 242 and
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11 K. Maruyama, K. Sawada, J. Koike: ISIJ Intern.41 (2001), 641–653.
Fundamental aspects of creep deformation and deformation mecha-
nism map; Ed. F. Abe, T-U. Kern, R. Viswanathan: Creep resistant
steels, Woodhead Publ. LTD., Cambridge, England, (2008), 265–278
12 F. Vodopivec, D. A. Skobir, B. Zuzek, M. Jenko: ISIJ Intern., Sub-
mitted for publication
13 W. Blum: Mechanisms of creep deformation in steel; in F. Abe, T-U.
Kern, R. Wiswanathan: Creep resistant Steels, Woodhead publ.,
Cambridge, 2008, 366–400
14 J. Hald, L. Korcakova: ISIJ Intern.43 (2003), 420–427
15 A. Gustafson, J. Agren: ISIJ Intern. 41 (2001), 356–360
16 K. Maruyama: Fundamental aspects of creep deformation and defor-
mation mechanism map; Ed. F. Abe, T-U. Kern, R. Viswanathan:
Creep resistant steels, Woodhead Publ. LTD., Cambridge, England,
(2008), 265–278
17 M. Taneike, M. Kondo, T. Morimoto: ISIJ Intern. 41 (2001), S111–S
115
18 E. Hornbogen: Einfluss von Teilchen einer zweiter Phase aus das
Zeitverhalten; W. Dahl, W. Pitch: Festigkeits- und Bruchverhalten
bei höheren Temperaturen, Verl. Stahleisen, Düsseldorf, (1980),
31–52
19 M.F. Asby: Proc. Sec. Int. Conf. On Strength of Metals and Alloys,
(1970), Am. Soc. Metals, ASM, Metals Park, Ohio, Ca, 507. Loc.
Cit. ref. 15
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21 J.H. Schröder, E. Artz: Scripta metall., 19 (1985), 1129–1134
22 J. Rösler, E. Artz: Acta metal., 38 (1990), 671–683
23 E. Artz, D.S. Wilkinson: Acta metall., 34 (1986), 1893–1898
24 E. Artz, J. Rösler: Acta metal., 36 (1988), 1053–1060
25 R. W. K. Honeycombe: The Plastic deformation of Metals, 2nd Ed.,
E. Arnold, 1985, 372
26 J. ^adek, V. [ustek, M. Pahutová: Mat. Sci. Eng. A225, (1997),
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638–642
28 B. Ule, A. Nagode: Mat. Sci. Techn. 23 (2007), 1367–1374
29 F. Vodopivec, J. Vojvodi~-Tuma, B. [u{tar{i~, R. Celin, M. Jenko:
Mat. Sci.Techn. 27 (2011), 937–942
30 D. A. Skobir, F. Vodopivec, L. Kosec, M. Jenko, J.Vojvodi~-Tuma:
Steel Res. Int. 75 (2004), 196–202
31 J. Vojvodi~-Gvardjan~i~, D. Kmeti~, R. Celin, B. Arzen{ek, F. Vodo-
pivec: Report NCRI 377/2007, Institute of Metals and Technology,
Ljubljana, Slovenia
D. A. SKOBIR BALANTI^ et al.: EFFECT OF CHANGE OF CARBIDE PARTICLES SPACING ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 559

LJ. TIHA^EK [OJI] et al.: STRESS-STRAIN ANALYSIS OF AN ABUTMENT TOOTH ...
STRESS-STRAIN ANALYSIS OF AN ABUTMENT TOOTH
WITH REST SEAT PREPARED IN A COMPOSITE
RESTORATION
NAPETOSTNO-DEFORMACIJSKA ANALIZA OPORNEGA ZOBA Z
ZAPORNIM SEDE@EM, IZDELANA S KOMPOZITNIM
POPRAVILOM
Ljiljana Tiha~ek [oji}1, Aleksandra M Lemi}1, Dragoslav Stamenkovi}1,
Vojkan Lazi}1, Rebeka Rudolf2,1, Aleksandar Todorovi}1
1Faculty of stomatology, University of Belgrade, Dr. Suboti}a 8, 11000 Beograd, Serbia
2University of Maribor, Faculty of Mechanical Engineering, Smetanova 17, 2000 Maribor, Slovenia
saskam@eunet.rs
Prejem rokopisa – received:2011-02-02 ; sprejem za objavo – accepted for publication: 2011-04-06
Placing a composite restoration on abutments for the removable of partial dentures gives favorable aesthetic results with
minimal intervention. The objective of this paper is to analyze the stress distribution of a tooth with occlusal rest-seat
preparation in the composite and compare it to the biomechanical behavior of an intact tooth, assuming that the stress and strain
distribution throughout the intact tooth provides the control scenario. For this finite-element study two different models were
designed. The first model was the three-dimensional (3D) model of an intact tooth, and the other one was a 3D model of a tooth
with a composite restoration and an appropriate occlusal rest-seat preparation. Load stimulations were performed when the rest
was fully seated on its corresponding rest seat and abutment tooth in order to obtain data about the biomechanical behavior of
the abutment tooth compared to the intact tooth’s stress-distribution pattern. The results of our analyses are presented and
analyzed qualitatively. The occlusal loading effect along the sound tooth exhibits a wider high-stress area, localized in
correspondence with the occlusal enamel, than the restored teeth. This is due to the rigidity of the enamel. The reduction in the
stress values occurs in the composite restoration, which is less rigid. Its lower rigidity allows larger cusp movements. The
stress-distribution pattern of the restored tooth with the rest seat showed that introducing an occlusal restoration does not differ
from the intact tooth globally, but locally. Our findings indicate that the composite rest-seat restoration absorbs the loading, so
reducing the stresses inside the tooth’s structure.
Key words: composite rest seat, abutment tooth, stress distribution, finite element method (FEM)
Postavitev kompozitnega popravila opore za odstranljive dele zobovja daje dobre estetske rezultate z minimalno intervencijo.
Cilj tega ~lanka je analizirati porazdelitev napetosti na zobu z zapornim sede`em s kompozitno zaporo in jo primerjati z
biomehanskim vedenjem intaktnega zoba s predpostavko, da odlo~a razdelitev napetosti in deformacije intaktnemu zobu. Za to
{tudijo sta bila na podlagi kon~nih elementov razvita dva razli~na modela. Prvi je tridimenzionalen (#D) model intaktnega zoba,
drugi pa je 3D-model zoba s kompozitno restoracijo in ustrezno pripravo zapornega sede`a, da bi tako dobili podatke o
biomehanskem vedenju mosti~nega opornika in o porazdelitvi napetosti v intaktnem zobu. Rezutati so prikazanai in analizirani
kvalitativno. Zaporna obremenitev povzro~i {iroko podro~je visoke napetosti v zdravem zobu zaradi stika s sklenino in ne
restoriranega zoba. To je posledica velike togosti sklenine. Manj{e so napetosti v manj togi kompozitni restoraciji, kar omogo~a
ve~je premik opornega vrha. Porazdelitev napetosti restoriranega zoba z opornim sede`em ka`e, da uporaba zaporne restoracije
ne vpliva globalno, ampak lokalno. Ugotovitve ka`ejo, da popravilo zob z oporo absorbira obremenitev, ker zmanj{a napetosti v
strukturi zoba.
Klju~ne besede: kompozitna opora, oporni zob, porazdelitev napetosti, FEM
1 INTRODUCTION
The preparation of occlusal rest seats on the
supporting, corresponding surfaces of the abutment teeth
is widely recommended as a way of promoting axial
load1. A removable partial denture (RPD) designed and
manufactured in that manner fulfils the functional,
prophylactic and aesthetic demands placed upon it.
Occlusal rest-seat preparation demands cutting the
enamel tissue in order to achieve "spoon-shaped" depres-
sions with the proper dimensions.2 The integrity of the
remaining tooth structure is deteriorated from the
biomechanical point of view with a resultant change in
the intact tooth’s pattern of stress and strain distribution.
The change of the natural tooth’s biomechanical balance
could lead to increased potential for further trauma and
eventual loss of the remaining tooth’s structure.
Although many studies with different methodologies
have been implemented and performed, no clear cut and
clinically relevant conclusions were drawn for the
minimal dimensions of the cavity preparation that would
minimize the tooth-fracture potential when subjected to
occlusal loading. In particular, when an RPD abutment
tooth is considered where most occlusal forces are
distributed to the abutment from the occlusal rest to the
rest seats in the tooth/tooth- or tooth/mucosa-supported
RPD. When planning prosthetic restorations one is
facing the presence of two different biological tissues
and the need for an even distribution of the occlusal and
other forces on the periodontal tissue of the remaining
teeth, and in the mucoperiosteum on the edentulous
alveolar ridge. The stress distribution throughout an RPD
Materiali in tehnologije / Materials and technology 45 (2011) 6, 561–566 561
UDK 616.31:539.3 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)561(2011)
and the supporting tissues may be evaluated easily using
the finite-element method (FEM) 3.
In order to manufacture an RPD that fulfills both the
functional and prophylactic demands, a non-invasive
modality treatment for preparing the abutment tooth may
be suggested. It is a good option for preparing the
supporting dental tissues for receiving the elements of
the RPD. The minimal appropriate intervention in that
case could be placing a composite restoration, as stated
by Shimizu & Takahashi 4. Using a highly filled compo-
site resin for the restoration of an abutment preparation
for removable partial dentures gives favorable aesthetic
results with the minimum intervention.4
The discussed treatment modality highlights the
objective of this paper, which is to analyze the
stress-strain distribution of the tooth with an occlusal
rest-seat preparation in a composite filling and compare
it to the biomechanical behavior of an intact tooth,
assuming that the stress-strain distribution throughout
the intact tooth provide the control scenario.
2 MATERIALS AND METHODS
The problem of the biomechanical behavior of a
complex structure with irregular geometry such as a
tooth could be analyzed using the finite-element method
(FEM). Applications of the FEM are expanding rapidly,
not only in the field of engineering but in science
globally, especially as it appears that the FEM is a useful
and convenient method for solving problems related to
macrostructures, but also might give a precise insight
into the problems related to microstructures.5
For this study two different models were created. The
first model was a three-dimensional (3D) model of an
intact tooth – Model 1 (Figure 1), and the other was a
3D model of a tooth with a composite filling and an
appropriate occlusal rest-seat preparation – Model 2
(Figure 2). The 3D model of the second upper premolar
was generated by combining the inner and outer
geometry profiles obtained from literature data.6,7 The
mentioned morphologic details and dimensions were
used to define a series of planes at different levels with
the outlines of the tooth cross-section at each level.
(Figure 1) The distance between the sliced planes was
three slices in one mm, where over two hundred planes
were generated. As this was time-consuming data for the
software processing, the distance between the slices was
increased afterwards to 2 mm for the crown portion of
the model, and 3 mm for the root portion of the model.
The PRO-ENGINEER (Parametric Technology
Corporation / Needham Massachusetts) and SOLID
WORKS (Solid Works Corporation) software packages
were employed for these procedures. The next step was
to describe the relations between the planes, from where
the solid model was constructed by connecting these
contours. The dentine part of the tooth structure’s crown
and root portions were modeled separately from the
enamel shell. Afterwards, the enamel and dentine were
assembled into the final model of the intact tooth
(Figure 2).
Subsequently, a model of a tooth with a composite
filling and an appropriate occlusal rest seat was obtained,
where additional cutting planes were defined in order to
perform the cavity preparation. The occlusal rest pre-
paration’s size and location were chosen to conform to
the standard cavity-design recommendations for a rest
seat. The rest seat was 1.5 mm deep, occupying one half
of the mesio-distal dimensions of the tooth crown and
was approximately one third of the crown in the
bucco-lingual direction. The recommended dimensions
were adopted from literature recommendations.8 The
final FEM model of the restored tooth was realized by
LJ. TIHA^EK [OJI] et al.: STRESS-STRAIN ANALYSIS OF AN ABUTMENT TOOTH ...
562 Materiali in tehnologije / Materials and technology 45 (2011) 6, 561–566
Figure 2: Solid model of a tooth with rest-seat preparation in a
composite
Slika 2: Trdni model zoba z opornim sede`em iz kompozita
Figure 1: Solid model of an intact tooth
Slika 1: Trdni model intaktnega zoba
changing the element material properties in the zone of
the cavity preparation.
Table 1: Young’s modulus and Poisson’s ratio of the materials used in
this study9,10
Tabela 1: Youngov modul in Poissonovo razmerje za materiale,
uporabljene v tej raziskavi9,10
Young’s modulus Poisson’s ratio
Enamel 84.1 GPa 0.33
Dentin 14.7 GPa 0.31
Composite resin 10.0 GPa 0.30
Co-Cr alloy 154.0 GPa 0.33
All the materials were considered to be isotropic and
homogeneous and were assigned the appropriate physi-
cal properties according to literature data9,10.The material
properties are listed in Table 1. The dental pulp was
modeled as a void because of its negligible stiffness and
strength.11 The cement was not included in the models
due to its small dimensions and because it had no influ-
ence on the analysis.12 Perfect bonding was considered
between the enamel and the composite.
Both models were constrained at the top, 2 mm from
the cement-enamel junction, representing the normal
level of the alveolar crest and, in that way, the boundary
conditions in this stress-strain analysis were defined.
Therefore, the influence of the periodontal ligament
(PDL) was not involved in the study.
Creating the finite-element mesh (pre-processing) for
the described models was performed using the NA-
STRAN program (Noran Engineering Inc / Westminster
Ca), where all the subsequent procedures (processing
and post-processing) were done. The intact tooth-model
mesh consisted of 15 092 finite elements with 23 300
nodes, while the mesh for the second model had 15 089
finite elements with 23 334 nodes. Each model was
meshed by structurally solid elements defined by 10
nodes and having three degrees of freedom in tetrahedral
bodies.
The loads were determined with an emphasis on load
intensity, the direction and the point of the loading.
Because of the variety of occlusal forces (differing
among individuals, types of food chewing, conditions of
occlusion) this study adopted vertical loading with 250 N
in intensity. The existence of possible horizontal forces
were ignored, where the RPD was assumed to be a rigid
and stable appliance resisting the horizontal component
of the masticatory forces13. As the analysis was consi-
dered linear in nature, sliding and friction phenomena
that might occur between a rest and an abutment tooth
were also ignored. Load stimulations were performed
when the rest was fully seated on its corresponding rest
seat and abutment tooth. The loads for both models were
applied vertically at right angles to the inner aspects of
the cusp slopes, away from the cusp tip. The simulated
direction and intensity of the loads represent the loading
pattern found in the centric occlusion, as stated by Rees
and Jacobsen 16.
3 RESULTS
Based on the assumptions involved in the study and
the fact that computer simulations simplified the real
problems, the results of the study might be different to
the values of stresses encountered by teeth in real
situations. Therefore, the results were presented and
considered qualitatively, not quantitatively, in order to
offer more insight into the general influence of the
prosthetic devices placed on teeth.
The results of the study are presented graphically as
maps of the stress distribution within both investigated
models. The total displacement (translation), maximum
principal stress 	max, and minimum principal stress 	min
were evaluated for each of the models.
The total displacement of the intact tooth after
occlusal loading is shown in Figure 3. The largest
displacement values are recorded at the cusp tip, due to
tooth structure deformation encountered as a result of
loading. Obviously, a lot of deformation happens in the
points where the load is applied and therefore the
greatest displacement values are observed there. Moving
away from the loading point along the longitudinal axis
of the tooth in an apical direction, the displacement
decreases. This is probably the mechanism for the
occlusal load amortization within the intact tooth’s
structure. Such findings are partly recognized as a
consequence of the applied boundary conditions.
The values for 	max and 	min are also found to be the
highest at the occlusal portion of the tooth. The con-
centration of stresses is higher at the location of the
loading. These stresses rapidly decrease in the occluso-
gingival direction. Close to the cement-enamel junction
(CEJ) the stresses again increase and concentrate at that
location. The pattern of stress distribution throughout the
rest of the tooth structure shows a reasonably symme-
trical distribution with the exception of the occlusal
surface and the location close to the CEJ. (Figures 4 and
5)
As shown in Figure 6, the total translational dis-
placement of the tooth with rest seat preparation is
greatest at the cusp tip, as a result of the whole structure
deformation. The occlusal loading of the model with
rest-seat preparation induced a large deformation of the
tooth structure. The observed total displacement values
are similar to the values obtained for the intact tooth
model. Similarly, as with the intact tooth model, the total
displacement values are the largest at the location of the
loading where great deformation happens. The displace-
ment decreases in an apical direction as the distance
from the loading location increases.
The general trend of stress distribution through the
tooth with a rest-seat preparation is similar to that of an
intact tooth, but some regional variations were observed.
The concentration of the stresses 	max and 	min is
higher occlusally at the location of the loading, but very
high stresses are concentrated at the bottom (pulpal
LJ. TIHA^EK [OJI] et al.: STRESS-STRAIN ANALYSIS OF AN ABUTMENT TOOTH ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 561–566 563
LJ. TIHA^EK [OJI] et al.: STRESS-STRAIN ANALYSIS OF AN ABUTMENT TOOTH ...
564 Materiali in tehnologije / Materials and technology 45 (2011) 6, 561–566
Figure 8: Distribution of the minimum principal stress of the tooth
with the occlusal rest-seat preparation in the composite restoration
Slika 8: Porazdelitev najmanj{ih glavnih napetosti v zobu z opornim
sede`em s kompozitno restoracijo
Figure 5: Distribution of the minimum principal stress of the intact
tooth model
Slika 5: Porazdelitev najmanj{ih glavnih napetosti v modelu intakt-
nega zoba
Figure 4: Distribution of the maximum principal stress of the intact
tooth model
Slika 4: Porazdelitev najve~jih glavnih napetosti v modelu intaktnega
zoba
Figure 3: Total translational displacement of the intact tooth model
Slika 3: Totalni premik modela intaktnega zoba
Figure 7: Distribution of the maximum principal stress of the tooth
with the occlusal rest-seat preparation in the composite restoration
Slika 7: Porazdelitev najve~jih maksimalnih napetosti v zobu z zapor-
nim sede`em in kompozitno restoracijo
Figure 6: Total translational displacement of the tooth with occlusal
rest-seat preparation in a composite
Slika 6: Popoln premik zoba z opornim sede`em, izdelanim iz kom-
pozita
floor) of the rest-seat preparation (Figures 7 and 8). The
trend of a symmetrical stress distribution through the rest
of the tooth structure with decreasing intensity is also
observed in this model. Again, the stresses rapidly
increase close to the CEJ and reach their highest peak
concentration at the CEJ. The stresses 	max and 	min
resulting from the application of the occlusal load on the
tooth with rest-seat preparation have lower values than
those obtained after the intact tooth loading.
4 DISCUSSION
This investigation was focused on the mechanical
behavior of a critical system such as the second upper
premolar restored by a class II cavity preparation (where
an accordant amount of dentine and enamel tissue is lost
and the integrity of the structure is altered by placing an
occlusal rest).
The relative behavior of composite restoration and
the abutment tooth under the mastication load is shown,
although some assumptions were made in order to sim-
plify the calculations. Despite their intrinsic anisotropic
nature, dentine and enamel can be assumed to be homo-
geneous and isotropic because their anisotropy belongs
to the microscopic scale, whereas the tooth model is
macroscopic.14,15 Furthermore, all the materials were
considered to be elastic throughout the entire defor-
mation, which is a reasonable assumption for brittle
materials in non-failure conditions.16
From the mechanical point of view, considering
dentine as elastic and isotropic is an acceptable assump-
tion, while enamel does not show such behavioral
characteristics. Enamel is a more anisotropic material
and some authors have modeled the enamel as aniso-
tropic14,16 Although such enamel modeling was more
orthotropic, where the principal material axes coincided
with the direction of enamel prisms, they were assumed
to be perpendicular to the enamel-dentine junction.
Following the recommendations of Darendeliler 14 and
Versluis 15, materials which comprise the tooth structure
could be modeled as isotropic and homogeneous,14,18
especially because there is still no competent literature
data on dental structure inhomogeneity and anisotropy.
Considering the occlusal loading effect alone, the
sound tooth exhibits a wider high stress area, localized in
correspondence to the occlusal enamel, than the restored
teeth. This is due to the rigidity of the enamel. The
reduction in stress occurs in the composite restoration,
which is less rigid, and its lower rigidity allows greater
cusp movements. The composite restoration will only
cause deformation of the surrounding tooth structure
through a well-bonded interface. So, the average stress in
the entire structure is lower, but the stress values in the
buccal and lingual cusps are higher. However, placement
of a composite restoration with included a rest seat does
not deteriorate the tooth’s ability to withstand occlusal
loading. The tooth with a composite restoration was less
sensitive to occlusal loading than the intact one, probably
due to the cusp reinforcement achieved by dentine and
enamel bonding.
However, the concentration of stresses along the
pulpal floor must undoubtedly be considered. The
presence of cracks in the enamel introduced during
cavity preparation may be a primary factor contributing
to the stress concentration and eventual restoration
failure. Also, some variance in the enamel mineralization
between the surface and deep portions might interfere in
homogeneous stress distribution throughout the restored
tooth. The degree of enamel mineralization decreases as
one approaches the amelo-dentinal junction, with a
decreased elastic modulus19. Therefore, it seems that the
peak stresses found in the tooth model with composite
restoration with an included rest seat may be attributed to
different causes.
These preliminary considerations, however, do not
take into account setting composite shrinkage. These
findings correlate with the statement of Ausiello 20 that
more rigid composites lead to lower cusp movements
under occlusal loading, but exhibit a higher pre-loading
effect. A good composite for restoration has to balance
the two opposite effects. In this way, a low pre-load on
the cusps can be accepted in order to reach a sufficient
restoration rigidity.20 The size and configuration of the
composite restoration affects the amount of tooth defor-
mation due to resin polymerization shrinkage. In a small
restoration (Class I and small Class II), deformation
could hardly be visualized, which means that it was
consistently lower or close to the measurement reso-
lution.21 The cavity designed for the rest seat in the study
was of smaller dimensions than the standard Class II
restoration, allowing us to speculate that the obtained
deformation was a consequence of the applied loading
by the occlusal rest and may not be partly the result of
resin polymerization shrinkage. The study adopted the
absolute bonding between the tooth and composite resto-
ration, which may be the reason for the high stress
concentration found along the pulpal floor. An infinitely
rigid interface layer produces high stress areas all around
the tooth-restoration interface. Accompanied by the
polymerization shrinkage, stress and occlusal load stress
exists as a very complex biomechanical system. An
appropriate way to limit the intensity of all the stress
transmitted to the remaining tooth tissues under occlusal
loading by the rest would be the proper selection of the
adhesive layer. Ausiello22 stated that a substantially
thicker layer of a more flexible adhesive (lower elastic
modulus) would partially absorb the composite defor-
mation.22
5 CONCLUSION
Within the limitations of the finite-element simu-
lation study, the results suggest that a composite resto-
ration with a rest seat included reduces the generation of
LJ. TIHA^EK [OJI] et al.: STRESS-STRAIN ANALYSIS OF AN ABUTMENT TOOTH ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 561–566 565
stresses inside the tooth structure. Findings indicate that
a composite restoration absorbs the loading and, due to
its resilient nature, acts as a cushion beneath the occlusal
rest. The simplifications used in the study have also been
shown to affect the results. The detailed modeling of
dental hard-tissue anisotropy, along with the more
precise modeling of the adhesive interface, should be
considered as a goal for any future study on this topic.
The potential of the system composite-occlusal rest
seat is that it represents the minimum intervention
procedure and acts as an absorbent of occlusal loading.
Furthermore, the eventual use of a composite restoration
may be suggested as a way of preparing abutments for
receiving elements of the RPD. However, clinical trials
are required to ensure that a composite restoration with
rest seat included can survive under long-term clinical
conditions.
6 REFERENCES
1 O’Grady J, Sheriff M, Likeman P. A finite element analysis of a
mandibular canine as a denture abutment. European Journal of
Prosthodontics and Restorative Dentistry, 4 (1996) 3, 117–121
2 Sato Y, Shindoi N, Koretake K, Hosokawa R. The effect of occlusal
rest size and shape on yield strength. Journal of Prosthetic Dentistry,
89 (2003), 503–507
3 Todorovic A, Radovic K, Grbovic A, Rudolf R, Maksimovic I,
Stamenkovic D. Stress analysis of a unilateral complex partial
denture using the finite-element method. Mater. Tehnol., 44 (2010) 1,
41–47
4 Shimizu H, Takahashi Y. Highly filled composite partial coverage
restorations with lingual rest seats and guide planes for removable
partial dentures. Journal of Prosthetic Dentistry, 99 (2008),73–74
5 Avdiaj S, Setina J, Syla N. Modeling of the piezoelectric effect using
the finite-element method (FEM.) Mater. Tehnol., 43 (2009) 6,
283–291
6 Ash MM, Nelson SJ. Wheeler’s Dental Anatomy, Physiology and
Occlusion. Eighth edition Elsevier Science, 2003, 230–238
7 Schillinburg HT, Jacobi R, Brackett SE. Fundamentals of Tooth
Preparations. Second printing Quintessence Publishing Co. 1991,
13–15
8 Biomaterials Properties Database, University of Michigan
Quintessence Publishing, 1996. Available from World WideWeb:
http://www.lib.umich.edu/libhome/Dentistry.lib/Dental_tables.html
9 Ensaff, H., D. M O’Doherty and P. H. Jacobsen. The influence of the
restoration-tooth in light cured composite restoration: a finite
element analysis. Biomaterials, 22 (2001), 3097–3103
10 Rubin C., Krishnamurthy N., Capilouto E., Yi H. Stress analysis of
the human tooth using a three-dimensional finite element model.
Journal of Dental Research, 62 (1983), 82–86 2
11 Tanaka M, Naito T, Yokota M, Kohno M. Finite element analysis of
the possible mechanism of cervical lesion formation by occlusal
force. J Oral Rehabilitation, 30 (2003), 60–67
12 Rees JS. The effect of variation in occlusal loading on the develop-
ment of abfraction lesions: a finite element study. J Oral Rehabili-
tation, 29 (2002), 188–193
13 Muraki H, Wakabayashi N, Park I, Ohyama T. Finite element contact
stress analysis of the RPD abutment tooth and periodontal ligament.
Journal of Dentistry, 32 (2004), 659–665
14 Darendeliler, S. Y., Alacam, T., Yaman, Y., Analysis of stress distri-
bution in a maxillary central incisor subjected to various post and
core applications. Journal of Endodontics, 24 (1998), 107–111
15 Versluis, A., Douglas, W. H., Cross, M, Sakaguchi, R. L. Does an
incremental filling technique reduce polymerization shrinkage
stresses? Journal of Dental Research, 3 (1996), 871–878
16 Rees, J. S., Jacobsen, P. H. Modelling the effects of enamel aniso-
tropy with the finite element method. Journal of Oral Rehabilitation,
22 (1995), 451–454
17 Yettram A. L., Wright K. W. J. Pickard H. M. Finite element stress
analysis of the crowns of normal and restored teeth Journal of Dental
Research, 55 (1976) 6, 1004–1011
18 Versluis A. Does an incremental filling technique reduce polyme-
risation shrinkage stresses? Journal of Dental Research, 75 (1996),
871–878
19 Meredith N, Sheriff DJ, Setchell DJ. Measurement of the micro-
hardness and young’s modulus of human enamel and dentine using
an indentation technique. Archives of Oral Biology, 41 (1996),
539–541
20 Ausiello P., Apicella A. Davidson C. L., Rengo S. 3D-finite element
analyses of cusp movements in a human upper premolar, restored
with adhesive resin-based composites Journal of Biomechanics, 34
(2001), 1269–1277
21 Tantbirojn D., Versluis A., Pintado M. R., DeLong R., Douglas W H.
Tooth deformation patterns in molars after composite restoration.
Dental Materials, 20 (2004), 535–542
22 Ausiello P., Apicella A. Davidson CL. Effect of adhesive layer
properties on stress distribution in composite restoration – a 3D finite
element analysis. Dental Materials, 18 (2002), 295–303
LJ. TIHA^EK [OJI] et al.: STRESS-STRAIN ANALYSIS OF AN ABUTMENT TOOTH ...
566 Materiali in tehnologije / Materials and technology 45 (2011) 6, 561–566
V. KLEISNER et al.: IDENTIFICATION AND VERIFICATION OF THE COMPOSITE MATERIAL PARAMETERS ...
IDENTIFICATION AND VERIFICATION OF THE
COMPOSITE MATERIAL PARAMETERS FOR THE
LADEVÈZE DAMAGE MODEL
IDENTIFIKACIJA IN VERIFIKACIJA PARAMETROV
KOMPOZITNEGA MATERIALA ZA MODEL LADEVÈZE
Václav Kleisner, Robert Zem~ík, Tomá{ Kroupa
University of West Bohemia in Pilsen, Department of Mechanics, Univerzitní 22, 306 14, Plzeò, Czech Republic
kleisner@kme.zcu.cz
Prejem rokopisa – received:2011-02-01 ; sprejem za objavo – accepted for publication: 2011-04-14
In this investigation we examine the properties of a layered composite material and verify the Ladevèze material model
implemented in PAM-CRASH software. The complex material model incorporates plasticity, failure and damage mechanisms
and is suitable for dynamic phenomena, such as crash tests. The experimental tests were performed on appropriate laminated
specimens made from unidirectional, pre-impregnated, composite fiber (prepregs) – coupons with axially oriented fibers,
coupons with fibers at 45°, and ±45° cross-ply laminates. The tests included simple tensile tests to fracture and cyclic tensile
tests. Numerical models were created for the finite-element analysis using shell elements. A mathematical optimization was
then used to minimize the error between the experimental and numerical results in terms of load-displacement curves for all the
tested configurations by varying the material characteristics.
Keywords: composite, identification, carbon, fiber, epoxy, plasticity, experiment, finite-element analysis
Identifikacija lastnosti plastastega kompozitnega materiala in verifikacja modela Ladeveze za material s PAM-CRASH-
sofverom. Kompleksen model materiala vklju~uje plasti~nost, prelom in mehanizem po{kodbe ter je primeren za dinami~ne
fenomene kot preizkus trka. Preizkusi so bili izvr{eni na primernih laminatnih vzorcih, izdelanih iz enosmernih
predimpregniranih kompozitnih vlaken (prepreg) – kuponov z osno orientiranimi vlakni, kuponov z vlakni pod kotom 45° in
kri`nimi laminati ±45°. Preizkusi so obsegali enostavne raztr`ne in cikli~ne natezne preizkuse. Pripravljeni so bili numeri~ni
modeli za analizo po metodi kon~nih elementov z uporabo lupinastih elementov. Matemati~na optimizacija je bila nato
uporabljena za zmanj{anje napak med eksperimentalnimi in numeri~nimi rezultati s krivuljami obremenitev – pomik za vse
preizku{ene konfiguracije s spremembami karakteristik materiala.
Klju~ne besede: kompozit, identifikacija, ogljikova vlakna, epoksi, plasti~nost, preizkusi, kon~na elementna analiza
1 INTRODUCTION
Composite materials are modern materials with
advantageous strength- and stiffness-to-mass ratios com-
pared to classical materials, such as steel or aluminum 1,2.
Namely, the carbon-fiber-reinforced plastic composites
consisting of continuous carbon fibers and a matrix can
have similar or better strength than steel structures and
they can have similar or less weight than aluminum
structures. As their properties are highly oriented
(generally anisotropic), the greatest strength is achieved
in the direction of the fibers. This can be utilized
especially in the case of the design of components with
excessive loading in a specific direction.
Composite materials are increasingly used in the
aerospace and automotive industries for the reason
mentioned above. Numerical simulations help to design
the desired components or complex structures, including
the possibility to optimize the fiber orientations or
lay-ups. Nevertheless, it is important to know the correct
material parameters and to use the appropriate material
model. This material data must be obtained from
experimental measurements. An integral part of any
material model is the failure/damage prediction possi-
bility. Many material models have been proposed so far,
but none of them is perfect or universal 6. The basic
failure criteria, such as the maximum stress, maximum
strain and others, are not interactive criteria. This means
that there is no relation between the stress components in
different directions. In this respect, the so-called inte-
ractive criteria, such as Tsai-Wu 1, are more suitable for
crash simulations. On the other hand, the disadvantage is
that we cannot distinguish between the matrix and fiber
failure, which is important in an impact simulation. The
most recent failure criteria (the so-called direct mode
criteria), such as Puck 8 or LaRC 3, use the advantages of
both types 9.
The Ladevèze material model 5 in the PAM-CRASH
software 7 is implemented only for a multi-layered, thin
shell element and transient analysis (i.e., the explicit
code). It includes the following modes of failure of a
composite material: debonding, micro-cracking, delami-
nation, and fiber breaking. The Ladevèze damage model
also includes inelastic material deformations caused by
the matrix-dominated loading. The plasticity of the
matrix cannot be neglected in general and the effect is
best seen, for example, in the case of cyclic loading.
Materiali in tehnologije / Materials and technology 45 (2011) 6, 567–570 567
UDK 669.018.9:620.1/.2 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)567(2011)
2 MATERIAL AND DAMAGE MODELS
The constitutive relationship for materials with a
linear response is usually written in the form of the
extended Hooke’s law 1. The constitutive relationship of
the Ladevèze material model can be written with similar
formulae, except that elastic constants are herein
modified by additional damage parameters or functions
4,5. The crucial relations are summarized in Table 1. The
superscript 0 denotes the initial values (damage free) of
the material constants. The quantities d11, d22 and d22
represent the fiber damage in tension, matrix damage,
and fiber-matrix debonding damage, respectively. The
effect of d12 is shown in the relation of the actual (G12)
and initial (G012) values of the shear moduli. The shear
damage function Y12 is derived from the strain energy Ed
for an anisotropic material, where YC and Y0 are the
critical shear damage limit and the initial shear damage
threshold, respectively. The parameter YR represents the
shear failure.
Another important improvement to the composite
material model is obtained by the inclusion of the matrix
plasticity behavior. This is incorporated by changing the
yield stress during the cyclic loading. The yield stress is
given by R(
P), which is a function of the initial yield
stress R0, the plastic deformation 
P and the hardening
coefficients , m. This represents a power-law approxi-
mation of the experimental curve.
The fiber tensile damage (longitudinal damage) is
characterized by the initial (
i11) and ultimate (
u11) fiber
tensile damage strains.
3 EXPERIMENT AND SIMULATIONS
In this study, laminated composite coupons made of
HexPly 913C prepregs with Tenax HTS 5631 carbon
fibers are tested (see Figures 1–3). The material
characteristics needed for the numerical models are
obtained from the experimental data. The detailed
description of the measurement can be found in 7. It
consists of three types of tests:
• simple tensile test on [0]8 laminates,
• simple tensile test with load/unload cycles on [±45]2S
laminates,
• simple tensile test on [45]8 laminates.
Simple [0] tensile test
The tensile test was conducted on UD composite
coupons with the [0]8 fiber composition (see Figure 1).
The coupons were loaded by displacement (speed
1 mm/min) until rupture. The force–displacement curve
was measured, see Figure 4.
The initial Young’s modulus E011, the initial fiber
failure value 
i11 and the critical fiber failure value 
u11
were assessed from the data obtained using Hooke’s law.
The averaged experimental results were used directly
in the material model within the corresponding numeri-
cal simulation. The results of the simulation are in a good
V. KLEISNER et al.: IDENTIFICATION AND VERIFICATION OF THE COMPOSITE MATERIAL PARAMETERS ...
568 Materiali in tehnologije / Materials and technology 45 (2011) 6, 567–570
Tabela 1: Relacije modela Ladevèze za lupinaste elemente4,5
Table 1: Ladevèze model relations for shell elements 4,5
Figure 3: Fractured [45]8 specimen
Slika 3: Prelomljen vzorec [45]8
Figure 2: Fractured [±45]2S specimen. The position and orientation of
the cracks is emphasized
Slika 2: Prelomljen vzorec [±45]2s. Poudarjena sta polo`aj in
orientacija razpok
Figure 1: Fractured [0]8 specimen
Slika 1: Prelomljen vzorec [0]s
agreement with the experimental data (see Figure 4).
The constants 
i11 and 
u11 have similar values as the
whole cross-section ruptured at the same time.
Cyclic [±45]2S tension test
The composite coupons (Figure 2) were loaded by a
cyclic loading – 6 cycles (load/unload) with increasing
load amplitude (700 N, 800 N, 900 N, 1000 N, 1100 N
and 1200 N) and the force–displacement curves were
obtained. The nonlinear behavior and the plasticity of the
composite material can be clearly seen from the results.
This phenomenon is given by the plastic behavior of the
matrix or fiber-matrix interface. The stress and strain
vectors in the principal material directions (the fiber
direction and the transverse fiber direction) must be
calculated from the experimental data using the relations
for the stress/strain transformation for each lamina.
Consequently, it is possible to calculate the actual shear
modulus G12.
The material parameters responsible for the nonlinear
response of the numerical model were optimized using
the PAM-OPT tool to minimize the error between the
simulated and experimental data. Relatively good
agreement between the experimental and the simulated
curves was obtained; however, the maximum force in
this case was not correctly predicted.
Simple [45]8 tension test
For the validation of the shear failure parameter YR a
simple tension test on the [45]8 laminate was performed
(Figure 3). This parameter will ensure that the material
fails when the load exceeds a certain limit Figure 5.
Recalculation of the cyclic test with the new shear
failure parameter in the material model led to a
significant improvement of the correlation with the
experimental data. The comparison of the resulting
curves is shown in Figure 6. The resulting values of all
the parameters of the Ladevèze model used are
summarized in Table 2.
4 CONCLUSION
The combination of three types of experimental
measurements and numerical simulations in the
finite-element code PAM-CRASH was performed. A
V. KLEISNER et al.: IDENTIFICATION AND VERIFICATION OF THE COMPOSITE MATERIAL PARAMETERS ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 567–570 569
Figure 6: Load–displacement curves from [±45]2S test
Slika 6: Krivulji obremenitev – pomik za preizkus [±45]2S
Tabela 2: Identificirane karakteristike materiala
Table 2: Identified material characteristics
Figure 5: Load–displacement curves from the [45]8 test
Slika 5: Krivulji obremenitev – pomik za preizkus [45]8
Figure 4: Load–displacement curves from the [0]8 test
Slika 4: Krivulji obremenitev – pomik za preizkus [0]8
mathematical optimization was used to obtain the
parameters of the used Ladevèze material model that
incorporates plasticity, damage and failure. The resulting
comparison of the numerical and experimental data in
terms of load–displacement curves shows a very good
agreement.
In future work, a similar investigation will be
performed on textile composites. The applicability of the
Ladevèze model will thus be tested on a material with
even more complex behavior.
Acknowledgement
The work has been supported by the research project
GA CR 101/08/0299 and the research project
GA CR 101/08/P091.
5 REFERENCES
1 Berthelot, J. M. (1999). Composite materials, Springer, New York
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als, Oxford University Press, Inc. (1994)
3 Dávila, C. G., Camanho, P. P. (2003). Failure Criteria for FRP
Laminates in Plane Stress, NASA/TM-2003-212663 report, NASA
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oxy non-crimp fabric composites including effects of fabric
pre-shear, Composites Part A: Applied Science and Manufacturing,
37 (2006), 11, 1983–2001
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for laminated composites, Compos. Sci. Technol., 43 (1992) 3,
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panels, Journal of Composite Materials, 42 (2008) 1, 25–44
7 PAM-CRASH 2007. Solver Notes Manual. ESI-Group, Paris
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means of physically based phenomenological models. Composites
Science and Technology, 58 (1998), 1045–1067
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Exercise, Composites Science and Technology, 64 (2004), 589–604
V. KLEISNER et al.: IDENTIFICATION AND VERIFICATION OF THE COMPOSITE MATERIAL PARAMETERS ...
570 Materiali in tehnologije / Materials and technology 45 (2011) 6, 567–570
M. HOSSEINALI BEYGI et al.: EVALUATION OF THE STRENGTH VARIATION OF NORMAL ...
EVALUATION OF THE STRENGTH VARIATION OF
NORMAL AND LIGHTWEIGHT SELF-COMPACTING
CONCRETE IN FULL SCALE WALLS
OCENA VARIACIJE TRDNOSTI NORMALNEGA IN LAHKEGA
VIBRIRANEGA BETONA V POLNIH STENAH
M. M. Ranjbar1, M. Hosseinali Beygi2, I. M. Nikbin3 , M. Rezvani4, A. Barari5
1Department of Civil Engineering, Guilan University, Rasht, Iran
2Department of Civil Engineering, Babol University of Technology, Babol, Iran
3Department of Civil Engineering Islamic Azad University, Rasht Branch, Rasht
4Institute of Material Science, University of Duisburg-Essen, Universitätstrasse 15, 45141 Essen, Germany
5Department of Civil Engineering, Aalborg University, Sohngardsholmsvej 57, 9000 Aalborg, Aalborg, Denmark
ab@civil.aau.dk, amin78404@yahoo.com
Prejem rokopisa – received: 2011-03-26; sprejem za objavo – accepted for publication: 2011-06-10
The strength of cast concrete along the height and length of large structural members might vary due to inadequate compaction,
segregation, bleeding, head pressure, and material type. The distribution of strength within a series of full scale reinforced
concrete walls was examined using non-destructive testing. Self-compacting concrete (SCC) and lightweight self-compacting
concrete (LWSCC) with different admixtures were tested and compared with normal concrete (NC). The results were also
compared with results for standard cubic samples. The results demonstrate the effect of concrete type on the in situ strength
variation and the relationship to the strength of standard cube samples. Investigation of the strength variation along the height of
the wall showed that SCC mixes had better strength uniformity and that the NC mix had the greatest strength variation. There
were no significant strength differences between mixtures along the length of the walls. Furthermore, different admixture
replacements did not have a meaningful effect on the strength distribution.
Keywords: strength variation, self-compacting concrete, lightweight aggregate, ultrasonic pulse velocities, compressive
strength, structural walls.
Trdnost litega betona po vi{ini in dol`ini ve~jih gradbenih elementov lahko variira zaradi neprimernega vibriranja, segregacije,
iztekanja (solzenja), pritiska glave in vrste materiala. Porazdelitev trdnosti v seriji `elezobetonskih cementnih sten je bila
preiskana z uporabo neporu{nih preizkusov. Vibrirani beton (SCC) in lahek vibrirani beton (LWSCC) iz razli~nih zmesi sta bila
primerjana z normalnim betonom (NC). Rezultati so bili primerjani tudi z rezultati za standardne kockaste vzorce. Rezultati
dokazujejo u~inek vrste betona na in situ variacijo trdnosti in odvisnost od trdnosti standardnih kockastih vzorcev. Raziskava
trdnosti vzdol` vi{ine stene je pokazala, da razli~ne zmesi ne vplivajo pomembno na porazdelitev trdnosti.
Klju~ne besede: variacija trdnosti, vibrirani beton, lahek agregat, hitrost ultrazvo~nega impulza, tla~na trdnost, zidovi
konstrukcij
1 INTRODUCTION
In the past two decades, self-compacting concrete
(SCC) has been developed to build concrete structures.
Practical applications of SCC vary widely and have been
accompanied by numerous research studies. SCC is a
flowable concrete that fills spaces, especially in sections
with highly congested reinforced members and restricted
shapes.1 SCC has less stringent working and safety
requirements because of its negligible vibration factor.
The weight of SCC is the main cause of its compaction.
Its significant advantages, such as favorable mechanical
properties and durability, have made SCC a
high-performance concrete. Most of the probable
bleeding and segregation can be reduced with
viscosity-modifying agents in SCC mixes. Consequently,
due to the elimination of any voids, the maximum
density, material strength and concrete-steel rebar
bonding are improved. Furthermore, SCC provides easier
transition and pumping. Hence, the casting process is
faster for huge structural members.1,2,3
Structural lightweight concrete has a lower density in
place compared to normal weight concrete. The concrete
mixture is made with a lightweight aggregate. The main
application of structural lightweight concrete is to
decrease the dead load of concrete structures such as
high-rise buildings and long-span bridges, which then
allows the structural designer to reduce the size of piers,
footings, walls and other load-bearing elements.4
Furthermore, it can decrease the applied dynamic loads,
such as earthquake forces, which are directly related to
the dead load of a structure. It has been shown that a
decrease in the dimensions of structural members
compensates for the higher cost of lightweight concrete.5
Variations in the strength of concrete in a structure
and control samples of a similar age are fundamentally
due to differences in curing conditions and compaction.
These factors also affect the strength variations within
the depth of the members. Variations in a concrete
supply are due to the variety of batching, materials,
placing and handling methods, which are often restricted
by the quality control of a production, such as the
Materiali in tehnologije / Materials and technology 45 (2011) 6, 571–577 571
UDK 666.97:620.181 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)571(2011)
compliance testing of control samples. These factors are
clearly not related to the member type involved but lead
to random in situ variations. Previous research has shown
positional differences along the height of concrete
members cast in a deep form, and the lowest strength
occurred at higher positions due to the concrete pressure
head, aggregate settlement, bleeding and pores, which
cause inadequate compaction. Another reason for this
type of variation is the upward movement of water
through concrete while the material is still plastic and the
fact that upper layers protect the lower layers of concrete
from rapid drying.6–9
Considering the nature of self-compacting concretes,
which have a high consistency and flow ability, the
homogeneity of these concretes in comparison to normal
concrete that is compacted traditionally has been
evaluated. Previous studies performed by Zou and
Khayat on reinforced concrete beams, columns at large
scales and un-reinforced walls relative to small scale
walls made of self-compacting and normal concretes
showed that the members made of self-compacting
concrete had higher homogeneity but some discrepancies
within and between them.10,11
On the other hand, in many situations, the standard
cube sample strength may not indicate the actual strength
of the concrete members. Therefore, the concrete
strength of full scale reinforced concrete walls in service
sites was evaluated. The ultrasonic pulse velocity test
was used to investigate the strength variation and
uniformity of in situ concrete strength.
2 EXPERIMENTAL PROGRAM
To study the homogeneity of normal weight self-
compacting concretes and lightweight self-compacting
concretes, five (2.0 × 2.0 × 0.2) m reinforced concrete
walls were cast. The strength distribution in the
structural walls and their variations with the normal
weight concrete (NC) were also studied. The walls
contained reinforcing steel 12 mm in diameter and 25 cm
in length. Steel reinforcement details for the walls are
given in Figure 1. This reinforcement configuration was
selected to assess the effect of steel bars on the
compaction of self-compacting concretes.
Concrete mix with a specified cube strength of 28
MPa was considered, and Type I Portland cement
conforming to BS12 was used throughout the tests. The
lightweight aggregate used in this research was Leca
with a 24 h water absorption of 14 %. Other research has
shown that domestic Leca has lower levels of resistance
compared to its foreign counterparts.11 Hence, Leca was
used as a fine aggregate and substitute for part of the
natural sand. The natural river sand, 1.46 % water
absorption, was the other part of the fine aggregate. It
was washed before use in the mixture procedures. The
normal coarse aggregate was crushed stone with a
maximum size of 20 mm and 0.8 % water absorption,
except for the lightweight concretes, where the
maximum size was 10 mm with 0.5 % water absorption.
Limestone powder was used to maintain the viscosity of
the fresh mix and hence to reduce the segregation
phenomena. The chemical characteristics of the cement
and mineral admixture are presented in Table 1.
Superplasticizer Viscocrete 1 was introduced into the
mixes. The properties of the concrete mix used in this
study and the properties of the fresh concrete are shown
in Tables 2 and 3, respectively. SSCC and SLWSCC
(Light weight self-compacting concrete containing silica
fume) produced mixtures containing 7.0% silica fume by
weight of cement content. NSCC and NLWSCC (Light
weight self-compacting concrete containing nano-sized
SiO2) had mixtures containing nano-sized SiO2 with a
particle size of 15 nm in the amount of 3.0 % by weight
of cement content.
Full scale walls were cast in steel molds. To simulate
the site construction, workshop executives and technical
M. HOSSEINALI BEYGI et al.: EVALUATION OF THE STRENGTH VARIATION OF NORMAL ...
572 Materiali in tehnologije / Materials and technology 45 (2011) 6, 571–577
Figure 1: Details of reinforcement for the walls
Slika 1: Detajli utrditve stene
Table 1: Chemical properties of cement and mineral admixtures
Tabela 1: Kemi~na sestava cementnih in mineralnih zmesi
mass fractions, w/%
Cement Silica fume Limestonefiller Nano SiO2
SiO2 22.02 96.4 – 99.9
Al2O3 4.40 1.32 – –
Fe2O3 3.20 0.87 – –
CaO 64.9 0.49 – –
MgO 1.42 0.97 – –
SO3 1.67 0.10 – –
Na2O 0.27 0.31 – –
LOI 1.30 – – 0.1
P2O – 0.16 – –
MgCO3 – – 1.10 –
CaCO3 – – 98.1 –
staff were involved. The concrete was supplied in four
0.3 m3 batches. During the concreting of the reinforced
concrete walls, the different concretes were poured
directly from the top of the walls without any external
vibration into the framework except for the wall made of
normal concrete, which was compacted with a common
vibrator. Specimens were cured under wet textiles and
polythene sheets for 7 d. Five 100 mm cubes were also
cast from each of the four batches. Half of the cubic
specimens were subjected to curing conditions similar to
those of the walls, and other half were treated with the
standard curing regime.
3 TEST PROCEDURES
3.1 Ultrasonic Pulse Velocity (UPV) Test
Non-destructive tests of concrete are preferred
because measurements can be obtained without
destructive forces. The most generally used method is
the ultrasonic pulse velocity method. The test procedure
is based on the fact that the velocity of an ultrasonic
pulse wave in solid bodies depends on the modulus of
elasticity, Poisson’s ratio and the density of the material.
When the uniformity, density and homogeneity of the
concrete are good, ultrasonic pulse waves of higher
velocity can be observed. Generally, acoustic transducers
are used to generate ultrasonic pulses.12 When the pulse
wave passes through concrete, it undergoes various
reflections at the boundaries of materials with different
properties within the concrete body. The velocity of the
pulse waves is often independent of the geometrical
shapes of the materials through which they pass. The
pulse velocity test is a common method to assess
structural concrete. The actual pulse velocity wave
obtained depends mainly on the materials and mixture of
concrete.
After traveling a known path length (L) in the
concrete, the pulse is converted into an electrical signal
by a second electro-acoustical transducer, and an
electronic timing circuit enables the transfer time (T) of
the pulse to be recorded. The pulse wave velocity (V) is
obtained by
V
L
t
= (1)
where V/(km/s) pulse wave velocity, /cm length of path,
and /μs transfer time
The longitudinal ultrasonic pulse waves that leave the
transmitter travel in the direction normal to the
transducer surface according to BS 1881 13. To make a
reasonably accurate and relevant assessment of the
uniformity of the concrete strength in existing walls in
this study, the ultrasonic pulse velocity test was used as a
non-destructive test method. Ultrasonic pulse velocity
(UPV) tests on the walls at 21 d were performed using
Pundit equipment 14. The pulse velocity measurements
were taken directly through the thickness of the walls at
the grid locations indicated in Figure 2. The UPV test
locations included 21 points, which were located at the
top, middle and bottom levels at heights of (175, 100 and
25) cm, respectively, above the bottom surface of the
walls. These locations were chosen to test different levels
within the walls while satisfying the minimum edge
distance and spacing requirements and avoiding
reinforcing steel.
M. HOSSEINALI BEYGI et al.: EVALUATION OF THE STRENGTH VARIATION OF NORMAL ...
Materiali in tehnologije / Materials and technology 45 (2011) 4, 571–577 573
Table 2: Mixture proportions
Tabela 2: Sestava zmesi
Coarse
Agg
Fine
Agg.
(kg/m3)
Leca
0–3 mm
(kg/m3)
Limestone
powder
(kg/m3)
Water
(kg/m3)
Superpla-
sticizer
(kg/m3)
Nano
SiO2
(kg/m3)
Silica
fume
(kg/m3)
Cement
(kg/m3)
Mixture Type
10–20 mm
(kg/m3)
5–10 mm
(kg/m3)
648 349 815 – – 200 – – – 320 NC
384 469 873 – 270 181 8 – 22.4 300 SSCC
384 469 873 – 270 166 8.5 9.6 – 310 NSCC
– 250 382 245 290 198 9.8 – 30.8 440 SLW-SCC
– 250 382 245 290 205 10.2 14.1 – 456 NLW-SCC
Table 3: Fresh properties
Tabela 3: Lastnosti sve`ih vzorcev
Mix Type
Slump flow V/s
Funnel
L-Box
(h2 / h1)
Unit
weight
(kg/m3)T50/s
D/cm
Final
NC – – – – 2395
SSCC 2.2 68 7.2 0.93 2373
NSCC 2.5 65 7.8 0.94 2362
SLW-SCC 3.8 69 8.1 0.91 1840
NLW-SCC 4.2 63 7.9 0.90 1831
Figure 2: Test positions on the walls
Slika 2: Mesta preizkusov v stenah
4 RESULTS AND DISCUSSION
4.1 Compressive strength
To obtain the compressive strength and UPV, the
walls were tested at the ages of (3, 7, 14 and 28) d. The
compressive strengths of five 100 mm cube specimens
were tested at the ages of (3, 7, 14 and 28) d. All results
were the average of five 100 mm cubes. The UPV
measurements were repeated three times for each cube
specimen. The values found for the compressive strength
and UPV tests of the mixtures at different ages are
tabulated in Table 4.
4.2 The relationship between the ultrasonic pulse ve-
locity and the compressive strength
From the data listed in Table 4, which shows the
averages of the compressive strength and UPV test
results, a regression analysis for the pulse velocity versus
compressive strength of the cube samples for each
mixture was performed. The correlations were
established using an exponential curve model. The pulse
velocity functions and the correlation coefficient (R2) are
presented in Table 5. The high correlation coefficient of
the numerical formula indicates the suitability of the
functions.
Table 5: Relationships between compressive strength and in-place test
results
Tabela 5: Odvisnost med tla~no trdnostjo in rezultati in situ preiz-
kusov
Mixture Type
Regression equations
Pulse velocity function Correlationcoefficient (R2)
NC f'c = 0.0204 e1.8142V 0.9217
SSCC f'c = 1.906 e0.6679V 0.9412
NSCC f'c = 0.0801 e1.3998V 0.8914
SLWSCC f'c = 0.1857 e1.1872V 0.9873
NSLWSCC f'c = 0.7731 e0.845V 0.9646
4.3 Distribution of the concrete strength within con-
crete walls
The results of ultrasonic pulse velocity measurements
at different test points at the age of 28 d are presented in
Figure 3. The concrete compressive strength of the rein-
forced concrete walls was determined by measuring the
ultrasonic pulse velocity along the wall thickness and
converting it to the cube compressive strength using the
relevant function. The average of the ultrasonic pulse ve-
locity measurements and the estimated compressive
strength using the pulse velocity functions are summa-
rized in Table 6.
The strength distribution of the concrete at the levels
of (25, 100 and 175) cm from the bottom of the walls
was studied. For all mixes, the strength variation in the
walls showed that the bottom regions of the walls were
stronger than the top regions. Some trends were also pre-
viously observed in structural members of normal and
lightweight concrete and self-compacting concrete.5–11
The phenomenon of strength reduction along the height
of the members was probably due to several factors.
Toosi et al. studied the strength variation of normal con-
crete and showed that local segregation and bleeding oc-
cur under aggregates, which leads to microcracks and
voids beneath the aggregate surface. Therefore, the
paste-aggregate bond is weakened. They also showed
bond improvement because of the increase of the pres-
sure head in the lower layers.8 In the case of lightweight
concrete, a nonuniform density distribution also occurs
because of the porosity and floating of the lightweight
aggregates in fresh concrete. Consequently, lightweight
aggregates tend to move to the upper levels of concrete
M. HOSSEINALI BEYGI et al.: EVALUATION OF THE STRENGTH VARIATION OF NORMAL ...
574 Materiali in tehnologije / Materials and technology 45 (2011) 6, 571–577
Table 4: Hardened properties
Tabela 4: Lastnosti strjenih vzorcev
Mix Type
Compressive Strength (MPa) Ultrasonic Pulse Velocity (km/s)
3 days 7 days 14 days 28 days 3 days 7 days 14 days 28 days
NC 14.2 19.4 23.3 29.0 3.61 3.78 3.89 3.98
SSCC 26.9 29.1 29.8 33.0 3.96 4.08 4.11 4.27
NSCC 30.3 32.3 33.8 35.5 4.26 4.27 4.31 4.36
SLWSCC 22.5 23.4 24.0 25.3 4.04 4.08 4.09 4.14
NLWSCC 24.0 26.8 27.2 27.0 4.07 4.18 4.21 4.22
Figure 3: Ultrasonic pulse velocity measurements in different test
points of the reinforced concrete walls: (a) SSCC, (b) SLWSCC, (c)
NC, (d) NSCC and (e) NLWSCC
Slika 3: Meritve hitrost ultrazvoka v razli~nih to~kah oja~enih beton-
skih sten: (a) SSCC, (b) SLWSCC, (c) NC, (d) NSCC, in (e)
NLWSCC
members. This phenomenon is more intense if coarse
lightweight aggregate is used.15 Previous research has
shown that the tendency of segregation decreases with
increasing the mineral admixtures contents.16 The rela-
tive concrete strength variation at different levels with re-
spect to the concrete strength at the bottom level of the
mixtures is shown in Figure 4.
Figure 4 shows that the strength of middle and upper
levels in comparison to the lower level decreased by 8 %
and 14 % for NLWSCC, 10 % and 8 % for NSCC, 11.4
% and 16.5 % for SLWSCC, 11 % and 14.5 % for SSCC
and 21 % and 28 % for normal concrete. The results in-
dicated a similar trend for the normal weight self-com-
pacting concrete mixes (SSCC and NSCC) and light-
weight self-compacting concrete (SLWSCC and
NLWSCC), and there was a significant discrepancy be-
tween the self-compacted mixtures and the normal con-
crete mix. It is possible that the results could be different
due to the use of more Leca, especially coarse aggre-
gates. As predicted, because of the use of mineral admix-
tures and a fine lightweight aggregate in this study, a
lower tendency of segregation was observed in the
SLWSCC and NLWSCC mixtures. In other words, light-
weight self-compacting concrete had greater homogene-
ity than the normal concrete, but compared with normal
self-compacting concrete, it showed less homogeneity.
Because self-compacted mixtures contain a lower coarse
aggregate volume of smaller sizes compared to NC, the
size and quantity of microcracks are much lower in
self-compacted mixtures, which leads to paste-aggregate
bond improvement. As mentioned above, the effects of
head pressure and local segregation on the strength varia-
tion are much lower in self-compacted mixes than in NC.
Some previous reports indicated considerable statistical
discrepancies in the homogeneity of properties.10,11 How-
ever, in this study significant differences in strength vari-
ation between self-compacted concrete and NC were ob-
served. Khayat showed a similar trend of changes in the
M. HOSSEINALI BEYGI et al.: EVALUATION OF THE STRENGTH VARIATION OF NORMAL ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 571–577 575
Figure 5: Stress contour plots based on 28-d UPV measurements: (a)
NC, (b) NLWSCC, (c) NSCC, (d) SSCC, and (e) SLWSCC
Slika 5: Porazdelitev napetosti na podlagi 28-dnevnih meritev UPV:
(a) NC, (b) NLWSCC, (c) NSCC, (d) SSCC in (e) SLWSCC
Figure 4: In situ strength variations across the height of the walls
Slika 4: Variacije in situ trdnosti po vi{ini sten
Table 6: Estimated in situ strengths and test results (28 days)
Tabela 6: Ocenjene in situ trdnosti in rezultati preizkusov (28 d)
Level of
walls
NC SSCC NSCC SLWSCC NLWSCC
f'c (MPa)
UPV mea-
surements
(km/s)
f'c (MPa)
UPV mea-
surements
(km/s)
f'c (MPa)
UPV mea-
surements
(km/s)
f'c (MPa)
UPV mea-
surements
(km/s)
f'c (MPa)
UPV mea-
surements
(km/s)
Top 16.7 3.70 26.8 3.95 31.5 4.27 20.7 3.97 22.3 3.98
Middle 18.3 3.74 27.9 4.01 30.7 4.25 22.0 4.02 23.7 4.05
Bottom 23.2 3.87 31.2 4.19 33.4 4.31 24.9 4.12 25.8 4.15
Average 19.4 3.77 28.6 4.05 31.8 4.27 22.5 4.03 23.9 4.06
Note: f'c = Cube compressive strength.
concrete strength of normal self-compacting and normal
concrete walls.11 However, the walls that Khayat consid-
ered were not full scale or reinforced. Hence, the effects
of the reinforcement configurations on compaction were
not considered, and more studies are necessary. The in
situ strength variability is shown in the contour plot of
Figure 5. This contour plot provides further comparison
of the strength variation within the walls both across
their height and their length. Random variation along the
member length was noted. Figure 6 shows the strength
changes with specific trends along the wall height. How-
ever, this trend varied for different mixtures. The NC
wall exhibited more intensive discrepancies because of
the effect of the internal vibration on the compaction of
the concrete.
In Table 7, the wet-cured standard cubic compressive
strengths are compared with equal cubic strengths at dif-
ferent levels of the walls. Figure 6 shows that the in situ
strength varied from 81 % to 102 % of the wet-cured
standard 28 d strength for all four self-compacted con-
crete walls. However, this difference was about 62 % to
86 % for the normal concrete wall. The main reason for
the differences between the in situ and standard 28 d
strengths is the low sensitivity of self-compacted mix-
tures to curing conditions. This observation was previ-
ously reported by Pera et al.17 On the other hand, the
high discrepancy between the lightweight self-compact-
ing concretes and NC might be because the high initial
moisture content of the lightweight aggregates led to typ-
ical internal curing. Thus, low sensitivity to the curing
regime can be expected for lightweight self-compacting
concretes.
A previous study showed 75 % to 90 % differences
between the standard cube and in situ 28 d strengths for
the SCC mixture.11
4.4 Distribution of concrete strength along the walls
Figure 7 indicates the coefficient of variation (COV)
of strength for different heights along the experimental
walls. The COV of strength measurements ranged from 1
% to 4.1 % for self-compacted mixes and were limited to
10.2 % for the NC mix. Therefore, the self-compacted
mixtures had better uniformity along the length of the
walls.
5 CONCLUSIONS
The following conclusions were drawn from the re-
sults presented in this paper concerning the variation of
M. HOSSEINALI BEYGI et al.: EVALUATION OF THE STRENGTH VARIATION OF NORMAL ...
576 Materiali in tehnologije / Materials and technology 45 (2011) 6, 571–577
*: Average value of compressive strength
Figure 7: Strength variations along the length of the walls with re-
spect to the average value of each level
Slika 7: Variacije trdnosti po dol`ini sten v razmerju s povpre~no
vrednostjo na vsakem nivoju
Table 7: Comparison of cube strengths and estimated in situ strengths
Tabela 7: Primerjava trdnosti kock in ocenjenih in situ trdnosti
Mixture
type Level
f: Estimated in
situ cube
strength (MPa)
f'c: 28 day stan-
dard cube strength
in wet-cured con-
dition (MPa)
f
fc
'
NC
Top 16.7
27.0
0.62
Mid. 18.3 0.64
Bot. 23.2 0.86
SSCC
Top 26.8
32.8
0.82
Mid. 27.9 0.85
Bot. 31.2 0.95
NSCC
Top 31.5
32.7
0.96
Mid. 30.7 0.94
Bot. 33.4 1.02
SLWSCC
Top 20.7
25.5
0.81
Mid. 22 0.86
Bot. 24.9 0.97
NLWSCC
Top 22.3
28.0
0.97
Mid. 23.7 0.84
Bot. 25.8 0.92
Figure 6: In situ compressive strength in walls in relation to standard
28 d strength
Slika 6: In situ tla~na trdnost v stenah v odvisnosti od standardne
28-dnevne trdnosti
M. HOSSEINALI BEYGI et al.: EVALUATION OF THE STRENGTH VARIATION OF NORMAL ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 571–577 577
factors such as strength level, type of lightweight aggre-
gate and admixture content.
1. Strength variations across the height of the reinforced
concrete walls followed the general pattern of a rea-
sonably uniform distribution from top to bottom, with
the top region having a lower strength than the bot-
tom region. However, the magnitude of this variation
varied according to the concrete type.
2. The low in situ strength variations of all self-com-
pacted mixtures showed greater strength uniformity
than that of the NC mixture along the height of the
walls.
3. Due to the low sensitivity of self-compacted mixes to
curing conditions, smaller differences were observed
in the self-compacted mixes than the normal concrete
mix.
4. The COV of the strength measurements showed
better uniformity for all normal self-compacting and
lightweight self-compacting mixtures.
5. By replacing the silica fume with nano-sized SiO2 as
the admixture, no considerable changes were ob-
served in strength variation trends.
6 REFERENCES
1 Skarendahl, A., Petersson, O. Self compacting concrete: state-of-
the-art report of RILEM TC 174-SCC, Report 23, RILEM Pub-
lishers, Cachan, France, 2000
2 Koyata, H., Comman, C. R. 2005. Workability measurement and de-
veloping robust SCC mixture designs. In proceeding if Second north
American conference on the design and use of self consolidating
concrete (SCC) and the fourth International RILEM Symposium on
self compacting concrete. Addison, IL, USA: Hanley Wood Pub-
lishers, (2005), 799–805
3 N'ielsson, I., Wallevik, O.H. In: Wallevik O, Nielsson I., 2003.
Rheologycal evaluation of some empirical test methods – prelimi-
nary results proceedings of the 3rd international RILEM symposium
Reykjavik, Iceland. RILEM Publishers PRO, 33 (2003), 59–68
4 Videla C., Lopez, M. Effect of lightweight aggregate intrinsic
strength on lightweight concrete compressive strength and modulus
of elasticity. Mater de construction, 52 (2002), 23–27
5 Madandoust, R. Strength assessment of lightweight concrete. PhD
thesis, The University of Liverpool; 1990
6 Bungey J. H., Madandoust R. Strength variation in lightweight con-
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7 Toosi, M., Houde, J. Evaluation of strength variation due to height of
concrete members. Cem. Concr. Res., 11 (1981), 519–529
8 Toosi, M. Variation of concrete strength due to pressure exerted on
fresh concrete. Cem. Concr. Res., 10 (1980), 845–852
9 Kumar, R., Bhattacharjee, B. Porosity, pore size distribution and in
situ strength of concrete. Cem. Concr. Res., 33 (2002), 155–164
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of self compacting concrete in full- scale structural element, Cem.
Concr. Compos., 23 (2001), 57–64
11 Khayat, K. H, Manai, K., Trudel, A. In situ mechanical properties of
walls elements cast using self consolidating concrete, ACI Mater J.,
941 (1997), 491–500
12 Malhotra, V. M., Carino, N. J. CRC handbook on non-destructive
testing of concrete. CRC Press, 1997
13 Bungey, J. The testing of concrete in structures. London, Uk: Surrey
university Press, 1989
14 PUNDIT Ultrasonic concrete tester. C. N. S. Electronics LTD, 61–63
Holms road, London, NW5, England
15 Mindess, S., Young, J.F, Darwin, D. Concrete 2nd Prentice Hall, 2003
16 Demirboða, R., Örüng, I, Gül, R. Effect of expanded perlite aggre-
gate and mineral admixtures on compressive strength of low density
concrete. Cem. Concr. Res., 31 (2001), 1627–32
17 Pera, J., Husson, S., Guilhot, B. Influence of finely ground limestone
on cement hydration. Cem. Concr. Res. 21 (1999), 99–105

O. POPOVI] et al.: THE INFLUENCE OF BUFFER LAYER ON THE PROPERTIES OF SURFACE ...
THE INFLUENCE OF BUFFER LAYER ON THE
PROPERTIES OF SURFACE WELDED JOINT OF
HIGH-CARBON STEEL
VPLIV VMESNE PLASTI NA LASTNOSTI POVR[INSKIH ZVAROV
JEKLA Z VELIKO OGLJIKA
Olivera Popovi}1, Radica Proki} - Cvetkovi}1, Aleksandar Sedmak1, Galip Buyukyildrim2,
Aleksandar Bukvi}3
1Faculty of Mechanical Engineering, University of Belgrade, 11000 Belgrade, Serbia
2EWE, Istanbul, Turkey
3Ministry of defense, 11000 Belgrade, Serbia
opopovic@mas.bg.ac.rs
Prejem rokopisa – received: 2011-02-02; sprejem za objavo – accepted for publication: 2011-06-02
Surface welding with buffer layer is often in use because of its well-known properties of plasticity, or ability to slow crack
growth initiated. However, in modern surface welding technologies, buffer layer is rarely used. New classes of flux-cored and
self-shielded wires are recently developed and it is possible to achieve the requested properties of welded joints without buffer
layer. In this paper, for comparison, the high-carbon steel surface was welded with and without buffer layer. In both cases, it has
been used same surface process, but with different filler materials and equal heat input. The mechanical properties, total impact
energy, as its components, the fatigue threshold value of Kth, and the crack growth rate da/dN were determined. The results
obtained at room temperature show better properties of the sample surface welded with the buffer layer, but, with temperature
decrease a sharp decrease of toughness of the sample welded with buffer layer occured. Also, buffer layer didn’t change the
property of initiated crack in terms of crack growth rate. The construction from high-carbon steel are exposed to low
exploitation temperature and are used for prolong working time, thus the use of buffer layer in modern surface welding
technologies is not recommended.
Keywords: surface welding, buffer layer, welded joint, toughness, crack growth parameters
Povr{insko varjenje z vmesno (buffer) plastjo se ve~krat uporablja zaradi dobre plasti~nosti in sposobnosti za prepre~evanje rasti
nastale razpoke. Vendar se redkeje uporablja pri sodobnih tehnologijah povr{inskega varjenja. Nove vrste polnjene in
samoza{~itne `ice so bile razvite in je bilo tako mogo~e dose~i zahtevane lastnosti zvarov brez vmesne plasti. V tem delu je
opisana zavarjena povr{ina jekla z veliko ogljika z vmesno plastjo in brez nje. V obeh primerih je bil uporabljen enak proces z
enakim vnosom toplote, vendar z razli~nim varilnim materialom. Dolo~ene so bile mehanske lastnosti, skupna `ilavost in njene
komponente, prag utrujenosti Kth in hitrost napredovanja razpoke da/dN. Lastnosti pri sobni temperaturi so bolj{e pri vzorcu
povr{ine, ki je bil zavarjen z vmesno plastjo, vendar se je pri zni`anju temperature hitro zmanj{ala `ilavost vzorca, ki je bil
zvarjen z vmesno plastjo. Tudi vmesna plast ni spremenila hitrosti rasti za~ete razpoke. Konstrukcije iz jekla z veliko ogljika
obratujejo pri nizki temperaturi in se uporabljajo dolgo ~asa. Zato ni priporo~ena uporaba vmesne plasti pri modernih
tehnologijah varjenja povr{ine.
Klju~ne besede: varjenje povr{ine, vmesna plast, zvar, `ilavost, parametri rasti razpoke
1 INTRODUCTION
The main properties of high-carbon steels are high
hardness and strength and having a pearlitic microstruc-
ture, have a typically low toughness and crack growth re-
sistance also. Since in exploitation they are often ex-
posed to wear and rolling contact fatigue, parts become
unfit for service due to unacceptable profiles, cracking,
spalling etc. Surface welding is maintenance way to pro-
long the exploitation life of damaged parts.1 For surface
welding are mostly in use semi-automatic arc welding
processes, with flux-cored and self-shielded wires. Basic
difference between them is that the first requires an ex-
ternal shielding gas. In both cases, core material acts as a
deoxidizer, helping to purify the weld metal, generate
slag formers and by adding alloying elements to the core,
it is possible to increase the strength and provide other
desirable weld metal properties.2,3 These processes have
replaced slowly MMA process and they almost ideal for
outdoors in heavy winds. The result of flux-cored wire
application are higher quality welds, faster welding and
maximizing a certain area of welding performance.4 The
number of layers in surface welded joint depends of the
damage degree, most frequently it’s consists of three lay-
ers, sometimes with buffer layer, also. The buffer layer is
applied for the crack sensitive materials, what high car-
bon steel certainly is (high CE). The function of buffer
layer is to slow down the growth of initiated crack with
its own plasticity. Constructions, like railways, are ex-
posed to cyclic load and wear in exploatation, that the
crack initiate. Sometimes it is necessary to use a buffer
layer, which besides good affects, may have drawbacks,
also. Namely, the use of buffer layer slows down signifi-
cantly the surface welding process, due to replacement of
wires and settings of other welding parameters. Since, as
already noted, for surface welding are used mainly
semi-automatic and automatic processes, it significantly
Materiali in tehnologije / Materials and technology 45 (2011) 6, 579–584 579
UDK 621.791.05:669.1 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)579(2011)
extends the working time. New classes of flux-cored and
self-shielded wires are developed recently, and it is pos-
sible to achieve the requested properties of welded joints
without buffer layer.
2 EXPERIMENTAL PROCEDURE
The investigation was carried out with high carbon
steel with 0.52C-0.39Si-1.06Mn-0.042P-0.038S-
0.011Cu-0.006Al, having initial pearlitic microstructure
and tensile strength of 680–830 N/ mm2.
The surface welding of the testing plates was
perfomed with a semi-automatic process. As the filler
material, the self-shielded wire (FCAW-S) and flux-
cored wires (FCAW) with chemical compositions and
mechanical properties given in Table 1, were used. The
plates were surface welded in three layers; sample 1 with
FCAW-S without buffer layer; sample 2 with FCAW
with buffer layer, as shown in Table 1.
Since the CE-equivalent was CE = 0.644, the heat in-
put during welding was of 10 kJ/cm, the preheating tem-
perature was of 230 °C, and the controlled interpass tem-
perature was of 250 °C. Sample 1 was surfaced with one
type of filler material (self-shielded wire), while for sur-
facing of sample 2 two types of wires were used, but
both flux-cored: one for the buffer layer and the second
for the last two layers. As shielded gas for welding of
sample 2, CO2 was used. To evaluate the mechanical
properties, specimens for further investigation were cut
from surface welded joints.
3 HARDNESS
Hardness measurements were performed using the
100 Pa load. The hardness profiles of surface welded
joints are shown in Figure 1. The lowest hardness is for
the base metal (250–300 HV), being the hardness of nat-
urally cooled standard rails.5,6 In HAZ hardness increase
is noticable in both samples, due to complex heat treat-
ment and grain refinement.4 In sample 2 in the first sur-
faced layer, i.e. in buffer layer the hardness is decreased
sharply. The function of buffer layer is to stop the growth
crack initiates with own plasticity and lower hardness.
The hardness of II and III welded layers of both samples
are the highest and similar, due to influence of alloying
elements in filler materials, which shift transformation
points to bainitic region.4 The maximum hardness level
of 350–390 HV is reached in surface welded layers and
it provides improvement of mechanical properties and
wear resistance.4
4 TENSILE TESTS
The tensile tests were performed on a 2 mm thick
specimens. The room temperature mechanical properties
(ultimate tensile strength, UTS) of the surface welding
layers are shown in Figure 2. The basic requirement in
O. POPOVI] et al.: THE INFLUENCE OF BUFFER LAYER ON THE PROPERTIES OF SURFACE ...
580 Materiali in tehnologije / Materials and technology 45 (2011) 6, 579–584
Table 1: Chemical composition of filler materials
Tabela 1: Kemi~na sestava varilnih materialov
Sample
No. Wire designation
Wire
diam.
d/mm
Chemical composition, mass fractions, w/% Hardness,
HRCC Si Mn Cr Mo Ni Al
Sample 1 OK Tubrodur 15.43(self-shielded wire) 1.6 0.15 <0.5 1.1 1.0 0.5 2.3 1.6 30–40
Sample 2
1.layer
(buffer layer)
Filtub 12B
(flux-cored wire) 1.2 0.05 0.35 1.4 - - - - -
2. and 3.
layer
Filtub dur 12
(flux-cored wire) 1.6 0.12 0.6 1.5 5.5 1.0 - - 37–42
Figure 1: Hardness profiles along the joint cross-section of samples
Slika 1: Profili trdote na pre~nem prerezu vzorcev
Figure 2: Ultimate tensile strength of the surface welded joints
Slika 2: Raztr`na trdnost povr{inskih zvarov
welded structures design is to assure the required
strength. In most welded structures this is obtained with
superior strength of WM compared to BM (overmatch-
ing effect), and in tested case this is achieved7,8. The
highest UTS was found for the weld metal of sample 2
(1210 MPa), due to solid state strengthening by alloying
elements.9
5 IMPACT TESTING
The impact testing was performed according to EN
10045-1, i.e ASTM E23-95, with Charpy V notched spe-
cimens on the instrumented machine SCHENCK TRE-
BEL 150 J. Impact testing results are given in Table 2, 3
and in Figure 3 for base metal and HAZ at all testing
temperatures. The total impact energy, as well as crack
initiation and crack propagation energies, for weld metal
of both samples at all testing temperatures are presented
in Table 4 and in Figure 4.
The total energy of base metal is very low (5 J), due
to very hard and very brittle cementite lamellae in
pearlite microstructure,4 while the toughness of HAZ is
higher (11–12 J) and is similar for both samples at all
testing temperatures.
Table 2: Instrumented impact testing results of Charpy V specimens
for base metal and HAZ at all testing temperatures
Tabela 2: Rezultati instrumentiranih Charpyjevih preizkusov za
osnovni material in HAZ pri vseh temperaturah preizku{anja
Total impact energy, Eu/ J
20 °C –20 °C –40 °C
base metal 5 3 3
sample 1-HAZ 1 12 11 10
sample 2-HAZ 2 11 10 9
Table 3: Instrumented impact testing results of Charpy V surface weld
metal specimens at all testing temperatures
Tabela 3: Rezultati instrumentiranih Charpyjevih preizkusov za
vzorce V povr{inskih zvarov pri vseh temperaturah preizku{anja
sample 1-WM1 sample 2-WM2
20 °C –20 °C –40 °C 20 °C –20 °C –40 °C
Total impact en-
ergy, Eu/ J
29 23 17 34 14 11
Crack initiation
energy, Ein/ J
20 16 15 12 10 10
Crack propaga-
tion energy, Epr/ J
9 7 2 22 4 1
The total impact energy of samples 1 and 2 at room
temperature are significantly higher (29 J and 34 J) than
in base metal (5 J), as consequence of appropriate choice
of alloying elements in the filler material. The presence
of Ni, Mn and Mo promotes the formation of needled
bainitic microstructure and grain refinements, and in-
creases the strength and toughness also9. By analyzing
the impact energy values of sample 1, a change of tough-
ness in continuity is observed, with no marked drop of
toughness, and for all tested temperatures, crack initia-
tion energy is higher than crack propagation energy. This
O. POPOVI] et al.: THE INFLUENCE OF BUFFER LAYER ON THE PROPERTIES OF SURFACE ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 579–584 581
Figure 3: Dependence total impact energy, crack initiation and crack
propagation energy vs.temperature for a) weld metal of sample 1 and
b)weld metal of sample 2
Slika 3: Odvisnost skupne `ilavosti ter energije za~etka in napre-
dovanja razpoke od temperature za zvar vzorca 1 in zvar vzorca 2
Figure 4: Diagrams force-time, obtained by instrumented Charpy pen-
dulum for sample 1 and sample 2
Slika 4: Odvisnost sile od ~asa, dolo~ena z instrumentalnim Charpy-
jevim kladivom za vzorca 1 in 2
is the reason for the absence of significant decrease of
toughness. The highest value of total impact energy was
found for the sample 2 at room temperaure (34 J), which
is the only case when the initiation energy is lower than
propagation energy (12 J and 22 J, respectivelly). This
shown practically the buffer layer function. Namely, the
initiated crack during propagation comes to plastic buffer
layer, which slows down crack further growth. For this
reason, the crack propagation energy is the largest part of
total impact energy. However, at –20 °C, significant drop
of total impact energy is noticable (14 J) due to losing of
buffer layer plastic properties at lower temperatures. The
low-carbon wire (0.05 % C and 1.4 % Mn) has excellent
toughness, but and marked rapid drop on S-curve (de-
pendence toughness vs. temperature). Transition temper-
ature of this material above –20 °C is confirmed by the
obtained impact toughness results. The use of buffer
layer is reasonable if the exploatation temperature is
above –5 °C; on the contrary, at lower temperatures,
buffer layer is losing its function and the toughess is de-
creased.
Diagrams force-time, obtained by instrumented
Charpy pendulum, are given in Figure 5. As can be seen,
for the sample 1 the character of diagrams force-time
changed little by lower temperature. Namely, this mate-
rial at room temperature has diagram with marked rapid
drop, as consequence of unstable crack growth. After the
maximum load, a very fast crack growth is started, and it
is confirmed by the low value of crack propagation en-
ergy.10 On the contrary, on the sample 2 diagram at room
temperature, the presence of buffer layer is clearly
shown. The initiated crack, during its growth, comes to
buffer layer which temporary stops the further crack
growth and changes crack growth rate. The obtained ex-
perimental diagram doesn’t belong to any type, accord-
ing to standard EN 10045-1. This leads to toughness in-
crease, primarily crack propagation energy, and it is also
here the only case when the crack initiation energy is
lower than crack propagation energy.
6 CRACK GROWTH RATE
A basic contribution of fracture mechanics in fatigue
analysis is the division of fracture process to crack initia-
tion period and the growth period to critical size for fast
fracture7. Fatigue crack growth tests had been performed
on the CRACKTRONIC dynamic testing device in
FRACTOMAT system, with standard Charpy size speci-
mens, at room temperature, and the ratio R = 0.1. A stan-
dard 2 mm V notch was located in third layer of WM, for
the estimation of parameters for WM and HAZ, since
initiated crack will propagate through those zones. Crack
was initiated from surface (WM) and propagated into
HAZ, enabling calculation of crack growth rate da/dN
and fatigue treshold Kth.4 The results of crack growth
resistance parameters, i.e., obtained relationship da/dN
vs. K for sample 1 and for sample 2 are given in Figure
6. Parameters C and m in Paris law, fatigue threshold
Kth and crack growth rate values are given in Table 5
for both samples as obtained from relationships given in
Figure 6, for corresponding K values.
The behaviour of welded joint and its constituents
should affect the change of curve slope in the part of va-
lidity of the Paris law. Materials of lower fatigue-crack
growth rate have lower slope in the diagram da/dN vs.
K.7 For comparison of the properties of surface welded
joint constituents the crack growth rates are calculated
for different values of stress-intensity factor range K.
Bearing in mind that the weld metal consists of two lay-
O. POPOVI] et al.: THE INFLUENCE OF BUFFER LAYER ON THE PROPERTIES OF SURFACE ...
582 Materiali in tehnologije / Materials and technology 45 (2011) 6, 579–584
Figure 6: Diagram da/dN vs. K for sample 1 and sample 2
Slika 6: Diagram da/dN za vzorca 1 in 2
Table 4: Parameters C, m, Kth and crack growth rate values for all zones of surface welded joints
Tabela 4: Parametri C, m, Kth in hitrost rasti razpoke za vse dele povr{insko zvarjenih vzorcev
Zone of surface
welded joint
Fatigue thresh-
old Kth/
(MPa m1/2)
Parameter
C
Parameter
m
Crack growth rate (da/dN)/m
K = 15
MPa m1/2
K = 20
MPa m1/2
K = 30
MPa m1/2
sample 1
WM 1
9,5
4.45 
 10–13 3.74 1.11 
 10–8 - -
WM 2 3.78 
 10–13 3.61 - 1.88 
 10–8 -
HAZ 4.07 
 10–13 3.79 - - 1,61 
 10–7
sample 2
WM 1
8,9
4.63 
 10–13 3.87 1.65 
 10–8 - -
WM 2 3.85 
 10–13 3.88 - 2.07 
 10–7 -
HAZ 3.76 
 10–13 3.93 - - 1.18 
 10–6
ers (third layer is used for V notch), as referent values of
K were taken: K = 15 MPa m1/2 for WM1, K = 20
MPa m1/2 for WM2, and K = 30 MPa m1/2 for HAZ. It’s
important that all the selected values are within the mid-
dle part of the diagram, where Paris law is applied. In all
three zones of surface welded joint (WM2, WM1 and
HAZ), the sample 2 with buffer layer has a higher crack
growth rate than sample 1, i.e. the growth of initiated
crack will be slower in sample 1. This means that for the
same value of stress intensity factor K, the specimen of
sample 2 needs less number of cycles of variable ampli-
tude than the specimen of sample 1, for the same crack
increment.9 The maximum fatigue crack growth rate is
achieved in HAZ for both samples, when stress intensity
factor range approaches to plane strain fracture tough-
ness.
If a structural component is continuously exposed to
variable loads, fatigue crack may initiate and propagate
from severe stress raisers if the stress intensity factor
range at fatigue threshold Kth is exceeded.7 The fatigue
treshold value Kth for sample 2 (Kth = 8.9 MPa m1/2) is
lower than that for sample 1 (Kth = 9.5 MPa m1/2). This
means that the crack in sample 2 will be initiated earlier,
i.e. after less number of cycles, than in sample 1.
Values of fatigue threshold and crack growth rates
corespond to initiation and propagation energies in im-
pact testing, and in this case, good corelation is
achieved.9 Sample 1 has higher crack initiation energy
(20 J) and higher Kth (Kth = 9.5 MPa m1/2 for sample 1
and Kth = 8.9 MPa m1/2 for sample 2). With compa-
ration of crack propagation energy and crack growth
rate, it is hard to establish the precise analogy, as tough-
ness was estimated for the surface weld metal, whereas
crack growth rate for each surface welded layer. Gen-
erally, buffer layer didn’t show slow,, the initiated crack
growth, with aspect of crack growth rate, while this ef-
fect is obvious in the case of toughness, i.e. crack propa-
gation energy.
7 CONCLUSIONS
On the base of obtained experimental results and
their analysis, the following is concluded:
1. The experimental investigation of surface welded
joints with different weld procedures has shown, as
expected, significant differences on their perfor-
mance in terms of mechanical properties. But, in both
cases, it was shown, that in spite of poor weldability
of high carbon steel, they can be successfully welded.
2. The maximal hardness level of 350–390 HV is
reached in surface welded layers of both samples,
with equal hardness of base metal (250–300 HV).
The main difference appears in the first deposition
layer, where as expected, in sample 2 the hardness is
significantly lower (buffer layer). The obtained hard-
ness values ensure simultaneously the improvement
of mechanical and wear properties, and in the case of
a rail, represents maximal hardness preventing the
wheel wear.4 Similar results are obtained by tensile
testing. Sample 2 has slightly higher ultimate tensile
strength (1360 MPa) than sample 1 (1210 MPa) due
to solid solution strengthening by alloying elements.
3. The greatest differences are found in impact proper-
ties. The highest value of total impact energy of sam-
ple 2 at room temperaure (34 J) was obtained only in
the case when the initiation energy was lower than
propagation energy (12 J and 22 J, respectivelly).
However, at –20 °C, the drop of total impact energy
is significant (14 J), due to lowering of buffer layer
plastic properties at lower temperatures. The transi-
tion temperature of this material is above –20 °C, and
it was confirmed by obtained impact toughness re-
sults. The use of buffer layer is beneficial for exploat-
ation temperature above –5 °C. On the contrary, at
lower temperatures, buffer layer loses its function and
toughess decreases. On the contrary, for sample 1 the
change of toughness is continous and without marked
drop of toughness (29 J at 20 °C and 23 J at –20 °C).
At all tested temperatures, the crack initiation energy
is higher than crack propagation energy. This may be
the reason for the absence of significant decrease of
toughness and that should be kept in mind during de-
sign and exploitation.
4. Results show that sample 2 has higher crack growth
rate (1.65 · 10–8) than sample 1 (1.11 · 10–8), and
lower fatigue treshold value Kth (8.9 MPa m1/2 for
sample 2 and 9.5 MPa m1/2 for sample 1). This means
that the crack in sample 2 will be initiated earlier, i.e.
after less number of cycles, than in sample 1, and that
a less number of cycles is needed to reach the critical
size.
5. Values of fatigue threshold and crack growth rates
corespond to initiation and propagation energies in
impact testing. In the case of fatigue treshold and
crack initiation energy, good correlation was
achieved. Sample 1 has higher crack initiation energy
(20 J) and higher Kth (9.5 MPa m1/2) than sample 2
(12 J and Kth = 8.9 MPa m1/2). On the contrary,
buffer layer didn’t show decrease of initiated crack
growth rate, as this effect is obvious in the case of
toughness, i.e. crack propagation energy. Since the
constructions from high-carbon steel are used at low
temperature, and bearing in mind the extended work-
ing time, in modern surface welding technologies, the
use of buffer layer is not recommended.
Acknowledgement
The research was performed in the frame of the na-
tional project TR 35024 financed by Ministry of Science
of the Republic of Serbia.
O. POPOVI] et al.: THE INFLUENCE OF BUFFER LAYER ON THE PROPERTIES OF SURFACE ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 579–584 583
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E. [VARA FABJAN et al.: CORROSION STABILITY OF DIFFERENT BRONZES IN SIMULATED URBAN RAIN
CORROSION STABILITY OF DIFFERENT BRONZES IN
SIMULATED URBAN RAIN
KOROZIJSKA STABILNOST RAZLI^NIH BRONOV V UMETNEM
KISLEM DE@JU
Erika [vara Fabjan, Tadeja Kosec, Viljem Kuhar, Andra` Legat
National Building and Civil Engineering Institute, Dimi~eva 12, 1000 Ljubljana, Slovenia
erika.svara@zag.si
Prejem rokopisa – received: 2011-04-16; sprejem za objavo – accepted for publication: 2011-06-06
Copper and high copper alloys tend to passivate in humid air. In clean humid air, cuprite slowly transforms to black tenorite. If
atmosphere contains aggressive species, acid rain might effect the formation of different corrosion products. The patina that
forms upon exposure to urban acid rain also depends of the composition of the base alloy. In the present study three different
alloys were investigated: leaded bronze, usually used for sculptures, unleaded bronze as an alloy without an impact on the
environment, and new type of bronze alloy, silicon bronze.
The electrochemical measurements were performed to investigate the different bronze properties in simulated urban acid rain
that contained carbonates, sulphates and nitrates, acidified to pH 5. Morphological characteristics of the three different bronzes
were studied and SEM/EDX analysis of the corrosion products was performed.
It was found that silicon bronze has higher corrosion resistivity than unleaded bronze, the former having higher corrosion
resistivity than leaded bronze. In addition, time dependant electrochemical impedance spectroscopy measurements showed that
polarization resistances for silicon bronze and unleaded increased with time, whereas it decreased for leaded bronze. The
corrosion layer on silicon bronze is more compact and thinner due to homogeneous microstructure.
Keywords: bronze, metallography, corrosion, simulated urban rain
Baker ter bakrove zlitine se na vla`nem zraku pasivirajo, prihaja do nastanka t. i. patin. Najprej se tvori rde~kast kuprit, ki se
nato po~asi spreminja v ~rn tenorit. ^e pa so v atmosferi agresivne zvrsti, lahko nastali kisli de` vpliva na tvorbo razli~nih
korozijskih produktov. Patina, ki se tvori zaradi izpostavitve kislemu de`ju, je odvisna tudi od sestave zlitine. V {tudiji smo
preiskovali tri razli~ne vrste bronastih zlitin: bron z vsebnostjo svinca, pogosto uporabljen pri ulivanju ve~jih bronastih
spomenikov, bron brez vsebnosti svinca ter novo vrsto brona s silicijem.
Elektrokemijske preiskave smo izvedli v simulirani raztopini kislega de`ja, ki je vsebovala sulfate, karbonate in nitrate. Dolo~ili
smo mikrostrukturne zna~ilnosti posameznih bronov ter morfologijo korozijskih produktov.
Ugotovili smo, da je najbolj korozijsko odporna zlitina bakra s silicijem, sledi recentna zlitina brez svinca, najslab{e korozijske
lastnosti pa ima zlitina s svincem. Nadalje smo z elektrokemijsko impedan~no spektroskopijo pokazali, da se korozijske
lastnosti zlitine s svincem s ~asom slab{ajo, korozijske lastnosti zlitine brez svinca ter bron s silicijem pa se s ~asom
izbolj{ujejo. Korozijski produkti na silicijevi zlitini so kompaktni predvsem zaradi homogene mikrostrukture.
Klju~ne besede: bron, metalografija, korozija, simulirana raztopina kislega de`ja
1 INTRODUCTION
Copper and copper alloys may corrode different ways
depending on the type of atmosphere that the bronze ob-
jects are exposed to. In rural areas where atmosphere is
clean, the bronze surface first turns reddish forming cu-
prite and finally it transforms to black tenorite. In gen-
eral, urban and marine atmosphere are more aggressive,
and versatile corrosion products form on non protected
copper or bronze. Numerous studies have been involved
in the subject 1–28.
The bronze problematics is very wide in its area, it
covers studies of different bronze types 3,7,11,20 and natural
patinas that form on copper or bronze 1,8–10,19,24,25,28. The
main concern in some studies is protection of natural and
synthetic artificial patinas 4,5,14,15,18,19. Much work is de-
voted to different corrosion inhibitors 5,14,16–18,27 next to
different basic electrochemical studies 6,22,23.
The corrosion products that form in particular type of
bronze are strongly dependant on the type of bronze, the
type of the atmosphere, the chemical and physical prepa-
ration of surface, patinations and protections. The two
scientific works were result of EUREKA– 2210 Euro-
care bronzart project, where different cast bronzes for
production of bronze artworks were studied 7,11.
The seven different types of bronze with different
contents of copper, tin, zinc, lead, silicon and nickel
were employed in the study by Gallese et al. 11. It was
shown that the bronze alloy with minimal content of Pb
(Cu–4Sn–5Zn) showed very good characteristics, as well
as the bronze alloy with less zinc and more tin
(Cu–9Sn–1.5 Zn) and alloy with the presence of Ni
(Cu–9Sn–2Zn–3Ni). All these alloys showed improved
corrosion characteristics when compared to reference al-
loy (Cu–5Sn–4Zn–5Pb). This traditional alloy that con-
tains lead is recognized as problematic due to lead toxic-
ity.
In another study, the electrochemical properties of al-
loy Cu–5Sn–5Zn–5Pb (G85) were compared to SI3
(Cu–8Sn–3Si) in simulated acid rain and the corrosion
characteristics were evaluated after exposure in simu-
lated urban–industrial and marine environment. It was
found out that silicon bronze exhibited better corrosion
resistance in more aggressive marine environment 7.
Materiali in tehnologije / Materials and technology 45 (2011) 6, 585–591 585
UDK 669.35'6:620.193 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)585(2011)
It is of big importance to develop and investigate
bronzes that would perform well in different corrosive
environments. The information could provide more
knowledge and appropriate selection of suitable materi-
als for the application in the field of Cultural Heritage.
The selection of proper alloy in different environments
could be improved to minimize visual change of bronze
surface as well as unwanted appearance of corrosion
products. Novelty feature of this study is the evaluation
of the difference in corrosion performance of older and
newer type of alloys.
The main objectives of the present study were to
electrochemically investigate the difference of the three
different bronze types: Cu–Zn–Sn and Cu–Zn–Sn–Pb
and Cu–Sn–Si. Potentiodynamic techniques as well as
electrochemical impedance spectroscopy were used for
the study of electrochemical behaviour and passive lay-
ers that formed in simulated urban rain at open circuit
potential. The morphological examination was per-
formed. The SEM/EDX spectroscopic investigation was
made in order to examine the corrosion products that had
formed during 35 days immersion in simulated urban
rain solution.
2 EXPERIMENTAL
2.1 Material and surface preparation
The bronze samples were cast in a sand mould of di-
mension of (100 × 100 × 5) mm. Three different bronzes
were chosen to be investigated: CuSnZn, CuSnZnPb and
CuSnSi. CuSnZn bronze is currently used for casting of
big bronze sculptures, bronze CuSnZnPb contains some
lead, which is advised to be omitted in newly casted
sculptures, and new type of bronze that contains silicon:
CuSnSi. The composition of the three different bronzes,
determinated by portable X-ray fluorescence analysers
X-MET5100, Oxford Instruments, UK, is given in Table
1.
Samples were sectioned from 5 mm plates in the
form of discs of 15 mm diameter used as working elec-
trodes. Prior to measurement, the specimens were
abraded with 800 and 1000 grid emery paper. Finally, the
samples were ultrasonically cleaned in distilled water
and then well dried.
2.2 Electrochemical measurements
The electrochemical measurements were performed
in a solution of 0.2 g/L Na2SO4, 0.2 g/L NaHCO3 and 0.2
g/L NaNO3. It was acidified with H2SO4 to pH 5 in order
to simulate the acid rain which is frequently found in
polluted urban environments (this atmosphere was desig-
nated: "simulated urban rain"). Four different electro-
chemical measurements were done for three different
types of bronzes.
A three–electrode corrosion cell was used with a vol-
ume of 350 cm3. The working electrode was embedded
in a Teflon holder, and had an exposed area of 0.785 cm2.
Saturated calomel electrode (SCE, E = 0.2415 V) was
used as reference electrode and two stainless steal rods
served as a counter electrode. For electrochemical tests a
Gamry 600 potentiostat/galvanostat, expanded with a
Gamry Instruments framework module was used.
Following initial 1 h stabilization at open circuit po-
tential (OCP), linear polarization measurements at ±20
mV vs. OCP with a scan rate 0.1 mV/s were performed.
Finally, electrochemical impedance measurements, the
frequency range ranged from 65 kHz to 5 MHz at 10 cy-
cles per decade, with an ac amplitude of ±10 mV, were
monitored. The absolute impedance and phase angle
were measured at each frequency. The impedance mea-
surements were carried out after different times of im-
mersion (1 h, 4 h, 8 h and 12 h) in the electrolyte. All po-
tentials are reported with respect to the SCE scale.
2.3 SEM/EDX analysis
For SEM/EDX analysis the different alloys were im-
mersed in a simulated urban rain, pH 5 for 35 d under
stationary conditions. At the end of exposure, the speci-
mens were taken from the solutions, rinsed with distilled
water and dried. Surface morphology was inspected and
analyzed with a low vacuum scanning electron micro-
scope (SEM, JSM 5500 LV, JOEL, Japan) at acceleration
voltage of 20 kV. The microscope was equipped with en-
ergy dispersive spectrometer (Inca, Oxford Instruments
Analytical, UK). EDS analysis was performed at an ac-
celeration voltage of 20 kV.
2.4 Metalographic examination
Samples were first grounded up to grades 2000, then
they were polished up to 4000 and finally 0.5 μm paste
was used. Etching for uncovering the microstructure was
performed using solution of ammonium and hydrogen
peroxide for 3 min. The sample was then immediately
rinsed with alcohol and dried with air.
E. [VARA FABJAN et al.: CORROSION STABILITY OF DIFFERENT BRONZES IN SIMULATED URBAN RAIN
586 Materiali in tehnologije / Materials and technology 45 (2011) 6, 585–591
Table 1: The composition of the three different bronzes in mole fractions, x/%
Tabela 1: Sestava razli~nih bronastih zlitin v molskih dele`ih, x/%
Bronze alloy /
Composition Cu Sn Zn Pb Si Al P Fe the rest
CuSnZn 87.0 5.5 3.6 0.1 – 0.6 – 0.1 3.1
CuSnZnPb 87.1 6.6 1.1 5.0 – – 0.1 0.1 0
CuSnSi 85.5 9.6 0.1 0.1 2.5 – 0.1 0.1 2.0
A CARL ZEISS AXIO Imager M2m optical metallo-
graphic microscope was used to inspect the micro-
structure of specimens. Metallographic specimens were
prepared and investigated in the longitudinal and trans-
verse direction of castings.
3 RESULTS AND DISCUSSION
3.1 Metalographic examination
Unleaded bronze, CuZnSn, revealed a non–homoge-
neous material consisting of dendrites of –tin and zinc
eutectics in the matrix of –copper. CuSnZn bronze
many microstructural imperfections, such as shrinkage
cavities, pores and non–uniform dendrite distribution
(Figure 1a and 1b).
Microstructure of CuSnZnPb bronze consisted of
similar eutectics and matrix as CuSnZn bronze. Lead
produced by the eutectic reaction was distributed
inter–dendritically in copper as small globules. Some
shrinkage cavities (black inter–dendritic network) could
be observed in the microstructure (Figure 1c and 1d).
Silicon bronze, denoted as CuSnSi, is a very homoge-
neous material in both longitudinal and transverse direc-
tion (Figure 1e and 1f). It consisted of dendrites of sili-
con and –tin eutectics in matrix of –copper. It was
reported that higher tin content in the alloy might im-
prove corrosion resistance of the bronze 7.
3.2 Electrochemical measurements
In order to evaluate corrosion properties of the three
different bronzes, electrochemical experiments were con-
ducted in a simulated urban rain, pH 5. For further evalu-
ation of corrosion behaviour, the samples were immersed
in the simulated urban rain solution for 35 d. After that,
the surface was examined by EDX/SEM technique.
3.2.1 Open circle potential measurement
During the stabilization process, the open circuit po-
tential was measured as a function of time. Figure 2 rep-
resents open circuit potential curves of all three investi-
gated bronzes, immersed in the urban rain solution with
pH 5. All curves show a similar electrochemical behav-
iour. The corrosion potential, EOC at the beginning of ex-
posure (up to 500 s) did not change much with time for
CuSnZn and CuSnSi. For leaded bronze CuSnZnPb it
moved to slightly more positive values at the beginning
of exposure. Then, the corrosion potential in all cases
moved to more negative values of potentials. After one
hour of immersion it stabilized at –0.025 V for bronze
CuSnZn and at a slightly more positive potential for
bronze CuSnZnPb at –0.020 V. Corrosion potential for
CuSnSi stabilized at –0.030 V after 1 h of immersion.
The observed decrease of the value of EOC in time might
be a result of formation of adsorbed layer at interface
bronze/electrolyte due to carbonate, sulphate and nitrate
ions in simulated urban rain. However, Eoc evolution is
quite regular, indicating a stable process occurring
through a relative stable layers that formed on investi-
gated bronze surfaces.
E. [VARA FABJAN et al.: CORROSION STABILITY OF DIFFERENT BRONZES IN SIMULATED URBAN RAIN
Materiali in tehnologije / Materials and technology 45 (2011) 6, 585–591 587
Figure 2: Open circuit potential measurements for the three different
bronzes: CuSnZn, CuSnZnPb and CuSnSi, immersed in simulated ur-
ban rain solution, pH 5.
Slika 2: Meritve pri potencialu odprtega kroga za tri razli~ne brone
CuSnZn, CuSnZnPb in CuSnSi, potopljene v simulirano raztopino
mestnega de`ja, pH 5.
Figure 1: Metalographic images of the three different bronzes
(CuSnZnPb– a and b, CuSnSi–c and d and CuSnZn– e and f), exposed
surface direction–longitudinal (a, c, e) and cross–sections–transverse
side (b, c, f). All shown images are etched, except b) and f), which are
polished.
Slika 1: Metalografski posnetki treh razli~nih vrst brona (CuSnZnPb–
a in b, CuSnSi–c in d ter CuSnZn– e in f). Longitudinalna smer je
izpostavljena povr{ina brona (a, c, e) in prerez brona (b, c, f). Broni na
prikazanih posnetkih so jedkani, razen na posnetkih b) in f), kjer je
prikazana bru{ena povr{ina.
3.2.2 Polarization resistance (Rp) measurements
One hour after immersion of CuSnZn, CuSnZnPb
and CuSnSi in simulated urban rain, pH 5, the linear po-
larization measurements at a scan rate of 0.1 mV/s were
executed (Figure 3). The slope of the curve poten-
tial–current is the polarization resistance value as de-
scribed in Stearn Geary equation (1):
R
E
jEp ( )Δ
Δ
Δ→
=0 (1)
The average measured values for the three bronzes
are presented in Table 2. The lowest value of polariza-
tion resistance (Rp = 2.3 k cm2) corresponded to leaded
bronze, CuSnZnPb and the highest value of polarization
resistance corresponded to silicon bronze, CuSnZnSi (Rp
= 4.6 k cm2). The lowest value of polarization resis-
tance for CuSnZnPb exhibited the highest corrosion sus-
ceptibility of leaded bronze, whereas values for Silicon
bronze CuSnSi showed higher corrosion resistance in
simulated urban rain than leaded bronze and bronze
CuSnZn. Similar observation was found by Chiavari and
coworkers 7. In their study, leaded bronze and silicon
bronze exhibited similar electrochemical behaviour at
initial stages but after 7 d, the corrosion rate of leaded
bronze was decreased. It was ascribed to formation of
uniform and stable passive layer on silicon bronze in
comparison with passive layer on leaded bronze.
Namely, the passive layer on silicon bronze hindered the
oxygen diffusion to the copper surface.
3.2.3 Electrochemical impedance spectroscopy (EIS)
measurements
Figure 4 represents the Nyquist diagrams for the
three different bronze samples: CuSnZn, CuSnZnPb and
CuSnSi at different immersion times during 12 h of im-
mersion in simulated urban rain solution with pH 5.
E. [VARA FABJAN et al.: CORROSION STABILITY OF DIFFERENT BRONZES IN SIMULATED URBAN RAIN
588 Materiali in tehnologije / Materials and technology 45 (2011) 6, 585–591
Table 2: Corrosion potential, Ecorr and polarization resistance values,
Rp, deduced from linear polarization measurements
Tabela 2: Korozijski potencial Ekor in vrednosti polarizacijskih upor-
nosti Rp, dobljenih iz meritev linearne upornosti
Material Ecorr /V Rp/(k cm2)
CuSnZn –0.032 3.2
CuSnZnPb –0.029 2.3
CuSnSi –0.038 4.6
Table 3: The estimated polarization resistance from EIS measure-
ments in dependence of the time, values of Rp/(k cm2)
Tabela 3: Ocenjene vrednosti polarizacijskih upornosti iz meritev EIS
v razli~nih ~asovnih obdobjih Rp/(k cm2)
Immersion time 1 h 4 h 8 h 12 h
Rp (CuSnZn) 15 13 23 25
Rp (CuSnZnPb) 3.4 3.4 3.1 2.9
Rp (CuSnSi) 5.2 5.0 5.9 9.2
Figure 3: Polarization resistance curves for CuSnZn, CuSnZnPb and
CuSnSi immersed in simulated urban rain, pH 5 at a scan rate 0.1
mV/s
Slika 3: Meritve linearne upornosti za CuSnZn, CuSnZnPb in CuSnSi,
potopljene v simulirano raztopino mestnega de`ja, pH 5, pri hitrosti
preleta 0,1 mV/s
Figure 4: EIS response as Nyquist plots for different bronzes: a)
CuSnZn, b) CuSnZnPb and c) CuSnSi after different immersion times
in a simulated urban rain solution, pH 5.
Slika 4: EIS-odziv v obliki Nyquistovih diagramov za razli~ne brone
CuSnZn, CuSnZnPb in CuSnSi po razli~nih ~asih potopitve v simuli-
rano raztopino mestnega de`ja, pH 5.
Nyquist diagram corresponding to bronze CuSnZn
was characterized by high frequency intercept at the
abscise axis and broad and semi-depressed loop at lower
frequencies (Figure 3). One time constant was observed
at 1 h immersion which then evolves with two clearly re-
solved time constants at 12 h immersion time. Estimated
polarization resistances (Rp) were determinated from im-
pedance modulus at the lowest value of frequencies, as
presented in Table 3. During 12 h immersion, impedance
values for bronze CuSnZn increased with time. Thus, the
estimated polarization resistance increased with immer-
sion time. Stable layer of protective products had
formed.
The Nyquist diagram for CuSnZnPb (Figure 4) was
characterized by high frequency intercept at the absise
axis, broad depressed loop at middle frequencies and
straight line at low frequencies. The high frequency in-
tercept is relatively high due to low conductivity of the
urban acid rain solution. The shape of impedance re-
sponse indicated the undefined time constant as observed
from Bode diagrams (not shown). Also, the phase shift is
very low for all times of immersion. Impedance response
is not highly dependant of the immersion time, but a
slight decrease is observed, as also presented in Table 3.
Nyquist diagram for CuSnSi showed similar impedance
behaviour to the sample CuSnZn including similar dia-
gram shape and increasing value of impedance response.
Also, the estimated polarization resistance value in-
creased with immersion time (Table 3).
Thus, our results showed that the corrosion rate of
unleaded bronze and silicon bronze decreased with time,
whereas corrosion rate for leaded bronze increased with
time. Impedance spectroscopy results show that unleaded
bronze showed slightly better corrosion characteristics
due to bigger estimated polarization resistances when
compared to silicon bronze. However, the appearance of
silicon bronze corrosion products is favourable over un-
leaded bronze, as shown in the next chapter of results.
Our results are similar to results reported by Chiavari
et al. 7. They investigated the silicon bronze (3 % Si) and
leaded bronze (5 % Pb) in acid rain solution pH 3.1 by
conducting different electrochemical measurements at
different immersion times in acid rain solution pH 3.1.
They found out that corrosion rate of silicon bronze de-
creased after long immersion time due to formation of
stable layer of protective corrosion product, whereas in
the case of leaded bronze the corrosion rate increased.
3.3 Surface characterization by SEM / EDX analysis
After 35–day immersion in simulated acid rain solu-
tion pH 5 under stagnant conditions, the samples of
CuSnZn, CuSnZnPb and CuSnSi were examined by
Scanning Electron Microscopy (SEM) and quantitatively
analysed by EDS analysis (Figure 5). Differences of pa-
tina formation were observed already with a naked eye
(not shown). Patina on unleaded bronze CuSnZn exhib-
ited orange type compact corrosion product, whereas pa-
tina on leaded bronze after 35 day immersion in urban
acid rain was brown–reddish and uneven. Patina that
formed on silicon Bronze CuSnSi looked like smoky
brown colour and very compact and thin.
The SEM image of sample CuSnZn after 35 d of ex-
posure to urban acid rain (Figure 5a) showed homoge-
neously coated surface with corrosion products. The
EDS analysis indicated the presence of mainly copper,
oxygen and carbon (mole fractions x: 41 %, 35 % and
10 %), whereas Zn, Sn, S, Si were present as minor ele-
ments. Due to high percentage of oxygen in the analysed
layer, the bronze surface presumably contained other
corrosion products apart from cuprite, Cu2O. The possi-
ble proposed mineral was brochantite 16. The SEM image
of leaded bronze, CuSnZnPb (Figure 5b), after 35 day
immersion in acid rain solution exhibited two different
E. [VARA FABJAN et al.: CORROSION STABILITY OF DIFFERENT BRONZES IN SIMULATED URBAN RAIN
Materiali in tehnologije / Materials and technology 45 (2011) 6, 585–591 589
Figure 5: SEM image of a) CuSnZn b) CuSnZnPb and c) CuSnSi af-
ter 35 d of immersion in simulated urban rain solution pH 5
Slika 5: SEM-prikaz a) CuSnZn b) CuSnZnPb and c) CuSnSi po 35 d
izpostavitve v simulirani raztopini mestnega de`ja, pH 5
areas, denoted as area No. 1b and area No. 2b. Area No.
1b contained mainly carbon, less oxygen and copper, and
nitrogen and lead in trace amounts. That indicated to
copper carbonate as possible corrosion product. Whereas
darker area, denoted as area No. 2b, was rich in copper
(x/% : 51 % Cu, 2 % Sn, 27 % C, 9 % N and 11 % O)
and contain more tin than area No. 1. The darker area
could be the base alloy, not yet entirely covered by cor-
rosion products. Cuprite and SnO2 were proposed as the
most possible oxidation products.
The SEM micrograph of CuSnSi (Figure 5c) showed
three different areas, denoted as Area No. 1c, Area No.
2c and Area No. 3c. All analysed areas examined con-
tained similar elements but in different ratios. Corrosion
products in the shape of white spherical particles (Area
No. 1) are rich in oxygen, carbon, nitrogen and sulphur.
The formation of sulphates (brohantite Cu4SO4(OH)6 or
naukarite Cu4(SO4)4(CO3)(OH)6·48 H2O) was proposed.
Grey coloured area, denoted as Area No.2c, is rich in tin,
silicon and copper and less rich in oxygen and sulphur,
indicated the formation of SiO2 and SnO2. Area No.3c is
in composition similar to area No. 1c, but contained less
sulphur.
It is difficult to unambiguously differentiate the cor-
rosion products, grown during 35 d immersion in urban
acid rain on the three different bronzes as the corrosion
layers are too thin to detect the EDS signal just from cor-
rosion products. However, the visual inspection of the
patina layers on the three different bronzes showed that
the patinas were different as the result of the base alloy
composition.
Chiavari et al 7 who studied leaded bronze and silicon
bronze in climatic chamber with chlorides and sulphates,
showed that the main corrosion product on leaded bronze
was brochantite, and that the layer on Silicon bronze was
thinner and more compact. Also, similarly to our case,
the corrosion products on leaded bronze were more
non-homogeneous and thicker.
As the result of microstructural characteristics of the
three different bronzes, it can be concluded that surface
imperfections, non homogeneous distribution of corro-
sion products formed in the case of leaded bronze
CuSnZnPb. Even distribution of corrosion products on
Silicon bronze is also a result of homogeneous distribu-
tion of metallographic phases on silicon bronze, whereas
unleaded bronze CuSnZn behaved somehow in between.
4 CONCLUSIONS
Three different bronzes were investigated in simu-
lated urban rain solution, pH 5: recent bronze CuSnZn,
leaded bronze CuSnZnPb and newer representative of
bronzes, silicon bronze CuSnSi.
Microstructural investigation showed that recent
bronze CuSnZn consisted of –tin and zinc eutectics in
–copper. Inter dendritic network of shrinkage cavities
were observed as an alloy imperfection. Similar observa-
tion occurred for leaded bronze, where lead as globules
intermixed in –copper of the bronze alloy. Silicon
bronze has extremely oriented microstructure which is
very regular. It consisted of dendrites of silicon and –tin
eutectics in matrix of –copper.
Electrochemical investigation of different bronzes in
simulated urban rain solution showed that the leaded
bronze is most likely to corrode the most, followed by
recent bronze CuSnZn. The highest corrosion resistivity
among the bronzes was found for silicon bronze,
CuSnSi.
Time dependence during 12 h immersion was fol-
lowed by electrochemical impedance spectroscopy mea-
surements. It was showed that corrosion resistance de-
creased with time for leaded bronze, whereas it increased
for unleaded bronze CuSnZn and silicon bronze CuZnSi.
Morphological observation of corrosion products that
formed during 35 d immersion in stagnant solution of ur-
ban acid rain showed that the patina layer is different for
the three different bronzes. It is non–homogeneous for
leaded bronze and very thin and compact for silicon
bronze. It is believed that alloy composition and mi-
cro-structural characteristic have a great impact on elec-
trochemical behaviour and formation of oxidation prod-
ucts–bronze patina layer. It affects visual appearance as
well as chemical composition of corrosion products.
Unleaded bronze can successfully replace the use of
leaded bronze where also better corrosion characteristics
were found. Silicon bronze can be used in environments
where only small changes of appearance of corrosion
products are expected.
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25 Strandberg, H., Reactions of copper patina compounds–I. Influence
of some air pollutants. Atmospheric Environment, 32 (1998) 20,
3511–3520
26 Strehblow, H. H.; Titze, B., The investigation of the passive behav-
iour of copper in weakly acid and alkaline solutions and the exami-
nation of the passive film by esca and ISS. Electrochimica Acta, 25
(1980) 6, 839–850
27 Varvara, S.; Muresan, L. M.; Rahmouni, K.; Takenouti, H., Evalua-
tion of some non–toxic thiadiazole derivatives as bronze corrosion
inhibitors in aqueous solution. Corrosion Science, 50 (2008) 9,
2596–2604
28 Watanabe, M.; Toyoda, E.; Handa, T.; Ichino, T.; Kuwaki, N.;
Higashi, Y.; Tanaka, T., Evolution of patinas on copper exposed in a
suburban area. Corrosion Science, 49 (2007) 2, 766–780
E. [VARA FABJAN et al.: CORROSION STABILITY OF DIFFERENT BRONZES IN SIMULATED URBAN RAIN
Materiali in tehnologije / Materials and technology 45 (2011) 6, 585–591 591

D. KEK MERL et al.: MORPHOLOGY AND CORROSION PROPERTIES PVD Cr-N COATINGS ...
MORPHOLOGY AND CORROSION PROPERTIES PVD
Cr-N COATINGS DEPOSITED ON ALUMINIUM ALLOYS
MORFOLOGIJA IN KOROZIJSKE LASTNOSTI CrN
PVD-PREVLEK, NANESENIH NA ALUMINIJEVE ZLITINE
Darja Kek Merl1, Ingrid Milo{ev1, Peter Panjan1, Franc Zupani~2
1Jo`ef Stefan Institute, Jamova 39, 1000 Ljubljana, Slovenia
2University of Maribor, Faculty of Mechanical Engineering, Smetanova 17, 2000 Maribor, Slovenija
darja.kek@ijs.si
Prejem rokopisa – received:2011-04-07; sprejem za objavo – accepted for publication: 2011-05-30
The attempt to find an alternative coating for corrosion protection of Al- alloys was made. PVD coatings are one of the possible
alternatives for replacement of ecological unfriendly chromate coatings. Chromium-nitride (Cr-N) and Ni/Cr-N coatings were
sputtered on aluminium substrates (AA7075 and cladded AA2024). Surface and sub-surface characterizations were performed
by AFM and SEM. Special attention was given to defects incorporated into coatings, since they play important role in the
corrosion protection of the coating/substrate systems. The cross-sections through the typical defects were performed by ion
beam milling incorporated into the SEM. The Vickers hardness of the Cr-N with and without layer of Ni on both substrates was
determined. After the coatings deposition, the values of Vickers hardness (10 mN load) increase for 10 to 100-fold compared to
the substrates. The corrosion behaviour of Cr-N and Ni/Cr-N thin films was investigated in near neutral 0.1 M solution of NaCl
using potentiodynamics electrochemical measurement. Cr-N and Ni/Cr-N coatings shift the corrosion potentials to more positive
values. The best corrosion resistance among the tested coating/substrate systems were found for Ni/Cr-N on AA7075 substrate.
Keywords: Al-alloys, corrosion properties, CrN films, FIB, PVD coatings
Ena od mo`nih alternativ za zamenjavo ekolo{ko neprijaznih kromatnih prevlek na aluminijevih zlitinah so PVD-prevleke. V
prispevku opisujemo pripravo in karakterizacijo prevlek Cr-N in Ni/Cr-N na aluminijeve podlage (AA7075 in AA2024).
Povr{insko in podpovr{insko karakterizacijo CrN-prevlek smo izvedli z vrsti~no elektronsko mikroskopijo (SEM) in
mikroskopom na atomsko silo (AFM). Posebna pozornost je bila namenjena karakterizaciji defektov v prevlekah, saj igrajo
pomembno vlogo pri korozijski za{~iti sistema prevleka/podlaga. Defekte smo karakterizirali z vrsti~nim mikroskopom (SEM),
ki je dodatno opremljen s fokusiranim ionskim curkom (FIB) za odstranjevanje materiala. Dolo~ili smo trdoto prevlek po
Vickersu na obeh vrstah aluminijevih podlag. Korozijske lastnosti prevlek smo merili v 0,1 M raztopini NaCl s
potenciodinamskimi krivuljami. Cr-N- in Ni/Cr-N-prevleke premaknejo korozijski potencial proti pozitivnim vrednostim.
Za{~ita aluminijevih zlitin je bolj{a z dvojno prevleko Ni/CrN kot samo s prevleko CrN.
Klju~ne besede: Al-zlitina, korozijske lastnosti, CrN, FIB, PVD-prevleke
1 INTRODUCTION
Aluminium alloys are very important engineering
materials employed in a variety of applications, which
include automotive, constructive, chemical, petroche-
mical, cryogenics, transportation equipment, etc. Among
other attractive properties, these materials exhibit
relatively high strength to weight ratio, good corrosion
resistance and high thermal conductivity. However, due
to their low hardness, wear and abrasion resistance, the
application of these materials to sliding part is quite
limited. Conventionally, the improvement of the surface
properties of aluminium alloys, particularly those
involved in aircraft and aerospace application, like
AA2024 and AA7075, has been the use of electrolytic
hard chromium deposition and conversion chromate
coating. However, the strong environmental regulations
issued in the past few years to decrease significantly the
emissions of the highly toxic hexavalent chromium, have
led to the development of different technologies also able
to improve the surface properties of such materials, but
more friendly from the environmental point of view.1–5 On
the other hand, conversion chromate coating possesses
advantageous properties, like self-healing effect in case
of mechanical coatings damage.
PVD technologies offer a promising alternative for
the production of cost-effective, quality coatings with
dry, clean and environment-friendly technology fully
supported by legislation on environmental protection.
The properties of various PVD coatings (metals, alloys,
nitrides and oxides, carbides) have been well docu-
mented and systematically presented.6,7 Ceramic mate-
rials have oriented covalent bonds; such materials are
hard, brittle, chemical inert and with high melting-point
temperature. Their wear resistance is excellent. Because
of these properties, PVD coatings are in many cases a
serious alternative to electrochemical coatings. For
example, 1–5 μm thick layer of PVD coatings can
replace 250 μm thick layer of hard chrome.
In the recent years, many efforts have been oriented
to the development of anticorrosive coatings deposited
by reactive sputtering in order to protect aluminium
alloys like AA2024 and AA7075 used in aircraft appli-
cation. Among conventional nitride and carbide PVD
hard coatings, TiN and AlN sputtered on aluminium
alloys were investigated. Diesselberg et al.8 sputtered
Materiali in tehnologije / Materials and technology 45 (2011) 6, 593–597 593
UDK 669.71:620.193:621.9 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 45(6)593(2011)
TiNx varying concentration of nitrogen and bias voltage
to get different microstructructure of TiNx. The lower
concentration of nitrogen and the deposition at higher
bias shows a positive effect on the corrosion protection.
AlNx sputtered coatings were investigated by Schafer
and Stock9. AlNx coatings on Al-substrate with higher
concentration of nitrogen shift the pitting potential to
more positive values. However, salt spray test showed
opposite effect of nitrogen concentration; the coatings
with less concentration of nitrogen are more stable. Liu
et al. 10 studied effect of nitrogen ion implanted into
aluminium substrate on the mechanical properties of TiN
films. Implantation of nitrogen ion, increase the adhesion
of TiN coating, therefore improve surface properties of
aluminium by forming 80 nm thick AlN gradient layers.
Corrosion properties were not reported.
On the other hand, sputtered chromium and reac-
tively sputtered chromium nitride (CrN) act as very
promising candidates due to their high internal resistance
to corrosion, which arises from the instant formation of
an oxide layer on the surface at room temperature.11,12,13
In this paper, we present the investigation of PVD
coating CrN and Ni/CrN for corrosion protection of
aluminium alloys AA2024-cladded and AA7075. CrN
thin films on aluminium substrates were prepared by
sputtered deposition at low temperature. Surface and
sub-surface characterization were performed on the
coatings. Mechanical properties such as hardness and
roughness were determined for each coating. The
corrosion testing of PVD coatings, using electrochemical
methods (potentiodynamic measurement), were
performed in the 0.1M solution of NaCl.
2 EXPERIMENTAL
2.1 Substrate preparation
As substrates, the disks of aluminium alloys
(AA2024 cladded and AA7075) were used. All the
Al-substrates were chemically cleaned with a mildly
alkaline cleaner Ridoline 1 % in an ultrasonic bath at 60
°C for 4 min, followed with distilled water bath and
ethanol bath and drying with N2. To study the effects of
the surface preparation on corrosion performance of
PVD coatings, the grinding (1000 and 4000 SiC papers)
and polishing (0.25 μm diamond paste) of the substrate
surface were performed in some cases. Ion etching to
remove native oxide layer usng Ar + H2 mixture (15 min)
in vacuum chamber was performed in some cases.
2.2 Coating depositions
All coatings were prepared by sputtering at 150 °C in
the depositing system with termionic arc (Sputron,
Balzers). Cr-N and Ni/Cr-N were deposited on the
substrates with and without ion etching. Additional
interface layer of amorphous carbon (a-C) were
performed to reduce the roughness of the substrates. The
thickness of Cr-N coatings was 2,5 μm and Ni/Cr-N
coating about 3 μm.
The vacuum chamber was evacuated to a base
pressure of approximately 2 mPa. Nitrogen (purity
99.995 %) flow remained constant at 9.5 cm3. During
deposition the bias voltage and the substrate temperature
were -30 V and 150 °C
2.3 Characterisation
Field emission scanning electron microscope (Zeiss
Supra 35 VP) was used for study of the defect morpho-
logy in planar surface view and cross-sectional fracture
view.
Focused ion beam (FIB) workstation14 was used to
prepare cross-section through the defects. We used FIB
integrated in FEI QUANTA 200 3D microscope. Ion
beam was used to remove precise sections of material
(close to the selected defect) from the specimen surface
by sputtering.
The surface morphology and roughness of the sub-
strates and coating/substrate systems was also examined
by atomic force microscope (Solver PRO).
Hardness was determined using the nanoindenter
Fischerscope H100C. The load was varied between
10–1000 mN.
The electrochemical corrosion behaviour was studied
using potentiodynamic polarisation tests in a solution of
0.1 M NaCl. Autolab three-electrode corrosion cell was
used, with the working electrode embedded in a Teflon
holder. An Ag/AgCl electrode served as a reference
electrode and carbon rods as counter electrodes. The
working electrode (WE) was a substrate with an area of
0.785 cm2 that was coated with Cr-N and Ni/Cr-N. The
polarisation curves were measured after 1 h of stabili-
zation at the corrosion potential so that a quasi-stable
potential was reached. The curves were obtained by
sweeping the potential from the cathode to the anode.
The sweep-rate setting was 1 mV/s using an EG&Par
PC-controlled potentiostat /galvanostat Model 263 and
Powersuite software.
3 RESULTS AND DISCUSSION
3.1 Microstructure
A typical surface morphology of PVD-hard coatings
Ni/CrN on AA-7075 alloy is shown on Figure 1a.
Top-view image is possible to distinguish typically
growth defects for PVD-coatings: a) spherical droplets-
cone structure, coming above the surface, typical wide
size from 1 to few μm, b) pin- holes apparently going to
the substrate, reaching the size up to the few μm c) big,
not deep craters of wide size of 10–40 μm.
High surface roughness of Al-alloy substrate is
shown on cross section image on Figure 1b. Surface
roughness of AA-7075 substrate measured by AFM was
Ra  14 nm, AA2024–cladd Ra  11.3 nm. After
D. KEK MERL et al.: MORPHOLOGY AND CORROSION PROPERTIES PVD Cr-N COATINGS ...
594 Materiali in tehnologije / Materials and technology 45 (2011) 6, 593–597
deposition surface roughness increased for around Ra 
34 nm on AA7075 substrate and Ra  26 nm on
AA2024-cladd.
Microstructure characteristics of PVD-coatings,
where surface defects have to be taken into account, play
important role in the determination of corrosion pro-
perties of thin films. A few authors pay attention to the
defect density and try to explain the origin of the growth
defect on the PVD-coatings. To determine the origin of
the defects the FIB cross section analysis were made on
the two typical defects shown on Figure 2. The first
defect (Figure 2a–d) has the form of a cone, while the
second one (Figure 2e–h) resembles a pinhole. The
question was weather these defects extend through the
whole coating or not, and what is the origin for their
nucleation. After consecutive serial sectioning, we found
that the first defect started to grow in the middle of the
coating due to the incorporation of a foreign particle.
Figure 2c shows that the region under the cone is not
completely filled with material. The second defect
extends through the whole coating and originates in a
small hole in the substrate (Figure 2e). It is well known
that PVD processes have a poor ability to cover a small
hole due to shadowing effect. Figure 2e–h clearly shows
that a small crater on the substrate surface cannot be
covered completely and that a pinhole extending through
the whole coating was formed. The corrosion takes place
on such defects, while solution can reach the base mate-
rial. The origin of the pinholes could be the corrosion
products (wet cleaning) or dust particles that they are not
removed from the substrate surface during the deposi-
tion. The high roughness of selected substrates could
result the same effects. The cone microstructure,
appearing quite an accident anywhere in the growing
layer, (Figure 2b) are due to inclusions coming on the
growing layer from the vacuum chamber, i.e. drops of
metal target.
The origin of the big not deep craters (Figure 1a) can
be explained by the falling away of the particles
(corrosion products due to wet cleaning, dust particles)
from the surface, soon after the starting the deposition.
3.2 Mechanical characterization
The Vickers hardness of the Cr-N, Ni and Ni/Cr-N
coating, with deposited on AA2024 cladd and AA7075
substrates was determined. The bare substrate
D. KEK MERL et al.: MORPHOLOGY AND CORROSION PROPERTIES PVD Cr-N COATINGS ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 593–597 595
Figure 3: Hardness of the substrates and coatings for a) AA2024-
cladded b) AA7075
Slika 3: Trdote po Vickersu pripravljenih prevlek: a) podlaga
Al-AA2024, b) podlaga Al-AA7075
Figure 2: FIB cross section analysis were made on the two typical
defects of coatings a-d) cone structure and e-h) pin-hole
Slika 2: SEM posnetek prerezov (narejenih s FIB) dveh tipi~nih napak
v prevleki; a)–d) konusna oblika napake in e)–h) luknjica (pin-hole)
Figure 1: A top view of a) Cr-N coating on aluminium alloy AA2024
cladded substrate b) cross-section of Ni/Cr-N/Al-alloys
Slika 1: SEM posnetek a) mikrostrukture povr{ine Cr-N-prevleke na
aluminijevi zlitini AA2024, b) prereza Ni/CrN/Al-zlitina
AA2024-cladd (Figure 3a) has a hardness of 40 HV and
AA7075 substrates of 200 HV (Figure 3b). After the
coating deposition, the values of Vickers hardness,
measured at 10 mN, increase for 10 to almost 100-fold
compare the substrates and then progressively dimi-
nishes at higher loads. In both cases, this is only a
surface effect due to small thickness of the coatings
(Figure 3a and Figure 3b).
3.3 Corrosion characterisation
The results from the potentiodynamic measurements
of the CrN and Ni/Cr-N coatings on the substrates of
aluminium alloys (AA2024-cladd and AA7075) in the
0.1M aqueous sodium chloride solution are shown in
Figure 4. After 1 h of stabilization at the open-circuit
potential, the corrosion potential (Ecorr) for the
AA2024-cladd substrate in was approximately –0.67 V
(Figure 4a). Following the Tafel region, the alloy
exhibited a narrow range of passivation, with breakdown
potential Eb of about –0.63V, due to native oxide layer
formed on pure (cladded) aluminium. In the case of the
AA7075, the Ecorr in 0.1 M NaCl was approximately –
0.70V.
Ecorr shifted to a more positive potential with the
application of the Cr-N and Ni/Cr-N coatings, reaching
-0.62 V and -0.63 V versus on the AA2024-cladd
(Figure 4a). The Ecorr of the CrN and Ni/Cr-N coatings
on the AA7075 reached a value of -0.67 V and -0.64 V
(Figure 4b).
The corrosion current density is often used as an
important parameter to evaluate the kinetics of corrosion
reactions. Corrosion protection is normally proportional
to the corrosion current density (io) measured via
polarization. In our case, where PVD coatings of Cr-N
are chemically not reactive, the corrosion current density
indicates pores in the coatings, where the electroche-
mical reaction of the substrate takes place. The io values
from Figure 4 show that the Cr-N coating on on both
Al-alloys corroded faster than the Ni/CrN coating under
open-circuit conditions; this means that the active area of
the substrate due to coating porosity is higher in the case
of Cr-N than in the case of Ni/Cr-N coatings. However,
the overall corrosion process was still dominated by
locally active dissolution, which occurred as the
substrate was exposed via coating porosity.
Cr-N coating did not improve the corrosion pro-
perties for aluminium substrates in the extent as is well
known for corrosion protection of steel substrate6. The
effects of the surface preparation (grinding, polishing,
etching) and interface layers were investigated further on
AA7075 as is shown on Figure 5. Performance of Cr-N
coatings can be slightly improved with the proper surface
preparation (grinding, polishing) as is shown with the
green curves (green and yellow curves). No effect of
etched and non-etched substrate, with (Ar+H2) plasma in
the vacuum chamber before the deposition, was observed
D. KEK MERL et al.: MORPHOLOGY AND CORROSION PROPERTIES PVD Cr-N COATINGS ...
596 Materiali in tehnologije / Materials and technology 45 (2011) 6, 593–597
Figure 5: PD curves of Cr-N coating with the different surface
substrate preparation as indicate in the figure
Slika 5: PD-krivulje Cr-N z razli~no predpripravo podlag, kot so
ozna~ene na sliki
Figure 4: PD curves of various coatings, as indicated on figures, for a)
AA2024-cladded substrate b) AA7075 substrate
Slika 4: PD-krivulje prevlek, ozna~enih na sliki za a) podlage
Al-AA2024, b) podlage Al-AA7075
Table 1: Corrosion potential (Ecorr) and corrosion current density (io)
after 1 h stabilisation for uncoated and coated substrates on Figure 4.
material Ecorr/V io/(μA cm–2)
AA7075 –0.707 2.10
AA7075/CrN –0.679 0.99
AA7075/Ni/NiCrN –0.645 0.41
AA2024-cladd –0.674 0.18
AA2024-cladd /CrN –0.631 0.38
AA2024-cladd /Ni/NiCrN –0.624 2.80
on the PD performance of Cr-N protective layer (Figure
5). The deposition of an interface layer could influence
the performance of the coatings as well. Thin interface
layer of a-C (amorphous carbon) was sputtered in order
to reduce the high roughness of the substrate. A slight
improvement was observed. PD curves show that
although there is some effect of surface preparation on
the corrosion performance of Cr-N, Cr-N alone is not
suitable coating for corrosion protection of soft alumi-
nium substrates.
4 SUMMARY
The deposition of Cr-N and Ni/Cr-N was carried out.
Physical, surface analytical and electrochemical
measurements were performed on coated and uncoated
AA2024-cladded and AA7075 substrates. Cr-N and
Ni/Cr-N coatings increase the surface hardness of both
Al-alloys substrates to 10-fold. At the measure loads of
10 mN.
FIB cross section analyses were made on the two
typical defects of coatings (cone structure and pin-hole).
Cone microstructure, appearing quite an accident in the
growing layer, are due to inclusions coming on the
growing layer from vacuum chamber (drops of metal
target, dust, etc). A small crater (probably due to surface
roughness) on the substrate surface cannot be covered
completely due to shadowing effect and that a pinhole
extending through the whole coating was formed. The
corrosion takes place on such defects, while solution can
reach the base material.
Corrosion behaviour of Cr-N and Ni/Cr-N thin films
was investigated in near neutral 0.1 M solution of NaCl
using potentiodynamics electrochemical measurement.
Cr-N and Ni/Cr-N coatings shift the corrosion potentials
to the more positive values. The best corrosion resistance
among the tested coating/substrate systems were found
for Ni/Cr-N on AA7075 substrate.
Acknowledgements
The work was supported by the Slovenian Research
Agency (project No. L2-9189).
5 REFERENCES
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ASM International, Materials Park, OH, 1994
2 K. O. Legg, M. Graham, P. Chang, F. Rastagar, A. Gonzales, B. D.
Sartwll, Surf. Coat.Technol., 81 (1996), 99
3 R. L. Twite, G. P. Bierwagen, Prog . Org. Coat., 33 (1998), 91
4 P. M. Natishan, S. H. Lawrence, R. L. Foster, J. Lewis, B. D.
Sartwell, Surf. Coat.Technol., 130 (2000), 218
5 H. J. Gibb, P. S. J. Lees, P. F. Pinsky, C. B. Rooney, Am. J. Ind.
Med., 38 (2000), 15
6 B. Blushan, B. K. Gupta, Hard Coating, Handbook of Tribology,
McGraw Hill, New York, 1991, Chapter 14
7 G. Jetson, Handbook of Thin Film Process Technology, Institute of
Physics, Bristol, 1996
8 M. Diesselberg, H-R. Stock, P. Mayr, Surf. Coat.Technol., 177–178
(2004), 399
9 H Schaefer, H-R. Stock, Corr. Sci., 547 (2000), 953
10 Y. Liu, L. Li, M. Xu, Q. Chen, Y. Hu, X. Cai, P. K. Chu, Surf. Coat.
Technol., 200 (2006), 2672
11 B. Navin{ek, P. Panjan, I. Milo{ev, Surf. Coat.Technol., 116–119
(1999), 476
12 D. K. Merl, M. Cekada, P. Panjan, M. Macek, Electrochimica Acta,
49 (2004), 1527
13 M. ^ekada, P. Panjan, D. Kek Merl, M. Ma~ek, Mater. Tehnol., 37
(2003) 5, 213–216
14 J. M. Cairney, P. R. Munroe, M. Hoffman, Surf. Coat.Technol., 198
(2005), 165
D. KEK MERL et al.: MORPHOLOGY AND CORROSION PROPERTIES PVD Cr-N COATINGS ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 593–597 597

F. KAVICKA et al.: THE EFFECT OF ELECTROMAGNETIC STIRRING ON THE CRYSTALLIZATION ...
THE EFFECT OF ELECTROMAGNETIC STIRRING ON
THE CRYSTALLIZATION OF CONCAST BILLETS – II.
VPLIV ELEKTROMAGNETNEGA ME[ANJA NA KRISTALIZACIJO
KONTINUIRNO ULITIH GREDIC – II.
Frantisek Kavicka1, Karel Stransky1, Bohumil Sekanina1, Josef Stetina1, Vasilij
Gontarev2, Tomas Mauder1, Milos Masarik3
1Brno Technological University, Technicka 2, 616 69 Brno, Czech Rep.
2University of Ljubljana, A{ker~eva 12, 1000 Ljubljana, Slovenia
3EVRAZ VITKOVICE STEEL, a. s., Stramberska 2871/47, 709 00 Ostrava, Czech Rep.
kavicka@fme.vutbr.cz
Prejem rokopisa – received:2011-02-01 ; sprejem za objavo – accepted for publicatio: 2011-03-03
The solidification and cooling of a continuously cast slab and the simultaneous heating of the mold is a very complicated
problem of three-dimensional (3D) transient heat and mass transfer. The solving of such a problem is impossible without
numerical models of the temperature field of the concasting processed through the concasting machine. Experimental research
and measurements have to take place simultaneously with the numerical computation, to be confronted with the numerical
model and make it more accurate throughout the process. An important area of the caster is the secondary cooling zone, which
is subdivided into thirteen sections. In this zone, where the slab is beginning to straighten, the breakout of the shell can occur at
points of increased local chemical and temperature heterogeneity for the steel, from increased tension as a result of the bending
of the slab and also from a high local concentration of non-metal and slag inclusions. The changes in the chemical composition
of the steel during the actual concasting are particularly dangerous. In the case of two melts, one immediately after the other,
this could lead to an immediate interruption in the concasting and a breakout. The material, physical, chemical and
technological parameters, which differed in both melts, were determined. If the dimensionless analysis is applied for assessing
and reducing the number of these parameters, then it is possible to express the level of risk of the breakout as a function of five
dimensionless criteria.
Keywords: concast slabs, oscillation marks, hooks, chemical composition, breakout, criteria, electromagnetic stirring,
crystallization
Strjevanje in ohlajevanje kontinuirno ulitega slaba in isto~asno ogrevanje kokile je zelo zapleten problem pri tridimenzionalnem
(3D) prenosu toplote in mase. Re{itev takega problema je nemogo~a brez uporabe numeri~nih modelov temperaturnega polja pri
kontinuirnem ulivanju. Eksperimentalne raziskave in meritve se morajo dogajati isto~asno z numeri~nim ra~unom, ne le zaradi
primerjave z numeri~nim modelom, ampak tudi zaradi ve~je natan~nosti samega procesa. Pomembno podro~je livnega stroja je
t. i. sekundarna hladilna cona, ki je razdeljena v trinajst presekov. V sekundarni hladilni coni lahko nastane preboj skorje zaradi
pove~ane lokalne kemijske in temperaturne heterogenosti jekla, porasta napetosti v slabu zaradi upogibanja in velikih lokalnih
koncentracij nekovinskih vklju~kov in vklju~kov `lindre. Posebno nevarne so spremembe kemijske sestave jekla med dejanskim
kontinuirnim ulivanjem. V primeru dveh talin, ki sledita ena takoj za drugo, lahko nastane takoj{nja prekinitve kontiulivanja in
preboja. Dolo~eni so bili fizikalni, kemijski in tehnolo{ki parametri snovi, v katerih se obe talini razlikujeta. Z uporabo
brezdimenzijske analize za ocenitev in zmanj{anje {tevila teh parametrov, je mogo~e izraziti stopnjo tveganja preboja kot
funkcijo petih brezdimenzijskih meril.
Klju~ne besede: kontinuirno uliti drogovi, nihajo~e oznake, kemijska sestava, preboj, merila, elektromagnetno me{anje,
kristalizacija
1 INTRODUCTION
Oscillation marks are transverse grooves forming on
the surface of the solidifying shell of a concast slab. The
course of the individual marks is rough and perpen-
dicular to the direction of the movement of the slab. The
formation of the marks is sometimes the result of the
bending of the solidifying shell during the oscillation of
the mould, which depends on the frequency and the
amplitude of the oscillation and on the casting speed.
The hooks are solidified, microscopically thin, surface
layers of steel 1–3 covered with oxides and slag, and their
microstructures are different to that of the solidifying
shell. The formation of the oscillation marks and hooks
is related. The depth of the oscillation marks and also the
shape, size and the microstructure of the hooks vary
irregularly. An increasing extent of these changes leads
to a defect in the shape of a crack, which reduces the
thickness of the solidified shell of the slab upon its exit
from the mould and causes a dangerous notch. In the
secondary-cooling zone, where the slab is beginning to
straighten out, a breakout of the steel can occur at points
of increased local chemical and temperature hetero-
geneity of the steel, from increased tension as a result of
the bending of the slab and also a high local concen-
tration of non-metal, slag inclusions. The changes in the
chemical composition of the steel during the actual
concasting are particularly dangerous. The consequences
of this immediate operational change in the chemical
composition of the steel, which are not prevented by a
breakout system directly inside the mould, could lead to
an immediate interruption of the concasting and a
Materiali in tehnologije / Materials and technology 45 (2011) 6, 599–602 599
UDK 621.74.047:669.14.018 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 45(6)599(2011)
breakout at a greater distance from the mould than usual,
thus leading to a significant material loss and downtime.
2 INTERRUPTION OF CONCASTING
This case was recorded during the process of
concasting of (250 × 1530) mm steel slabs of quality A
with the mass fractions of carbon content 0.41 % and
9.95 % chromium content (melts 1 to 3) and quality B
steel with 0.17 % carbon content and 0.70 % chromium
content (melt 4). The casting of the first two melts of
quality A took place without any significant problems,
after the casting of the third melt of quality A, the fourth
melt of quality B followed. The change in the chemical
compositions of the steels of both qualities was carried
out very quickly by changing the tundish. Inside the
mould, steel B mixed with steel A of the previous melt.
The pouring continued for another 20 min, but then, af-
ter, in the unbending point of the slab, at a distance of
14.15 m away from the level of the melt inside the
mould, there occurred a breakout between the 7th and 8th
segments and the caster stopped. The difference in the
height between the level inside the mould and the break-
out point was 8.605 m. This tear in the shell occurred on
the small radius of the caster. A 250-mm-thick sample
was taken from the breakout area using a longitudinal
axial cut (Figure. 1). The structure of this sample was
examined and using the Bauman print the distribution of
sulphur was analyzed too. The numbers 1 to 11 indicate
the positions of the samples in the places around the
breakout intended for the analysis. Simultaneously, sig-
nificant 25-mm sulphide segregations were discovered
(see Figure 1, position 6) – very heterogeneous areas
created by the original base material of the slab (melt 3),
the new material of the slab (melt 4) and between them
and also by the areas of mixed composition. Beneath the
surface of the slab, at a depth of 75–85 mm, there were
cracks and a zone of columnar crystals oriented towards
the surface of the slab on the small radius. This was
identical to the orientation of the groove, which gradu-
ally turned into a crack (Figure 1 – direction 4–6) and,
on the opposite surface of the slab, the hook which was
covered by the melt (position 8). In the first phase of the
analyses, the aim was to determine the material, physi-
cal, chemical and technological parameters, for which
both melts 3 and 4 differed (besides the already intro-
duced chemical composition). Table 1 contains the indi-
vidual parameters of both melts.
3 DIMENSIONLESS CRITERIA
If the method of dimensionless analysis is applied for
assessing and reducing the number of parameters in
Table 1 in the first approximation, then it is possible to
express the level of risk of breakout as a function of the
five dimensionless criteria contained in Table 2 (units m,
kg, s, K).
4 SUSCEPTIBILITY TO BREAKOUT –
BREAKOUT RISK
The risk of breakout grows in accordance with the
first criterion in direct proportion to the latent heat L
released from the mushy zone and inversely propor-
tionally to its dynamic viscosity . The second criterion,
i.e., the Strouhal number, includes transient, oscillation
movement, including the amplitude of the mould and
also, implicitly, a susceptibility to marks and hooks,
which precede the breakout. The third criterion has a
similar significance but, in addition, also includes the
dynamic viscosity. The first three criteria increase the
risk of breakout with melt 4 more than with melt 3. The
fourth criterion characterizes the reduction of the
load-bearing cross-section of the slab (by 28.1 % in
melt 3 and by 21.4 % in melt 4) by creating a mushy
zone, which indicates a greater risk of breakout in
melt 3. The last criterion considers the effect of the
mixture zone of melt 3 and a common effect of the
mixture zone of melts 3 and 4. The first three criteria are
of a dynamic nature and their product in melt 3 is 1.044
× 106, while in the fourth melt it is 1.502 × 106, i.e., the
mixture melt has a 50 % greater risk of breakout. The
product of all five criteria of the melts 3 and 4,
F. KAVICKA et al.: THE EFFECT OF ELECTROMAGNETIC STIRRING ON THE CRYSTALLIZATION ...
600 Materiali in tehnologije / Materials and technology 45 (2011) 6, 599–602
Figure1: Macrostructure of breakout
Slika 1: Makrostruktura zloma
considering their partial homogenization, is 1.078 × 105
in melt 4 and 6.498 × 104 in melt 3. The quotient of the
product for melts 3 and 4 is 0.603, which predicts a
reduced risk of breakout in melt 3. If the influence of
temperature on the surface of the slab in melt 3 and in
the place of the groove in melt 4, it is clear that the effect
of the groove during the straightening out of the slab is
connected with the tensile stress, then in the place of the
groove (Figure 1) the effect must have been compen-
sated for at a temperature of 1097 °C, i.e., at a tempe-
rature that is 163 °C higher than that of a completely
straight surface of the slab of melt 3. The data was
obtained from the investigation into the causes behind a
transversal crack that occurred in a different steel slab 4.
In order to clarify this, it was necessary to conduct a
series of ductility tests at temperatures ranging from
20 °C to the solidus temperature. Table 3 contains the
test results from temperatures that are close to the
temperatures in row 16 of Table 2. A comparison of the
mechanical values indicates that the tensile strength at
914.5 °C and the pulling force are 1.5 times greater than
at 1093.0 °C. In addition to this, there was a 8 605 m
column of melt working on the mushy zone at the point
of the breakout, where the mushy zone reached h S
max =
21.07 m from the level in the mould, i.e., at least 6.92 m
beyond the breakout point. It is therefore possible to
assume that the main factor that significantly increased
the risk of breakout was the superposition of the causing
F. KAVICKA et al.: THE EFFECT OF ELECTROMAGNETIC STIRRING ON THE CRYSTALLIZATION ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 599–602 601
Table 1: Parameters characterizing the concasting of melt 3 (quality A) and melt 4 (quality B)
Tabela 1: Parametri, ki karakterizirajo kontiulivanje taline 3 (kakovost A) in taline 4 (kakovost B)
Item # Parameter Symbol Units A – melt 3 B – melt 4
1 Pouring speed w m s–1 0.0130 0.0126
2 Dynamic viscosity  m–1 kg s–1 0.00570 TL 0.00562 TL
 = · m–1 kg s–1 0.00772 TS 0.00615 TS
3 Density  kg m–3 7560.7 7600.9
4 Latent heat of the phase change L m2 kg s–2 246 × 103 259 × 103
5 Specific heat capacity cp m2 s–2 K–1 632.6 611.0
6 Mould oscillation amplitude S m 0.006 ± 0.003 0.006 ± 0.003
7 Oscillation frequency f s–1 1.533 1.533
8 Solidus temperature TS °C 1427.0 1480.6
9 Liquidus temperature TL °C 1493.9 1512.3
10 Difference between the liquidus and solidus
temperatures TL – TS °C 66.9 31.7
11 Max. length of the isosolidus curve from the level* hS
max m 21.07 19.72
12 Min. length of the isosolidus curve from the level** hS
min m 19.92 18.69
13 Max. length of the isoliquidus curve from the level* hL
max m 14.50 16.20
14 Min. length of the isoliquidus curve from the level** hL
min m 13.70 15.20
15 The area of the mushy zone on half of the
cross-section of the breakout + Fmushy m
2 0.05366 0.04100
16 The surface temperature of the slab++ Tsurf °C 934 1097
Note (continued from table above): *) of the steel inside the mould to a position 0.650 m from the edges of the 1.53 m wide slab; **) of the steel inside the mould to
the centre of the slab; +) the overall area of half of the cross-section is Fslab = 0,19125 m
2 ; ++) in the material 15 mm around the groove (Figure 1). The data in Table 1
were established a) on the caster after breakout; b) from archived on-line results of the temperature model; c) by off-line modelling of the temperature field of melts 3
and 4 5.
Table 2: Dimensionless criteria characterizing the breakout
Tabela 2: Brezdimenzijska merila, ki ozna~ujejo zlom
Criterion
L f
c T S
⋅
p L Δ
ΔS f
w
⋅ 

ΔS f2 ⋅ F
F F
slab
slab mushy−
T T
T
L S
L
−
steel A 5124.78 1.179 172.77 1.3900 0.044782
steel B 6237.96 1.217 197.87 1.2729 0.056404*
Note: *) The maximum temperature difference inside the mixture zone (TL–B – TS–A)/TL–B
Table 3: Ductility testing at 1093.0 °C and 914.5 °C 5
Tabela 3: Preizkusi gnetljivosti pri 1093.0 °C in 914.5 °C 5
Sample
Testing
temperature
Tensile
strength Strength Diameter Contraction
Deformation
before breaking
Breaking
Work
°C N MPa mm % mm J
1 1093 817 28.9 3.90 58.0 12.0 7
2 914.5 1247 44.1 5.35 21.5 5.5 6
effects of the parameters occurring in the first four
criteria of Table 2.
5 DISCUSSION
Following a rapid change of the tundish, there was a
period of 20 min when there was a mixture of quality A
and quality B steels. The liquidus temperature 1493.9 °C
of quality A increased to 1512.3 °C and, simultaneously,
the latent heat of the phase change increased from
246 kJ/kg (quality A) to 259 kJ/kg (quality B). This led
to an increase in the temperature of the melt and to the
re-melting of the solidified shell of the original quality A
steel. Furthermore, there was an increase in the length of
the mushy zone (up to h S t melt−3 .
max – h S melt−3.
min = 21.07 –
13.70 = 7.37 m) and also in its temperature hetero-
geneity. The temperature of the mushy zone – following
the mixing of both qualities – could find itself anywhere
between the maximum temperature of the liquidus of
quality A and the minimum temperature of the solidus of
quality B (i.e., within the interval TL – B – TS – A= 1512.3 –
1427.0 = 85.3 °C. During the 20 min of pouring of the
quality B steel (the 4th melt), which began immediately
after the quality A steel (the 3rd melt), marks and hooks
formed as a result of the oscillation of the mould and
continued to form during the unbending of the slab
(Figure 1 – where the groove is 50 mm wide and
15–16 mm deep with an opening angle of 115°). The
tensile forces in the vicinity of this groove and the
re-melting of the solidified shell brought about the
breakout in the wall of the small radius of the slab in the
unbending point.
6 CONCLUSION
The changes in the chemical composition of the steel
during the actual concasting are particularly dangerous.
One way of reducing the risk of breakout and the
consequent shutdown of the caster is to modify the
values of the dimensionless criteria characterizing the
breakout, i.e., to select two consecutive melts of such
chemical compositions and the corresponding physical
and chemical parameters (from which the dimensionless
criteria are determined) that the criteria predict zero-
breakthrough.
Acknowledgements
GACR projects No. 106/08/0606, 106/09/0940, 106/
09/0969 and P107/11/1566.
7 REFERENCES
1 Badri A. et al.: A mold simulator for continuous casting of steel. part
ii. the formation of oscillation marks during the continuous casting of
low carbon steel. Metallurgical and Materials Transactions B, 36
(2005), 373–383
2 Thomas, B. G., J. Sengupta, Ojeda, C. Mechanism of hook and
oscillation mark formation in ultra-low carbon steel. In Second
Baosteel Biennial Conference, (May 25–26, 2006, Shanghai, PRC), 1
(2006), 112–117
3 Ojeda, C. et al.: Mathematical modeling of thermal-fluid flow in the
meniscus region during an oscillation cycle. AISTech Proceedings, 1
(2006), 1017–1028
4 Dobrovska, J. et al.: Analysis of a transversal crack in a steel slab.
Materials Science Forum, 567–568 (2007), 105–108 ©2008 Trans
Tech Publications, Switzerland
5 Stetina, J. et al.: Optimization of a casting technology of a steel slab
via numerical models. Proceedings 22nd Canadian Congress of
Applied Mechanics, Dalhousie University Halifax, Nova Scotia,
Canada, May 2009, 4
F. KAVICKA et al.: THE EFFECT OF ELECTROMAGNETIC STIRRING ON THE CRYSTALLIZATION ...
602 Materiali in tehnologije / Materials and technology 45 (2011) 6, 599–602
J. BA@AN et al.: WEAR OF REFRACTORY MATERIALS FOR CERAMIC FILTERS OF DIFFERENT POROSITY ...
WEAR OF REFRACTORY MATERIALS FOR CERAMIC
FILTERS OF DIFFERENT POROSITY IN CONTACT WITH
HOT METAL
OBRABA OGNJEVZDR@NEGA MATERIALA KERAMI^NIH
FILTROV Z RAZLI^NO POROZNOSTJO V STIKU Z VRO^O
KOVINO
Jiøí Ba`an1, Ladislav Socha1, Ludvík Martínek2, Pavel Fila2, Martin Balcar2,
Jaroslav Chmelaø3
1V[B - Technical University of Ostrava, FMME, Department of Metallurgy, 17. listopadu 15/2172, 708 33 Ostrava-Poruba, Czech Republic
2@ÏAS, a. s., Strojírenská 6, 591 71 @ïár nad Sázavou, Czech Republic
3KERAMTECH, s.r.o., Horská 139, 542 01 @acléø, Czech Republic
jiri.bazan@vsb.cz
Prejem rokopisa – received: 2011-05-31; sprejem za objavo – accepted for publication: 2011-09-02
This paper deals with an investigation of the development of refractory materials for the fabrication of ceramic filters for the
filtration of steel. Ceramic filters are used for increasing the cleanliness of steel and they must meet several strict requirements,
such as the ability to remove impurities, a resistance to sudden changes in temperature, a resistance to corrosion and erosion by
metal. The use of filters must not lead to an excessive reduction of the steel’s temperature, as this may lead to solidification of
steel and thus to filter clogging. That is why a special refractory material has been developed with reduced thermal capacity
caused by increased porosity. Tests were made in a laboratory of the Department of Metallurgy at V[B-TUO in order to
simulate the industrial conditions of the filtration of steel with a focus on the evaluation of the erosion and corrosion effects and
also on a determination of the resistance and service life of ceramic filters.
Keywords: steel, ceramic filter, refractory material, corrosion and erosion, porosity
V ~lanku je opisano raziskovanje v zvezi z razvojem ognjevzdr`nih materialov za kerami~ne filtre za filtriranje `eleza. Te filtre
se uporablja za izbolj{anje ~istosti jekla, zato morajo izpolnjevati ve~ strogih zahtev, kot npr.: odstranitev ne~isto~, odpornost
proti sunkovitim spremembam temperature in odpornost proti koroziji in eroziji zaradi kovine. Uporaba filtra ne sme preve~
zni`ati temperature jekla in povzro~iti njegovega strjevanja ter zato tudi ma{enja filtra. Zato je bilo razvit poseben ognjevzdr`en
material, ki ima zmanj{ano termi~no kapaciteto zaradi pove~ane poroznosti. Preizkusi so bili opravljeni v laboratoriju Oddelka
za metalurgijo V[B-TUO, da bi simulirali industrijske razmere filtriranja jekla s poudarkom na oceni korozijskih in erozijskih
u~inkov in za dolo~itev odpornosti in dobe uporabe kerami~nih filtrov.
Klju~ne besede: jeklo, keramika, ognjevzdr`ni material, korozija, erozija, poroznost
1 INTRODUCTION
The technology of filtration, i.e., the use of ceramic
filters in a gating system, is one of the possibilities for
enhancing the cleanliness of steel and the quality of
as-cast ingots. This makes it possible to achieve an
increase of the steel’s purity, a reduction of the
occurrence of non-metallic inclusions, a reduction of the
costs of repairs of defects, etc.
Ceramic filters are exposed to extreme working
conditions, such as sudden changes of temperature, the
resistance to corrosion and erosion caused by hot
metal or molten slags, etc. Ceramic filters are currently
commonly used for increasing the cleanliness of steel in
steel shops, where filtration is used for the removal of
non-metallic inclusions, particularly those of an
exogenous character and of the rest of slide valve nozzle
fill, e.g., during uphill casting, but also in the tundish
during the continuous casting of steel. Nowadays, a
whole series of structurally different types of ceramic
filters for the filtration of metals are manufactured. The
most commonly used types are strainer and foam filters 1.
The paper concentrates on the influence of the porosity
of refractory materials on their density and thus on the
reduction of their thermal capacity.
2 USE OF CERAMIC FILTERS DURING THE
CASTING OF STEEL INGOTS
The technology of the filtration of steel was tested in
industrial conditions in a steel shop at the company
@ÏAS, a. s., during the uphill casting of ingots through a
gating system in order to eliminate the occurrence of
inclusions and to ensure an improved purity of the steel.
The system for the casting of ingots consisted of a gate
stick, a stool and an ingot mould with a shrink head
(Figure 1). The application of the filtration system for
the casting of ingots consisted of the use of a series of
filters situated in a ceramic cartridge arranged in
succession. The cartridge is placed in the gating system
in the broadened channel of the stool (Figure 2).
The steel shop of @ÏAS, a. s., participated in the
design and realisation of the technical solution, including
Materiali in tehnologije / Materials and technology 45 (2011) 6, 603–608 603
UDK 669.1:620.193:620.1/.2 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 45(6)603(2011)
the manufacture of cartridges and filters. Ceramic foam
filters of 150 mm × 150 mm × 30 mm made of material
based on ZrO2, SiO2 + SiC, as well as of material based
on carbon, Al2O3 and SiO2 were then tested in this steel
shop. Due to problems during pouring (mechanical
damage and freezing of the metal) in the next stage the
filtration cartridges were modified and ceramic strainer
filters based on mullite with dimensions of 100 mm ×
100 mm × 20 mm and 133 mm × 133 mm × 20 mm were
used. The filtration cartridge was made of fireclay
material with the share of the mass fraction of Al2O3 >
61 % (Figure 3).
It was ascertained during industrial applications that
this arrangement is satisfactory; however, in this case too
during the flow of steel through the filter channels, the
steel was cooled and problems with its freezing occurred
as well. Subsequently, in order to minimise this problem,
the company KERAMTECH, s. r. o., developed and
tested a refractory material, in which its porosity was
purposefully increased (up to 10 % of material), which
led to a reduction of its mass and cooling effect.
3 DEVELOPMENT AND TESTING OF
REFRACTORY MATERIALS FOR NEW
CERAMIC FILTERS
The refractory material was developed by the com-
pany KERAMTECH, s. r. o., and a refractory material
consisting of a mullite-corundum mass with higher
contents of Al2O3 was chosen for the modification. Table
1 gives the chemical composition of this material. The
porosity of this material was increased by the addition of
organic mass in portions of (3, 5, 7.5 and 10) % in order
to reduce the material’s thermal capacity.
Table 1: Chemical composition of modified refractory material
Tabela 1: Kemi~na sestava modificiranega ognjevzdr`nega materiala
v masnih dele`ih, w/%
Chemical composition in mass fractions (w/%)
SiO2 Al2O3 Fe2O3 TiO2 K2O CaO MgO Na2O
21.0 75.0 0.8 0.6 0.8 less than 0.5
The specific thermal capacity of ordinary material is
1800 kJ K–1 kg–1. The addition of 1 % of organic mass to
the basic material reduces, by increased porosity, the
mass of the final product by 2 %, and thus also its
thermal capacity.
The modified refractory materials with increased
porosity were tested with experimental heats in a
laboratory of V[B – Technical University of Ostrava to
verify the erosion and corrosion effects and to determine
the resistance and service life of the new ceramic filters.
These experimental heats were supposed to simulate the
conditions during the industrial filtration of hot metal.
Two types of steel were used for all the experimental
heats, i.e., ordinary carbon steel and high-manganese
steel. Table 2 gives the chemical compositions of both
these steels, including the liquidus temperature.
Table 2: Chemical compositions of the used steels with liquidus tem-
peratures
Tabela 2: Kemi~na sestava uporabljenih jekel in likvidusna tempe-
ratura
Type of steel
Chemical composition (w/%)
Tl/°CC Si Mn P S
Carbon steel 0.44 0.23 0.67 0.019 0.016 1495
Manganese
steel 1.1-1.5
max.
0.70
12.0-
14.0
max.
0.10
max.
0.050 1375
J. BA@AN et al.: WEAR OF REFRACTORY MATERIALS FOR CERAMIC FILTERS OF DIFFERENT POROSITY ...
604 Materiali in tehnologije / Materials and technology 45 (2011) 6, 603–608
Figure 3: Filtration cartridge
Slika 3: Filtrirni tulec
Figure 2: Cross-section of casting system with the application of a
filtration cartridge
Slika 2: Prerez sistema za litje ingotov s filtrirnim tulcem
Figure 1: Cross-section of system for casting of ingots
Slika 1: Prerez sistema za litje ingotov
Samples with dimensions 10 mm × 10 mm × 100 mm
were made from the modified refractory material. For the
simulation of the usual industrial conditions, prior to
their insertion into the hot metal the samples were
pre-heated to a temperature of 350 °C for 10 min. This
tempering of the samples simulates the heating of the
casting system in industrial conditions. Before their
insertion into the reheating furnace and the start of the
experiment, the samples were weighed in order to
determine the mass loss. A comparison of the results
showed that the re-heating of the samples did not cause
any loss of mass. Afterwards, experimental heats were
carried out. The induction furnace served as a melting
unit. For the evaluation of corrosion and erosion
phenomena, the experiments were made afterwards with
use of both carbon and manganese steel at temperatures
of 1560 °C, 1600 °C and 1680 °C for 20 min, while the
porosity of the tested refractory materials was increased
by the addition of various organic masses 10 wt.%.
4 EVALUATION OF REFRACTORY MATERIALS
The evaluation of the exposed samples was made in
several steps. It started with a visual evaluation (photo-
graphs of the whole samples taken after the experiment)
followed by an evaluation of cross-sections with a focus
on the structure and the surface (edge) of the samples.
4.1 Visual evaluation of erosion and corrosion
With the experiments the refractory materials were
tested from the point of view of the influence on carbon
and manganese steels at the extreme temperature of 1680
°C for 20 min with contents of organic mass in the
refractory material in volumes of (3, 5, 7.5 and 10) %.
Figure 4 shows the results of the first series of experi-
ments.
It is evident from this figure that the addition of
organic mass in the quantity up to approx. 3 % had no
significant influence on the wear of the refractory
materials. However, larger additions of organic mass up
to 10 % brought about a distinct wear and deformation of
refractory material in both the carbon and manganese
steels. It is also evident that in the case of use of
manganese steel, the corrosion effects of refractory
materials were substantially higher than for the heats of
the carbon steel.
J. BA@AN et al.: WEAR OF REFRACTORY MATERIALS FOR CERAMIC FILTERS OF DIFFERENT POROSITY ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 603–608 605
Figure 5: Pictures of refractory material after exposition in carbon steel at temperatures of 1560 °C and 1600 °C for a period of 20 min
Slika 5: Posnetki ognjevzdr`nega materiala po 20-minutni izpostavi ogljikovemu jeklu pri temperaturah 1560 °C in 1600 °C
Figure 4: Pictures of refractory material after exposition in carbon and manganese steel at a temperature of 1680 °C for a period of 20 min
Slika 4: Posnetki ognjevzdr`nega materiala po 20-minutni izpostavi ogljikovemu in manganovemu jeklu pri temperaturi 1680 °C
J. BA@AN et al.: WEAR OF REFRACTORY MATERIALS FOR CERAMIC FILTERS OF DIFFERENT POROSITY ...
606 Materiali in tehnologije / Materials and technology 45 (2011) 6, 603–608
Figure 8: Comparison of pictures of cross-sections stressed at the temperature of 1600 °C in manganese steel taken with stereo-microscope and
scanning electron microscope
Slika 8: Primerjava posnetkov prereza, obremenjenega pri temperaturi 1600 °C v manganovem jeklu, pripravljenih s stereomikroskopom in z
vrsti~nim elektronskim mikroskopom
Figure 7: Comparison of pictures of cross-sections stressed at the temperature of 1600 °C in carbon steel taken with stereo-microscope and
scanning electron microscope
Slika 7: Primerjava posnetkov prereza, obremenjenega pri temperaturi 1600 °C v ogljikovem jeklu, pripravljenih s stereomikroskopom in z
vrsti~nim elektronskim mikroskopom
Figure 6: Pictures of refractory material after exposition in manganese steel at temperatures of 1560 °C and 1600 °C for a period of 20 min
Slika 6: Posnetki ognjevzdr`nega materiala po 20-minutni izpostavi manganovemu jeklu pri temperaturah 1560 °C in 1600 °C
On the basis of previous results the experiments were
carried out under modified conditions, again with
refractory materials with contents of organic mass of (3,
5, 7.5 and 10) % with use of carbon and manganese steel
with an interaction time of 20 min, but at temperatures of
1560 °C and 1600 °C. The objective of these tests was to
simulate the temperatures in practical conditions of the
uphill casting of steels into ingot moulds, as well as to
test the influence of various additions of organic mass on
the wear at these reduced temperatures. The results of
the experiments are shown in Figures 5 and 6.
An analysis of these figures revealed that a tempe-
rature of 1560 °C seems to be too low for carbon steels.
The liquidus temperature calculated on the basis of the
chemical composition of the steel is 1495 °C (see
Table 2). In this case freezing of the steel on the walls of
the ceramic samples occurred during testing. In the case
of an industrial application this would require an
increase in the pouring temperature from the usual 1560
°C (pouring temperature used at @ÏAS, a. s.) to approxi-
mately 1570–1575 °C (i.e., by about 10–15 °C).
However, with the same steel and temperature of
1600 °C, the refractory materials showed, with contents
of 5 % of organic mass, only minimum wear, and
slightly higher wear for a content of 7.5 %. Nevertheless,
higher contents of organic mass up to 10 % already had a
negative impact at this temperature. During the use of
manganese steel at the temperature of 1560 °C and at
calculated liquidus temperature of 1375 °C (see Table 2)
the degree of wear was higher for contents of organic
mass higher than 7.5 %. For this reason, experiments at a
temperature of 1600 °C were made without the sample
containing 10 % of organic mass. At the temperature of
1600 °C a minimum loss was determined in the same
(manganese) steel for the contents of 3 % of organic
mass in the refractory material 2.
4.2 Evaluation of cross-sections
Apart from a visual evaluation of the samples after
the experiments, their cross-sections were evaluated as
well. The evaluation itself was made on the basis of a
visual comparison of images taken using an Olympus
stereo-microscope and Tescan Vega scanning microscope
operating in the "fish eye" mode. The photos taken with
the stereo-microscope make it possible to determine the
depth of penetration, the material structure and also the
losses of material. The pictures taken with the scanning
microscope enable a determination of the cracks,
fissures, structure failures, and in some cases, also the
depth of the penetration. For an illustration, only the
samples with use of carbon and manganese steels for 20
min at the temperature of 1600 °C were used, with the
contents of organic mass in the refractory materials of (3,
5, 7.5 and 10) %.
The photographs of the cross-sections taken with the
stereomicroscope and the scanning microscope of the
carbon steel at the temperature of 1600 °C are shown in
Figure 7. The images show that the character and
morphology of the surfaces of the refractory materials
are similar. Refractory materials with contents of organic
mass up to 5 % showed minimum wear, and a slightly
higher wear was observed for the contents from 7.5 %.
Figure 8 shows pictures of the cross-sections taken with
the stereo-microscope and the scanning microscope of
the manganese steel at the temperature of 1600 °C. The
pictures show that for this steel the addition of organic
mass >5 % at the temperature of 1600 °C had a negative
impact, and it influenced not only the surface layers of
the refractory material, but also its central parts. Larger
additions had a very negative impact on the material’s
structure.
5 CONCLUSIONS
With the development of new ceramic filters intended
for the filtration of steel in the gating system during the
casting of ingots, experiments were carried out in
laboratory conditions with a focus on the verification of
the influence of porosity in refractory materials on the
erosion and corrosion caused by the steel. Afterwards,
various pictures were visually evaluated. This evaluation
made it possible to assess the influence of the additions
of the organic mass on the degree of wear of the
refractory material under various operating conditions.
The following conclusions may be drawn on the basis of
the results of the laboratory experiments:
– it is obvious from the results obtained in both series
that the manganese steel had a much greater
corrosive impact on the tested refractory materials
than the carbon steel,
– for the tested samples in the first part during the
contact with carbon and manganese steel at the
temperature of 1680 °C for 20 min, the addition of
organic mass up to approx. 3 % had no significant
influence on the wear of the refractory materials.
However, larger additions of organic mass up to
10 % caused a distinct wear and deformation of the
refractory material,
– the samples of refractory materials in the second part
were in contact with the carbon and manganese steel
for 20 min, but at temperatures of 1560 °C and 1600
°C. The temperature of the experiments at 1560 °C
was too low for the carbon steel, since freezing of
the steel on the walls of the samples occurred during
testing. In the case of the same steel and a tempe-
rature of 1600 °C the samples showed up to the
contents of 5 % of organic mass only a minimal
amount of wear. However, larger contents of organic
mass of 10 % had a negative influence at the
temperature of 1600 °C,
– for the manganese steel and a temperature of 1560
°C an increased degree of wear was found for con-
tents of organic mass exceeding 7.5 %. For this
reason, the experiments at the temperature of 1600
J. BA@AN et al.: WEAR OF REFRACTORY MATERIALS FOR CERAMIC FILTERS OF DIFFERENT POROSITY ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 603–608 607
°C were realised without the sample containing 10 %
of organic mass. The samples showed for the same
steel a minimum loss of refractory material at the
temperature of 1600 °C and contents of 3 % of
organic mass,
– it was determined from the pictures of the cross-
sections for the carbon steel and the temperature of
1600 °C, that up to the contents of 5 % of organic
mass the refractory materials showed only minimal
wear, while at the contents of 7.5 % and more of
organic mass they showed slightly increased wear.
However, the manganese steel had a negative impact
during the addition of organic mass >5 %, and this
phenomenon influenced not only the surface layers,
but also the central areas of the refractory material,
– on the basis of the obtained results the company
KERAMTECH, s. r. o., produces ceramic filters with
the addition of 5 % of organic mass under the
designation RK-5/5.
Acknowledgement
This work was carried out within the project
EUREKA MICROST OE08009 and the project
FR-TI1/222.
6 REFERENCES
1 Stránský, K., Ba`an, J., Horáková, D. Filtrace taveni `eleza
v prùmyslové praxi [Filtration of molten steel in practice]. Editor.
Ostrava, V[B-TU Ostrava, 2008, 113 pp. ISBN 978-80-248-1844-3
2 Ba`an, J., Socha, L., Martínek, L., Fila, P., Balcar, M., Lev, P. Effect
of Refractory Materials Porosity on Erosion and Corrosion of
Ceramic Filters by Molten Steel Treatment. Hutnické listy, 63 (2010)
2, 20–2
J. BA@AN et al.: WEAR OF REFRACTORY MATERIALS FOR CERAMIC FILTERS OF DIFFERENT POROSITY ...
608 Materiali in tehnologije / Materials and technology 45 (2011) 6, 603–608
M. KRGOVI] et al.: THE INFLUENCE OF THE MINERAL CONTENT OF CLAY FROM THE WHITE BAUXITE MINE ...
THE INFLUENCE OF THE MINERAL CONTENT OF
CLAY FROM THE WHITE BAUXITE MINE ON THE
PROPERTIES OF THE SINTERED PRODUCT
VPLIV VSEBNOSTI MINERALA GLINE IZ RUDNIKA BELEGA
BOKSITA NA LASTNOSTI SINTRANEGA PROIZVODA
Milun Krgovi}1, Ivana Bo{kovi}1, Mira Vuk~evi}1, Radomir Zejak2, Milo{
Kne`evi}2, Ratko Mitrovi}2, Biljana Zlati~anin1, Nata{a Ja}imovi}1
1University of Montenegro, Faculty of Metallurgy and Technology, D`. Va{ingtona bb, 81000 Podgorica, Montenegro
2University of Montenegro, Faculty of Civil Engineering, D`ord`a Va{ingtona bb, 20000 Podgorica, Montenegro
milun@ac.me
Prejem rokopisa – received: 2011-05-30; sprejem za objavo – accepted for publication: 2011-08-24
An investigation of the influence of the mineral content of clay from the White Bauxite Mine on the properties of the sintered
product is presented. The whole area of the white bauxite deposit (the "Bijele Poljane" mine) is characterized by the presence of
clays. To investigate the properties of the sintered product, two of the most present types of clays were used (marked as "B" and
"C"). The investigations of the properties of the sintered products on the basis of these clays involved the linear and volume
shrinkage, the total porosity and the compression strength. The sintering process was conducted at temperatures of 1000 °C,
1100 °C, 1200 °C, 1300 °C and 1400 °C.
Key words: clay, linear shrinkage, total shrinkage, compression strength, sintering, porosity
Raziskava vpliva vsebnosti minerala gline iz Rudnika belega boksita je predstavljena v tem delu. Karakteristi~na za celotno
podro~je le`i{~a belega boksita (rudnik Bele Poljane) je prisotnost gline. Lastnosti sintranega proizvoda so bile izvr{ene pri
dveh najve~ prisotnih vrstah gline (ozna~bi B in C). Dolo~ene so bile naslednje lastnosti sintranih proizvodov teh glin: linearno
in volumensko kr~enje, celotna poroznost in tla~na trdnost. Sintranje je bilo izvr{eno pri temperaturah (1000, 1100, 1200, 1300
in 1400) °C.
Klju~ne besede: glina, linearno kr~enje, skupno kr~enje, tla~na trdnost, sintranje, poroznost
1 INTRODUCTION
The investigated clay types from the White Bauxite
Mine appear in layers and have different mineral con-
tents:
– clays with bauxite minerals,
– illite-kaolinite clays,
– clays with a very heterogeneous mineral content.
This mineral content gives us the possibility to obtain
different ceramic products 1. The most important
differences in the mineral content are related mostly to
the content of bauxite minerals, clay minerals, iron
compounds, quartz and calcite 2. Depending on the
sintering temperature, this mineral content of the clays
influences the solid-state reactions, the polymorphic
transformations of the quartz and the liquid-phase
formation3. Apart from the degree of the sintering of the
ceramic mass, i.e., the firing regime, the mineral content
of the raw material also has an important influence on
the relations between particular microstructural ele-
ments. 4 The new crystal phases, i.e., the compounds
formed within the solid-state reactions during sintering,
are determined by the mineral content of the clays, as
well as by the previously mentioned factors. 5,6 It causes
important differences in the mineral content of the
sintered products. The content of bauxite minerals
(boehmite, gibbsite), iron compounds (hematite), clay
minerals (kaolinite, illite) in the investigated clay types
primarily determines the properties of the sintered
products.7,8 According to their refractoriness, the
investigated clay types are refractory (the refractoriness
is over 1500 °C).1,9 Neither flux, as a component that
decreases shrinkage, nor electrolytes were used for the
raw-material mixture formation, because the aim of the
investigation was to determine the influence of the
mineral content of the investigated clays on the
properties of the sintered product.
2 EXPERIMENTAL
The samples were formed by shaping of a plastic
mass in a mould corresponding to a parallelepiped with
dimensions of 7.7 cm × 3.9 cm × 1.6 cm. The
characterization of the investigated clays was made by a
determination of the mineral content, the chemical
content, the density, the humidity and the granulometry.
The chemical content was determined with a Perkin
Elmer 4000 atomic absorption spectrophotometer. The
granulometry of the clays was determined on a "Micro-
sizer 201C" VA INSTALT instrument. For the raw,
non-sintered products the linear and volume shrinkage
during the drying in air and in a dryer, to a constant
Materiali in tehnologije / Materials and technology 45 (2011) 6, 609–612 609
UDK 622.349.1:666.32/.36 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 45(6)609(2011)
mass, at a temperature of 110 °C was determined. The
samples were sintered at (1000, 1100, 1200, 1300 and
1400) °C.
For the sintered products the linear and volume
shrinkage during sintering, the total porosity and the
compression strength were determined. The volume
shrinkage was determined using the following equation:
C
V V
V
i
v =
−
⋅0
0
100(%)
V0/m3 – the volume before shrinkage
Vi/m3 – the volume after shrinkage
The total porosity was determined using the
following equation:
TP =
−
⋅
 

0
0
100
v
(%)
/(kg m–3) – the sample’s density
v/(kg m–3) – the sample’s volume mass
The compression strength was determined on a
novopress HPM 400 and the X-ray analysis was
performed on a Difractometer PHILIPS PW 1710. The
microscopic analysis of the sintered products was made
by scanning electron microscopy on a JEOL-JSM
6460LV-DEU (BAL-TEC SCD 005-SPVTTER
COATER).
3 RESULTS AND DISCUSSION
The results of the chemical analysis (Table 1) show a
slightly higher content of oxides, i.e., CaO and MgO, in
the clay "C" as a consequence of the higher content of
carbonates. The clay "C" contains more Fe2O3 in the
mass fraction, w (w = 2.3 %), with respect to the clay
"B" (w = 2.2 %). The content of TiO2 in clay "B" is
slightly higher (w = 0.8 %) compared to the clay "C" (w
= 0.7 %).
The mineral content of the clay "B", determined by
the X-ray analysis (Figure 1), shows the presence of
kaolinite, gibbsite, boehmite and anatase. The X-ray
analysis shows the presence of the following minerals in
the clay "C" (Figure 2): kaolinite, gibbsite, boehmite,
anatase, hematite, illite and clinochlore.
The results of the granulometric analysis show an
average grain size of 69.32 μm in the samples on the
basis of clay "B", while in the samples on the basis of
clay "C" the average grain size is 63.25 μm, under the
same milling conditions. The clay "B" has a higher
average grain size with respect to the clay "C" under the
same milling conditions, which is a consequence of the
mineral content of the clays.
The volume shrinkage of the products during
sintering (Figure 3) grows with the increase of the
temperature as a consequence of solid-state reactions,
polymorphic transformations of the quartz, the carbo-
nates’ dissolution, the glass-phase formation and the
closure of pores during sintering. The content of K2O in
clay "C" is slightly higher (1.0 %) with respect to clay
"B" (0.9 %) and it can be concluded that it does not have
an important influence on the differences in the
liquid-phase content, which accelerates the solid-state
reactions (the diffusion coefficient increases).
The granulometric content of clays "B" and "C" also
has an important influence on the volume shrinkage. The
average grain size is greater in clay "B".
The content of Fe2O3 also has an influence on the
volume shrinkage during sintering, and the content of
Fe2O3 in the clay „C" (2.3 %) is slightly higher compared
to the clay "B" (2.2 %).
M. KRGOVI] et al.: THE INFLUENCE OF THE MINERAL CONTENT OF CLAY FROM THE WHITE BAUXITE MINE ...
610 Materiali in tehnologije / Materials and technology 45 (2011) 6, 609–612
Table 1: Chemical composition of clay "B" and "C" in mass fractions, w/%
Tabela 1: Kemi~na sestava glin "A" in "B" v masnih dele`ih, w/%
Oxides SiO2 Fe2O3 Al2O3 CaO MgO K2O TiO2 lg. loss
Clay "B" 63,2 2,2 25,2 1,2 1,3 0,9 0,8 4,1
Clay "C" 64,2 2,3 24,8 1,3 1,4 1,0 0,7 3,8
Figure 2: X-ray diffractogram of "C" clay
Slika 2: Rentgenski difraktogram gline "C"
Figure 1: X-ray diffractogram of "B" clay
Slika 1: Rentgenski difraktogram gline "B"
The volume shrinkage increases with the increase of
the sintering temperature and the growth is more evident
in samples on the basis of clay "B" at higher tempera-
tures (1300 °C and 1400 °C), which can be explained by
the differences in the mineral and granulometric contents
of the clays and solid-state reactions (X-ray analysis of
sintered product at T = 1300 °C and 1400 °C, see Figure
6 and Figure 7).
The total porosity during sintering decreases with the
increase of the temperature (Figure 4). Apart from the
intense solid-state reactions at higher temperatures
(liquid-phase formation, diffusion coefficient increase),
the decisive factors for the decrease of the porosity at
higher temperatures are the chemical, mineral and granu-
lometric contents of the raw material. The granulometric
analysis shows that clay "B" has a larger average grain
size with respect to the clay "C".
Considering the small difference in the content of
alkali, they do not decisively influence the total porosity.
The total porosity at 1400 °C has smaller values in the
samples on the basis of clay "C", which is probably a
consequence of the grain fraction (smaller average grain
size) and the relations between the microstructural
elements.
Clay "B" has a higher average grain size with respect
to clay "C", which increases the total porosity and in this
way the granulometric content of the clays influences
indirectly on the compression strength. For the samples
on the basis of clay "B" at (1100, 1200, 1300 and 1400)
°C it is evident that there is an almost linear dependence
of the compression strength of the sintering temperature,
while in samples on the basis of clay "C" the
compression strength value at temperatures of (1100,
1200, 1300 and 1400) °C (Figure 5) increases only
slightly, which can be explained by the mineral content .
In the samples on the basis of clay "B" with an increase
M. KRGOVI] et al.: THE INFLUENCE OF THE MINERAL CONTENT OF CLAY FROM THE WHITE BAUXITE MINE ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 609–612 611
Figure 7: X-ray diffractogram of sintered product (clay "C", T = 1300
°C)
Slika 7: Rentgenski difraktogram sintranca (gline "C", T = 1300 °C)
Figure 4: Total porosity of products during sintering (clay "B" and
clay "C")
Slika 4: Skupna poroznost med sintranjem (glina "B" in "C")
Figure 5: Compression strength of product during sintering (clay "B"
and clay "C")
Slika 5: Tla~na trdnost sintranca med sintranjem (glina "B" in "C")
Figure 3: Volume shrinkage of product during sintering (clay "B" and
clay "C")
Slika 3: Volumensko kr~enje med sintranjem (glini "B" in "C")
Figure 6: X-ray diffractogram of sintered product (clay "B", T = 1300
°C)
Slika 6: Rengenski difraktogram sintranca (gline "B", T = 1300 °C)
of the temperature up to 1400 °C a high compression
strength can be obtained (above 90 MPa). With an
increase of the sintering temperature in samples on the
basis of clay "C" above 1200 °C the compression
strength increases slightly.
The X-ray structure analyses of the sintered products
in the samples on the basis of clay "B" and "C" show
important differences in the mineral content (Figure 6
and Figure 7). The crystal phases at the temperature of
1300 °C for samples on the basis of clay "B" are mullite,
hematite and ilmenit, while for samples on the basis of
clay "C" they are mullite, hematite and anatase. Mullite
and hematite, as crystal phases, are present in sintered
products even at temperatures of (1100, 1200 and 1400)
°C. The presence of the crystal phases is a consequence
of the solid-state reactions at these temperatures. The
microstructural analysis of the sintered products (Figure
8 and Figure 9, 10,000 × magnification) shows the
complexity of the structure (glass phase, crystal phase,
unreacted grains and pores).
4 CONCLUSIONS
On the basis of the performed investigations of the
influence of the mineral content of clays from the White
Bauxite Mine it is concluded that:
– The investigated clay types are very different for in
terms of their mineral and chemical compositions,
– the volume shrinkage is greater for samples on the
basis of clay "B" and this is explained by the
difference in mineral and granulometric content of
the clays and the solid-state reactions,
– the differences in the total porosity of the sintered
products on the basis of the investigated clays are
not significant,
– the values of the compression strengths of the
sintered products are very different: in the samples
on basis of clay "B", important values of the
compression strength are noted only at temperatures
above 1300 °C, while in samples on the basis of clay
"C" this is already above 1100 °C.
On the basis of the volume shrinkage, the total
porosity and the compression strength it is concluded
that a quality ceramic product can be obtained by
sintering clays from the White Bauxite Mine.
5 REFERENCES
1 Lj. Kosti}-Gvozdenovi}, R. Ninkovi}, Inorganic Technology, Faculty
of Technology and Metallurgy, University of Belgrade, Belgrade
(1997)
2 N. Ja}imovi}, The Influence of the Mineral Composition of clay
from the Mine "Bijele Poljane" on the Characteristics of Sintered
Products, Master Thesis,University of Montenegro, Podgorica
(2010), 4–15
3 M. Tecilazi} - Stevanovi}, Principles of Ceramic Technology, Faculty
of Technology and Metallurgy, University of Belgrade, Belgrade
(1990)
4 M. Krgovi}, M. Kne`evi}, M. Ivanovi}, I. Bo{kovi}, M. Vuk~evi}, R.
Zejak, B. Zlati~anin, S. \urkovi}, The properties of a sintered
product based on electrofilter ash, Mater. Tehnol., 43 (2009) 6,
327–331
5 B. @ivanovi}, R. Vasi}, O. Janji}, Ceramic Tiles, Monography,
Institute of Materials in Serbia, Belgrade (1985), 12–17
6 M. Krgovi}, N. Z. Blagojevi}, @. Ja}imovi}, R. Zejak, Possibilities of
using red mud as raw materials mixture component for production
of bricks, Research Journal of Chemistry and Environment, 8
(2004), 73–76
7 M. Krgovi}, N. Marstijepovi}, M. Ivanovi}, R. Zejak, M. Kne`evi},
S. \urkovi}, The influence of illite-kaolinite clays mineral content
on the products shrinkage during drying and firing, Mater. Tehnol.,
41 (2007) 4, 189–192
8 S. \urkovi}, The Possibilities of using Electrofilter Ash TE
"Pljevlja" as Raw Materials component Mixture for producting
sintered product, Master Thesis, University of Montenegro, Podgo-
rica (2008), 4–5
9 J. Griffiths, Minerals in Foundry Casting, Ind. Min., 272 (1990),
35–40
M. KRGOVI] et al.: THE INFLUENCE OF THE MINERAL CONTENT OF CLAY FROM THE WHITE BAUXITE MINE ...
612 Materiali in tehnologije / Materials and technology 45 (2011) 6, 609–612
Figure 9: Microstructure of sintered product (clay "C", T = 1300 °C,
magn. 10 000-times)
Slika 9: Mikrostruktura sintranca (gline "C", T = 1300 °C, pov.
10 000-kratna)
Figure 8: Microstructure of sintered product (clay "B", T = 1300 °C,
magn. 10 000-times)
Slika 8: Mikrostruktura sintranca (gline "B", T = 1300 °C, pov.
10 000-kratna)
S. TOROS et al.: EFFECT OF PRE-STRAINING ON THE SPRINGBACK BEHAVIOR OF THE AA5754-0 ALLOY
EFFECT OF PRE-STRAINING ON THE SPRINGBACK
BEHAVIOR OF THE AA5754-0 ALLOY
VPLIV PRENAPENJANJA NA POVRATNO ELASTI^NO
IZRAVNAVO ZLITINE AA5754-0
Serkan Toros, Mahmut Alkan, Remzi Ecmel Ece, Fahrettin Ozturk*
Department of Mechanical Engineering, Nigde University, Nigde, 51245, Turkey
fahrettin@nigde.edu.tr
Prejem rokopisa – received:2011-03-31 ; sprejem za objavo – accepted for publication: 2011-09-20
This study presents the effect of pre-straining on the springback behavior of the AA5754-0 aluminum-magnesium (Al-Mg) alloy
sheet under V bending by an experimental and finite-element simulation studies. Pre-straining ranges from 0 % to 11 % were
applied to the samples, which were bent on a 60° V-shaped die for the springback evaluation. Commercially available
finite-element software, ETA/Dynaform, was used to simulate the 60° V-die bending process. The dynamic explicit
finite-element method for pressing and the static implicit finite-element method for the unloading phase were used for the
simulations. The results from both the experiment and the simulation indicate that the pre-straining has no positive effect on the
springback compensation.
Keywords: pre-straining; springback; Al-Mg alloy; AA5754-O alloy
Delo obravnava vpliv prenapenjanja na povratno elasti~no izravnavo plo~evine iz aluminij magnezijeve zlitine AA5754-0 pri
V-upogibu eksperimentalno in s simulacijo s kon~nimi elementi. Za oceno povratne elasti~ne izravnave so bili prednapetosti v
razponu od 0 % do 11 % izpostavljeni vzorci, ki so bili upognjeni v V-utopu. Uporabljen je bil komercialno dosegljiv softver
ETA Dyna form za simuliranje upogibanja v 60° V-utopu. Za simulacijo sta bili uporabljeni eksplicitna metoda kon~nih
elementov za fazo tla~enja in stati~na implicitna metoda kon~nih elementov za razbremenitev. Rezultati preizkusov in
simulacije ka`ejo, da prenapetost nima pozitivnega vpliva na kompenzacijo izravnave.
Klju~ne besede: prednapetost, povratna izravnava, AlMg zlitina, zlitina AA5754-O
1 INTRODUCTION
In recent years, lightweight structures have been a
key target for automotive manufacturers in order to
reduce fuel consumption and carbon dioxide emissions
without sacrificing vehicle safety and performance.
Therefore, lightweight materials, particularly aluminum
alloys, have found more applications in auto-body
structures. Many industries, such as aerospace, defense
and ship building, prefer aluminum alloys because of
their relatively light weight to high strength ratios and
corrosion resistance 1–5. However, there are some
limitations in the usage of these materials in terms of low
formability at room temperature (RT) and springback.
The springback issue is the most common problem in the
forming operations of these lightweight materials
because of their low Young’s modulus. It can lead to
significant problems during assembly if the phenomenon
is not well controlled, and so the manufacturing costs
will increase 6–8. In bending operations, after the release
of the load, an elastic recovery occurs. The geometry of
the part becomes quite different than the desired shape.
The springback issue has been studied over the years to
compensate for the undesired shape errors and to identify
the effect of major factors, such as material parameters,
tooling geometry, and process parameters, on the amount
of springback, both experimentally and numerically.
Asnafi 9 examined the effects of process parameters on
the springback in the V-bending process by developing
theoretical models for stainless-steel sheets. Asnafi 10
also studied the springback characterization of steel and
aluminum sheets in double-curved autobody panels, both
experimentally and theoretically. He reported that this
springback could be reduced by increasing the blank
holder force (BHF), sheet thickness, and die radius and
decreasing the yield strength. One of the most important
material parameters that affect the amount of springback
is the Bauschinger effect. The Bauschinger effect is
normally associated with conditions where the yield
strength of a metal decreases when the direction of the
strain is changed. The basic mechanism for the Bau-
schinger effect is related to the dislocation structure in
the cold-worked metal. Gau and Kinzel 11 experimentally
investigated the Bauschinger effect in steel and alumi-
num alloys using a simple bending process. They
showed that the Baushinger effect on the springback of
an aluminum alloy (AA6111-T4) is very significant.
Chun et al. 12,13 also studied the Bauschinger effect on a
sheet-metal forming process that was subjected to cyclic
loading conditions for different hardening rules. More-
over, a similar method was also used by Xue et al. 14,15.
They developed a new analytical procedure to predict the
springback of circular and square metal sheets after a
double-curvature forming operation 14,15.
In recent years, finite-element analysis (FEA) soft-
ware packages have become very popular as a rapid and
Materiali in tehnologije / Materials and technology 45 (2011) 6, 613–618 613
UDK 669.715:539.3:519.68 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 45(6)613(2011)
cost-effective tool for sheet-metal forming processes.
The developed models for the materials which are used
in FEA software have significant effects on the accurate
prediction of the forming results. For a numerical
simulation of a springback analysis, the appropriate
hardening model and the plastic yield criterion that
properly describe the material behavior at a large strain
are needed. There have been several studies in the
literature to predict the springback of materials by using
FEA programs based on the different hardening models,
i.e., isotropic, kinematic, and mixed hardening models
for different yield criteria and tool geometry, material
properties, forming conditions, etc. 16–22. Laurent et al. 22
compared several plastic yield criteria to show their
relevance with respect to predicting the springback
behavior for a AA5754-0 aluminum alloy, both
experimentally and numerically. In their study, the
effectiveness of the isotropic and kinematic hardening
models, which are combined with one of the following
plasticity criteria – von Mises, Hill’48 and Barlat’91 –
are analyzed.
In the present study, the effects of pre-straining on
the flow stress and the springback behavior of the
5754-O Al-Mg alloy were investigated. In addition to the
experimental work, finite-element (FE) simulations were
also used for the springback predictions and the
comparison.
2 MATERIAL AND EXPERIMENTAL WORK
In this study, as-received Alcan 5754-O Al-Mg alloy
sheet in the O-temper with a thickness of 1.88 mm was
evaluated. The chemical composition of the material is
given in Table 1.
Table 1: Chemical composition of 5754 Al-Mg alloy (in mass
fractions, w/%)
Tabela 1: Kemi~na sestava AlMg zlitine 5754 (v masnih dele`ih,
w/%)
Mg Si Mn Fe Cu Al
3.17 0.112 0.51 0.1613 0.01 Balance
Initial material properties are given in Table 2.
Tensile and pre-straining tests were performed on a
Shimadzu Autograph 100 kN testing machine with a
data-acquisition system maintained by a digital interface
board utilizing a specialized computer program. The ma-
terial deformation was measured with a video-exten-
someter measurement system for tensile-test specimens
with a 50-mm initial gauge length. Tensile-test samples
were prepared according to the ASTM E8 standard in the
rolling, "diagonal" and transverse directions. The tests
were conducted at room temperature and a strain rate of
0.0083 s–1 (25 mm/min) in order to determine the initial
properties of the material, which are shown in Table 2.
The specimens in the rolling direction were also
pre-strained by the tensile testing machine for six differ-
ent pre-straining levels, ranging from 1 % to 11%, by an
increment of 2 %. The pre-straining was performed at a
constant deformation rate of 3 mm/min and then un-
loaded at the same deformation rate. It is generally
known that the alloy shows serrated hardening curves, as
commonly observed for 5XXX series aluminum alloy
sheets 23. The yield strength of the material was deter-
mined based on the 0.2 % proof stress. As mentioned be-
fore, the Young’s modulus of the material has consider-
able affects on the springback behavior of the materials.
However, determining the Young’s modulus by perform-
ing tensile tests is very difficult because of the machine
competence and software limitations. Research in the lit-
erature shows that the Young’s modulus varies with the
plastic deformation and therefore using a constant
Young’s modulus in springback calculations and simula-
tions means less accurate results 24,25.
The 60° V-shaped die bending test samples were also
prepared at a rolling direction in a rectangular shape of
30 mm × 200 mm. They were all pre-stained using the
S. TOROS et al.: EFFECT OF PRE-STRAINING ON THE SPRINGBACK BEHAVIOR OF THE AA5754-0 ALLOY
614 Materiali in tehnologije / Materials and technology 45 (2011) 6, 613–618
Table 2: Initial material properties of 5754 Al-Mg alloy
Tabela 2: Za~etne lastnosti AlMg zlitine 5754
YS
(0.2 %)/MPa UTS/MPa UE/% TE/% r
Lankford
Parameter n K/MPa
Rolling (0o) 118 296 19.5 22.3 0.712 0.7325 0.306 492.2
Diagonal (45o) 106 234 23.2 26.2 0.754 0.304 462.5
Transverse (90o) 108 234 21.8 24.3 0.710 0.302 464.7
UE: Uniform elongation; TE: Total tensile elongation
Figure 1: 60° V-shaped bending setup
Slika 1: 60° V-utop za upogibanje
tensile testing machine, in the same way as the tensile
test samples.
In this study the springback evaluation was made by
a 60° V-shaped die bending test, as shown in Figure 1.
The bending tests were performed at a 25 mm/min
deformation speed. The precision of the displacement
and force measurements of the punch are 0.001 mm and
±2 N, respectively. No lubrication was applied to the die
and blank surfaces. The punch was released after the
forming. No soak time was assigned. The springback
angle () was measured by a Mitutoyo 187-907
universal bevel protractor that has a ±5min measurement
accuracy.
3 FINITE-ELEMENT STUDY
There have been many numerical approaches to
defining the springback characterization of materials for
bending operations. In general, these approaches are
directly related to the material properties, which are the
strain hardening coefficient (n), strength coefficient (K),
thickness (t), anisotropy (R), Young’s modulus (E),
Baushinger effects, hardening models, etc. In this
research, finite-element modeling was considered in
addition to the experimental study.
The material properties, i.e., Young’s modulus (E),
yield stress (Y) and strength coefficient (K), were
obtained from uniaxial tensile test and modified for the
plane strain conditions using von Mises criterion. The
bending process was also analyzed based on a
consideration of the plane strain condition. Plane strain
bending is a major sheet-forming process and it is
practiced as air bending, U- and V-shaped die bending.
The deformation in a bending process can be pronounced
as a plane strain deformation. In a plane strain deforma-
tion, the sheet usually extends only in one direction. For
this reason, the plane strain condition was also studied
and compared with the tensile deformation in order to
see the effect of the deformation mode.
The new E’, Y’, and K’ for the plane strain calculation
are as follows:
E
E
v
l =
−1 2
Y
Yl =
2
3
K
Kl =
2
3
In the calculations and simulations it was assumed
that the n values that were obtained under prescribed
pre-straining conditions during the uniaxial tensile tests
do not change with the uniaxial and plane strain
conditions.
The 60° V-shaped bending process was modeled
using a commercially available ETA/Dynaform finite-
element simulation program, as shown in Figure 2. In
the model, the die and punch were considered as rigid
bodies, and the blank was a deformable body. A
Belytschko-TSAY shell element was used for the blank
and rigid tools in order to improve the effectiveness of
the nonlinear numerical computation.
Besides the element type, the number of elements can
also affect the accuracy of the simulation results. In the
study, 4012 elements were used with five integration
points through the thickness of the deformable sheet. An
adaptive mesh option was also used in order to reduce
the errors during the calculation of the springback. In the
adaptive mesh method the elements are subdivided into
smaller elements during the analysis. This subdivision of
the elements provides improved accuracy and in the
study a two-times adaptivity was applied to the model.
The implicit and explicit solutions are two methods
that are used for the springback simulations. In the
simulation, the explicit loading and implicit unloading
approaches were used to predict the springback characte-
rization of the material. The implicit solution is realized
by applying a reverse nodal force and an equivalent
iteration. Due to the large deformations in the sheet-
metal forming operations, the amount of springback is
relatively large so the implicit solution is able to meet
these kinds of convergence forming operations. When
the accuracy of the stress field after the forming is poor,
the convergence problem becomes more serious 26.
As a material model, the material Type 36 (MAT_3-
PARAMETER_BARLAT) was used. This model was
developed by Barlat and Lian 27 in 1989 for the modeling
of anisotropic materials under plane stress-strain condi-
tions. This material allows the use of Lankford para-
meters 28 for the definition of anisotropy. The criterion
can be expressed as:
aIK1 + K2Im + bIK1 – K2Im + cI2K2Im = 2em
where a, b and c are the material constants that depend
on the anisotropy, K1 and K2 are the invariants of the
stress tensor and m is the stress exponent that is
S. TOROS et al.: EFFECT OF PRE-STRAINING ON THE SPRINGBACK BEHAVIOR OF THE AA5754-0 ALLOY
Materiali in tehnologije / Materials and technology 45 (2011) 6, 613–618 615
Figure 2: Finite element modeling of V bending process
Slika 2: Modeliranje V-upogibanja s kon~nimi elementi
calculated based on the crystallographic texture and is
equal to 8 for FCC materials.
4 RESULTS AND DISCUSSION
In the experimental study, the pre-straining was
applied to 5754-O Al-Mg alloy sheets at the prescribed
values. The tensile load and extension data were
converted to true stress vs. true strain data, which were
obtained at different pre-straining values. The true stress
vs. true strain diagrams were plotted, as seen in Figure
3. Each pre-straining path can be seen on the graph. The
mechanical properties were measured for all pre-strained
conditions at RT and a 0.0083 s–1 strain rate. The data
was tabulated as seen in Table 3. As also seen in Table
3, the yield strength of the material was significantly
increased from 118 MPa to 259 MPa. The change is
twice that of the initial value. It is known that if the ratio
of UTS/Y is high, a high springback is generally
observed. Increased pre-straining creates a high elastic
energy, which is the opposite to steels, is thought to be
the major reason for a high springback observation. The
change in UTS is not a large amount. It has an increasing
tendency up to 9 % pre-straining, then it starts to
decrease.
E is the one of the most important material properties
that affects the springback of the materials. Some
researchers have pointed out the effects of a variable
Young’s modulus on the formability of the materials 29–33.
The change in Young’s modulus was plotted with respect
to the pre-straining value, as shown in Figure 4. As seen
from Table 3, the value was much lower than the known
value. Normally, it is very difficult to determine the
Young’s modulus with a tensile test. The purpose of the
study is to evaluate the springback variation with
pre-straining relatively. For this reason it is not paid
attention to the values. The same procedure was applied
for each condition, which means it does not make any
significant changes for the final outcome. The important
point is that the tendency of the E based on pre-straining
was determined. Figure 4 indicates two linear curves
and their equations. Although the Young’s modulus of
the material shows fluctuation with prescribed
pre-straining levels, the trend of the values is increasing
with plastic deformation.
Figures 5 and 6 show the variations of K and n with
respect to the prescribed pre-straining.
Similar behaviors were observed in Figures 5 and 6.
A minor increase was seen up to 3 % pre-straining and a
minor decrease was seen after that. K has the highest
value at 3 % pre-staining. As seen from the results, it is
very complicated to clarify the changes. It may be
S. TOROS et al.: EFFECT OF PRE-STRAINING ON THE SPRINGBACK BEHAVIOR OF THE AA5754-0 ALLOY
616 Materiali in tehnologije / Materials and technology 45 (2011) 6, 613–618
Figure 3: True stress vs. true strain curves for the 5754 Al-Mg alloy
Slika 3: Odvisnosti prava napetost – prava deformacija za zlitino
AlMg 5754
Table 3: Summary of material properties
Tabela 3: Povzetek lastnosti materiala
Pre-straining (%) Y/MPa UTS/MPa E/MPa K/MPa n
0 118.56 296.80 47.17 492.2 0.310
1 125.66 296.80 57.46 516.3 0.323
3 174.62 297.75 58.61 512.6 0.328
5 204.89 296.53 61.34 497.4 0.305
7 225.81 306.30 65.55 482.8 0.297
9 245.85 308.56 74.72 455.1 0.263
11 259.18 300.94 81.24 449 0.260
Figure 4: Variation of the Young’s modulus vs. pre-straining
Slika 4: Sprememba Youngovega modula v odvisnosti od prednape-
tosti
explained with the changes in the microstructure,
including the dislocation mechanism, porosities,
inclusion, and deformation rate, etc.
Finally, the evaluation of the springback was
investigated by experiment and finite-element methods.
All the results were plotted on the same graph and
displayed in Figure 7. The reference curve is the
experimental curve. Any prediction that is close to this
curve is considered to be a most accurate prediction.
The experimental data in Figure 7 reveals that the
springback is linearly increasing with increasing
pre-straining. Pre-straining has no positive contribution
on the springback compensation. It is related to the
increasing elastic energy during the pre-straining.
Finite-element predictions were determined for the
uniaxial and plane strain conditions. They were different
from each other. The predictions at 1 % pre-straining
were in good agreement with the experiments. But the
predictions at no pre-straining and 5 % were higher than
the experiments. At 3 % pre-straining, the FE prediction
for the plane strain condition is almost the same as the
experiment. At 7 % and over, the FE predictions for
uniaxial were in accord with the experiment, except for
the 11 % pre-straining. But the FE prediction for plane
stress condition was lower than the experiments. In
general, the FE predictions were close to the experi-
mental results. The simulation results that were obtained
for (1, 7 and 9) % pre-straining were close to the
experimental results for the uniaxial strain condition. All
these findings suggest that the pre-straining does not
help the springback compensation.
5 CONCLUSION
In this study, the effect of pre-straining on springback
was investigated for a 5754-O Al-Mg alloy, both
experimentally and numerically. It was found that the
springback was increased with increasing pre-straining.
The finite-element predictions were in partially good
agreement with the experimental data.
Acknowledgement
This work is supported by The Scientific and Tech-
nological Research Council of Turkey (TÜBÝTAK).
Project Number: 106M058, Title: "Experimental and
Theoretical Investigations of The Effects of Temperature
and Deformation Speed on Formability". TÜBÝTAK’s
support is gratefully acknowledged.
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S. TOROS et al.: EFFECT OF PRE-STRAINING ON THE SPRINGBACK BEHAVIOR OF THE AA5754-0 ALLOY
618 Materiali in tehnologije / Materials and technology 45 (2011) 6, 613–618
M. BALCAR et al.: HEAT TREATMENT AND MECHANICAL PROPERTIES OF HEAVY FORGINGS ...
HEAT TREATMENT AND MECHANICAL PROPERTIES
OF HEAVY FORGINGS FROM A694–F60 STEEL
TOPLOTNA OBDELAVA IN MEHANSKE LASTNOSTI TE@KIH
IZKOVKOV IZ JEKLA A694-F60
Martin Balcar1, Jaroslav Novák1, Libor Sochor1, Pavel Fila1, Ludvík Martínek1,
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 TU Ostrava – FMMI, 17. listopadu 15/2172, 708 33 Ostrava-Poruba, Czech Republic
3Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
martin.balcar@zdas.cz
Prejem rokopisa – received: 2011-05-31; sprejem za objavo – accepted for publication: 2011-09-21
The production of heavy steel forgings of micro-alloyed steels gives the possibility to obtain advantages associated with the
benefit of the application of micro-alloying elements and thermomechanical treatments for improving the mechanical properties
of forgings to level by sheets, strips and tubes. The paper presents the influence of quenching temperature on the mechanical
properties and microstructure of F60 steel according to ASTM A694. The verification of the influence of quenching temperature
contributes to an optimization of the method of micro-alloyed steel heat treatment. The steel’s microstructure and mechanical
properties after quenching constitute the initial and basic criteria to achieve the required mechanical properties after a properly
chosen tempering temperature.
Keywords: HSLA steel, A694 F605, quenching and tempering
Proizvodnja te`kih izkovkov iz mikrolegiranih jekel omogo~a, da se uporabijo prednosti mikrolegiranja in termomehanske
obdelave za doseganje mehanskih lastnosti pri trakovih, plo{~ah in ceveh. V ~lanku je predstavljen vpliv temperature kaljenja na
mehanske lastnosti in mikrostrukturo. Preveritev vpliva temperature kaljenja je del procesa opredelitve in optimizacije metode
toplotne obdelave mikrolegiranega jekla. Mikrostruktura in mehanske lastnosti po toplotni obdelavi so osnovni pogoj za
doseganje predpisanih lastnosti.
Klju~ne besede: HSLA-jeklo, A694 F605, kaljenje in popu{~anje
1 INTRODUCTION
As the requirements for the properties of structural
steel are increasing, the development of the use of mi-
cro-alloying elements, even in the field of the production
of forgings and castings, takes place. The production of
heavy steel forgings of micro-alloyed steels does not al-
low the use of the advantages associated with the benefit
of the application of micro-alloying elements and
thermomechanical treatment known from the production
of sheets, strips and tubes. The production of steel
forgings involves forming and heat-treatment processes,
which are significantly different than those for
thin-walled products (sheets, strips, tubes).
The development and verification of F60 steel pro-
duction and treatment technology according to ASTM
A694 in ZDAS, Inc. conditions constituted a number of
technological changes and the introduction of new pro-
cess elements in the field of steel making and
thermomechanical treatment. The verification of the in-
fluence of the quenching temperature on the properties
and the microstructure of F60 forged steel contributes to
the optimization of the HSLA steel-making technology
at ZDAS, Inc.
2 EXPERIMENTAL MATERIAL
The verification of the influence of quenching tem-
perature on the microstructure and mechanical properties
of modified F60 steel according to ASTM A694 made by
EOP/LF/VD technology was carried out on forged sam-
ples with dimensions 100 mm × 100 mm × 150 mm. The
basic chemical composition of the steel is shown in Ta-
ble 1.
The F60 steel modified according to ASTM A694 is
a typical low-carbon steel with the addition of the alloy-
ing elements, manganese, nickel and molybdenum.
Moreover, the steel is micro-alloyed with vanadium, alu-
minium and niobium. The content of other elements is at
the level of residuals.
Materiali in tehnologije / Materials and technology 45 (2011) 6, 619–622 619
UDK 669.14.018.298:621.785:620.17 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 45(6)619(2011)
Table 1: Basic chemical composition of HSLA steel F60 in mass fraction, w/%
Tabela 1: Osnovna kemi~na sestava HSLA jekla F60 v masnih dele`ih, w/%
C Mn Si P S Cr Ni Mo V Al Nb N
0.10 1.08 0.33 0.003 0.001 0.16 0.77 0.27 0.04 0.027 0.034 0.0037
After the forming process, the forgings were "anti-
flake" annealed at a temperature of 650 °C for a period
of 10 h and then normalized at a temperature of 930 °C
with air cooling.
3 LABORATORY HEAT TREATMENT
The heat treatment was carried out on forged steel
samples in the laboratory. The verification of the influ-
ence of the austenitization–quenching temperature (TA)
on the microstructure and mechanical properties was per-
formed for the temperature range of 880 °C to 940 °C
with water quenching, tempering TP = 620 °C and air
cooling. The sample markings and heat treatment were
carried out as follows:
Sample L1: TA = 880 °C/6 h/water + TP = 620 °C/8 h/air
Sample L2: TA = 890 °C/6 h/water + TP = 620 °C/8 h/air
Sample L3: TA = 900 °C/6 h/water + TP = 620 °C/8 h/air
Sample L4: TA = 910 °C/6 h/water + TP = 620 °C/8 h/air
Sample L5: TA = 920 °C/6 h/water + TP = 620 °C/8 h/air
Sample L6: TA = 930 °C/6 h/water + TP = 620 °C/8 h/air
Sample L7: TA = 940 °C/6 h/water + TP = 620 °C/8 h/air
4 MECHANICAL PROPERTIES OF F60 HSLA
STEEL
The samples for determining the achieved mechani-
cal properties and to evaluate the microstructure were
taken from the central zone of the forgings in the longi-
tudinal direction. In table 2 the requested level and at-
tained values of the mechanical properties of individual
F60 steel samples are shown.
The influence of the austenitization temperature on
the change in the mechanical properties of forged,
quenched and tempered F60 steel is visible from Table
2. It is obvious that the steel’s strength increases and a
significant toughness drop occurs with an increase of the
austenitization temperature. An austenitization tempera-
ture of over 910 °C causes the steel to become brittle.
5 MICROSTRUCTURE OF SAMPLES OF HSLA
STEEL ASTM A694 F60
Like in the case of the mechanical properties, the
steel microstructure was evaluated in the control zone of
samples.
The steel microstructure for the heat-treatment states
(TA = (880, 900, 920, 940) °C) is shown in Figure 2 to 5:
After quenching and tempering, the microstructures
of all the sample forgings are practically the same and it
consists of ferrite, bainite, granular pearlite and sorbite.
It is evident from the micrographs where the secondary
grain size can be compared more easily, that the second-
ary grain size does not change noticeably with an in-
crease of the quenching temperature. This is confirmed
M. BALCAR et al.: HEAT TREATMENT AND MECHANICAL PROPERTIES OF HEAVY FORGINGS ...
620 Materiali in tehnologije / Materials and technology 45 (2011) 6, 619–622
Figure 1: Forging specimen - 100 mm × 100 mm × 150 mm
Slika 1: Odkovek, 100 mm × 100 mm ×150 mm
Table 2: Mechanical properties HSLA steel F60 after different austenization temperatures
Tabela 2: Mehanske lastnosti HSLA-jekla F 60 po avstenitizacij pri razli~nih temperaturah
TA/°C Re/MPa Rm/MPa A5/% Z/% KV–46 °C/J AVG KV–46 °C/J
415–565 520–760 min. 20 min. 35 ø KV min. 63 ø KV min. 63
L1 880 548 639 21.6 76.0 299 300 217 272
L2 890 550 653 22.2 75.0 255 229 286 257
L3 900 561 653 21.8 75.0 213 217 218 216
L4 910 573 667 22.2 75.0 101 214 89 135
L5 920 576 662 23.0 76.0 189 137 27 118
L6 930 576 672 22.8 74.0 124 204 238 189
L7 940 576 671 22.4 75.0 86 153 31 90
Table 3: Austenitic grain size - HSLA steel F60 -ASTM E 112 –
LECO IA32
Tabela 3: HSLA-jeklo F 60 - ASTM E 112 - LECO IA32
TA/°C Grain size /μm
L1 880 11.3 ± 0.4
L2 890 11.2 ± 0.4
L3 900 10.7 ± 0.3
L4 910 11.3 ± 0.5
L5 920 11.4 ± 0.2
L6 930 9.4 ± 0.6
L7 940 10.3 ± 0.4
M. BALCAR et al.: HEAT TREATMENT AND MECHANICAL PROPERTIES OF HEAVY FORGINGS ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 619–622 621
Figure 4: Sample L5: TA = 920 °C/6 h/water + TP = 620 °C/8 h/air
Slika 4: Vzorec L5: TA = 920 oC/ 6 h/ voda + Tp = 620 oC/ 8 h/zrak
Figure 2: Sample L1: TA = 880 °C/6 h/water + TP = 620 °C/8 h/air
Slika 2: Vzorec L1: TA = 880 oC/ 6 h/ voda + Tp = 620 oC/ 8 h/zrak
Figure 5: Sample L7: TA = 940 °C/6 h/water + TP = 620 °C/8 h/air
Slika 5: Vzorec L7: TA = 940 oC/ 6 h/ voda + Tp = 620 oC/8 h/zrak
Figure 3: Sample L3: TA = 900 °C/6 h/water + TP = 620 °C/8 h/air
Slika 3: Vzorec L3: TA = 900 oC/ 6 h/ voda + Tp = 620 oC/8 h/zrak
by the results of the assessment of the austenite grain
size by the oxidation method according to ASTM E 112
– 97 using LECO IA32 image analysis. The results of
the austenitic grain-size assessment are shown in Table
3.
From the results in Table 3 it is not possible to estab-
lish the direct influence of quenching temperature on the
austenite grain-size change as all the samples show a
very fine grain size.
6 CONCLUSIONS
From the results of the experimental work we can see
the direct influence of the quenching temperature on the
mechanical properties of the F60 steel.
A slight increase of strength and a strong drop in im-
pact value was found for an increase of the quenching
temperature. The most favourable results of the mechani-
cal properties were attained with the quenching tempera-
tures 880 °C, 890 °C and 900 °C.
The steel microstructure after quenching and temper-
ing is similar for all samples and consists of ferrite,
bainite, granular pearlite and sorbite. The assessment of
the austenite grain size by the oxidation method con-
firmed the grain-size uniformity, when comparing exper-
imental samples, without any provable influence of the
quenching to a temperature of 920 °C.
A further optimization of the steel’s mechanization
properties and microstructure is expected after a verifica-
tion of the influence of the tempering temperature. Sub-
sequently, it will be possible to determine a complex op-
timized heat-treatment process for the HSLA steel
ASTM A694 F60.
In this paper the results obtained in the EUREKA
programme of the E!4092 MICROST project are pre-
sented. The project was realized with the financial sup-
port of the Ministry of Education, Youth and Sport of
the Czech Republic.
M. BALCAR et al.: HEAT TREATMENT AND MECHANICAL PROPERTIES OF HEAVY FORGINGS ...
622 Materiali in tehnologije / Materials and technology 45 (2011) 6, 619–622
R. PALANIVEL, P. KOSHY MATHEWS: THE TENSILE BEHAVIOUR OF FRICTION-STIR-WELDED ...
THE TENSILE BEHAVIOUR OF FRICTION-STIR-
WELDED DISSIMILAR ALUMINIUM ALLOYS
NATEZNE ZNA^ILNOSTI TORNIH POMI^NIH ZVAROV
RAZLI^NIH ALUMINIJEVIH ZLITIN
R. Palanivel1, P. Koshy Mathews2
1Faculty in Mechanical Engineering, Kalaivani College of Technology, Coimbatore, India
2Dean, Research and Development, Kalaivani College of Technology, Coimbatore, India
rpelmech@yahoo.co.in
Prejem rokopisa – received: 2011-04-01; sprejem za objavo – accepted for publication: 2011-07-27
Aluminium alloys generally have a low weldability with the traditional fusion-welding process. However, the development of
Friction Stir Welding (FSW) has provided an alternative, improved way of producing aluminium joints, in a faster and more
reliable manner. The FSW process has several advantages, in particular the possibility to weld dissimilar aluminium alloys. This
study focuses on the tensile behaviour of dissimilar joints of AA6351-T6 alloy to AA5083-H111 alloy produced by friction stir
welding. Five different tool pin profiles, such as Straight Square (SS), Tapered Square (TS), Straight Hexagon (SH), Straight
Octagon (SO) and Tapered Octagon (TO), with three different welding speeds (50 mm/min, 63 mm/min, 75 mm/min) have been
used to weld the joints. The effect of the pin profiles and the welding speed on the tensile properties was analyzed and it was
found that the straight square pin profile with 63 mm/min produced a better tensile strength then the other tool pin profiles and
welding speeds.
Key words: friction stir welding, aluminium alloys, tool pin profile, welding speed, tensile properties
Aluminijeve zlitine so slabo varive po tradicionalnih talilnih postopkih. Razvoj tornega pomi~nega varjenja (FSW) je ponudil
mo`nost priprave aluminijevih zvarov na hiter in zanesljiv na~in. Proces ima ve~ prednosti, predvsem mo`nost varjenja razli~nih
aluminijevih zlitin. Raziskava je bila osredinjena na natezno vedenje razli~nih zvarov zlitin AA6351-T6 in AA5083-H111,
pripravljenih s tornim pomi~nim varjenjem. Pet razli~nih profilov trna je bilo uporabljenih: raven kvadrat (SS), koni~en kvadrat
(TS), raven {estkokotnik (SH), raven osemkotnik (SO) in koni~en osemkotnik (TO) s tremi razli~nimi hitrostmi varjenja (50
mm/min, 63 mm/min in 75 mm/min). Dolo~en je bil vpliv oblike trna in hitrosti varjenja na natezne lastnosti. Ugotovljeno je, da
raven kvadraten trn pri hitrosti 63 mm/min da bolj{o trdnost kot drugi profili trna in druge hitrosti varjenja.
Klju~ne besede: pomi~no torno varjenje, aluminijeve zlitine, profil tornega trna, hitrost varjenja, natezne lastnosti
1 INTRODUCTION
Friction stir welding (FSW) is a solid-state welding
process developed by The Welding Institute (UK) in
1991, and now being used increasingly for joining
aluminium alloys, for which fusion welding is often
difficult. FSW uses a rotating tool with a probe travelling
along the weld path, and plastically deforms the
surrounding material to form the weld. Since the
material subjected to FSW does not melt and recast, the
resultant weld offers advantages over conventional
fusion welds, such as less distortion, lower residual
stresses and fewer weld defects.1–3 When developing
such a technology, one of the most important factors is
the possibility to join different aluminium alloys. 4 The
development of sound joints between dissimilar mate-
rials is a very important consideration for many
emerging applications, including ship building,
aerospace, transportation, power generation, as well as
the chemical, nuclear, and electronics industries 5. How-
ever, the joining of dissimilar materials by conventional
fusion welding is difficult because of the poor welda-
bility arising from the different chemical, mechanical,
and thermal properties of welded materials and the
formation of hard and brittle intermetallic compounds
(IMCs) on a large scale at the weld interface. The
absence of melting in friction stir welding (FSW)
provides a strong tendency to produce reliable dissimilar
joints. Amancio-Filho et al.6 determined the tensile
strength of dissimilar friction stir welded AA2024-T351
and AA6056-T4 as 56 % of the AA2024-T351 and 90 %
of the AA6056-T4. It is reported that the poor tensile
strength observed in these joints is due to the thermal
softening of the base metals, and the poor ductility
observed in these joints is due to the stress concentration
caused by the large difference in strength between the
base metals leading to confined plasticity and failure.
Cavaliere et al.[4] investigated the tensile behaviour of
dissimilar friction stir welded joints of the aluminium
alloys 2024-T3 and 7075-T6 and reported that both the
ultimate strength and the elongation of the dissimilar
joints are lower than the base metals 2024-T3 and
7075-T6. From the above literature review we can
conclude that very little research work has been carried
out on the dissimilar FS welding of aluminium alloys
and that the dissimilar friction stir welding of AA6351
and AA5083, which are widely used in aerospace, ship
building, and other fabrication industries,7 were not
investigated. Hence, the present research work focuses
Materiali in tehnologije / Materials and technology 45 (2011) 6, 623–626 623
UDK 669.715:621.791 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 45(6)623(2011)
on the tensile behaviour of dissimilar friction stir welded
joints of the aluminium alloys AA6351 and AA5083.
2 EXPERIMENTAL PROCEDURE
2.1 Manufacturing of FSW tools
Five different tools made of High Carbon High Chro-
mium steel (HCHCr) having different pin profiles of
Straight Square (SS), Tapered Square (TS), Straight
Hexagon (SH), Straight Octagon (SO) and Tapered Octa-
gon (TO) without draft were used to weld the FSW
joints. Each tool had a shoulder of diameter 18 mm, a
pin diameter of 6 mm and a pin length of 5.6 mm. The
shoulder–workpiece interference surface had 3 concen-
tric circular equally spaced slots of 2 mm depth on all
the tools. The FSW tools were manufactured using a
CNC turning center and a wire cut EDM (WEDM) ma-
chine to get an accurate profile. The tools were oil hard-
ened. The manufactured tools are shown in Figure 1.
2.2 Frictions stir welding of dissimilar aluminium
alloys
The aluminium alloys AA6351-T6 and AA5083-
H111 were selected for the dissimilar friction stir
welding process. The chemical compositions of the
materials AA6351-T6 and AA5083-H111 are presented
in Tables 1 and 2 and the mechanical properties of the
materials are presented in Table 3. Test plates of size
100 mm × 50 mm × 6 mm were prepared from rolled
sheets.
The experimental set up consists of a special-purpose
machine shown in Figure 2 with arrangements designed
for the friction stir welding. The plate AA 6351-T6 was
fixed with the advancing side and the AA5083 H-111
was fixed with the retreating side of the machine. The
vertical tool head can be moved along the vertical guide
ways (Z-axis). The horizontal table can be moved along
the X- and Y-axes and consists of mechanical fixtures to
hold the workpieces rigidly. The machine can be opera-
ted over a wide range of tool rotational speeds, welding
speeds and tool axial forces. Five different tool-pin
profiles were used to produce the joints. Using each tool,
three joints at three different welding speed levels and in
R. PALANIVEL, P. KOSHY MATHEWS: THE TENSILE BEHAVIOUR OF FRICTION-STIR-WELDED ...
624 Materiali in tehnologije / Materials and technology 45 (2011) 6, 623–626
Table 1: Chemical composition of AA6351-T6
Tabela 1: Kemi~na sestava zlitine AA6351-T6
Name of the Element Si Zn Mg Mn Fe Cu Ti Sn Ni Al
Masss fractions, w/%
(AA6351) 0.907 0.89 0.586 0.65 0.355 0.086 0.015 0.003 0.002 Balance
Table 2: Chemical composition of AA5083-H111
Tabela 2: Kemi~na sestava zlitine AA5083-H111
Element Si Zn Mg Mn Fe Cu Ti Al
Mass fracions, w/%
(AA5083-H111) 0.045 0.04 4.76 0.56 0.14 0.02 0.054 Balance
Figure 2: Experimental setup
Slika 2: Eksperimentalna naprava
Figure 1: Manufactured tool for FSW (Straight Square (SS), Straight
Hexagon (SH), Straight Octagon (SO), Tapered Square (TS), and
Tapered Octagon (TO)
Slika 1: Trni za pomi~no torno varjenje: raven kvadrat (SS), koni~en
kvadrat (TS), raven {estokokotnik (SH), raven osmokotnik (SO) in
koni~en osmokotnik (TO)
total 15 joints (5 × 3) were produced in this study. The
welding parameters are presented in Table 4. The joints
were visually inspected for exterior weld defects and
they were found to be free from any external defects. A
sample of a friction stir welded plate is shown in Figure
3. The tensile test specimens were prepared according to
the ASTM E8 standard and the transverse tensile
properties of the FS welded joints were evaluated using a
computerized (Universal Testing Machine) UTM. For
each welded plate, three specimens were prepared and
tested. The fracture occurred either on the retreating side
or the advancing side of the weld. Figure 4 shows the
fractured tensile specimen.
3 RESULTS
The effects of welding speed for various tool pin
profiles are shown in Figure 5. At the lowest (50
mm/min) and the highest welding speeds (75 mm/min) a
lower tensile strength was observed. This trend was
common in all the joints, irrespective of the tool pin
profile. The joint produced by the straight square pin
profiled tool exhibits a high tensile strength when
compared to the other joints. The joint produced by the
tapered octagon pin profiled tool had the lowest tensile
strength. The tensile strength of the joints welded using
the straight hexagon and the straight octagon pin profiled
tools did not differ significantly.
4 DISCUSSION
The increase in welding speed leads to an increase in
the tensile strength up to a maximum value, while a
further increase in the welding speed results in a
decrease of the tensile strength of the FS welded joints.
This is due to the increased frictional heat and insuffi-
cient frictional heat generated, respectively. 8 In general,
FSW at higher welding speeds results in a short exposure
time in the weld area with insufficient heat and a poor
plastic flow of the metal and causes some void-like
defects in the joints. The reduced plasticity and rates of
diffusion in the material may have resulted in a weak
interface. Higher welding speeds are associated with low
R. PALANIVEL, P. KOSHY MATHEWS: THE TENSILE BEHAVIOUR OF FRICTION-STIR-WELDED ...
Materiali in tehnologije / Materials and technology 45 (2011) 6, 623–626 625
Figure 5: Effect of welding speed and pin profile
Slika 5: Vpliv hitrosti varjenja in profila trna
Figure 3: FS Welded sample (950 r/min, 1 t, 63 mm/min straight
square pin profile)
Slika 3: Zvarjeni vzorec (950 r/min, 1 t, 63 mm/min, raven kvadraten
profil)
Table 3: Mechanical properties of the AA6351 and AA5083-H111
Tabela 3: Mehanske lastnosti zlitin AA6351-T6 in AA5083-H111
Base Material
Tensile
Strength
(MPa)
Yield
Strength
(MPa)
Percentage
of
elongation
AA6351 310 285 14
AA5083-H111 308 273 23
Table 4: Welding process parameters
Tabela 4: Parametri procesa varjenja
Process parameter Values
Tool rotational speed, r/min 950
Welding speed, mm/min 50,63,75
Axial force, t 1
Figure 4: Tested specimen for tensile strength
Slika 4: Pretrgani natezni preizku{anci
heat inputs, which result in faster cooling rates of the
welded joint. This can significantly reduce the extent of
the metallurgical transformations taking place during
welding and the local strength of the individual regions
across the weld zone. 9 The pin profile plays a crucial
role in the material flow and in turn regulates the
welding speed of the FSW process. The relationship
between the static volume and the dynamic volume
decides the path for the flow of plasticized material from
the leading edge to the trailing edge of the rotating tool
10. This ratio is equal to 1.56 for the straight square, 1.21
for the straight hexagon, 1.11 for the straight octagon,
2.04 for the tapered octagon and 3.51 for the tapered
square pin profiles. In addition, these pin profiles
produce a pulsating stirring action in the flowing
material due to the flat faces. The square pin profile
produces 63 pulses per second, the hexagon pin profile
produces 95 pulses per second and the octagon pin
profile produces 126 pulses per second, when the tool
rotates at a speed of 950 r/min. There is not much
pulsating action in the case of the octagonal and
hexagonal pin profiled tool because it almost resembles a
straight cylindrical pin profiled tool at this high rpm. In
the tapered pin profiled tools, the same principle affects
the material flow. Since the tapered square and tapered
octagon pin profile sweeps less material when compared
to that of the straight square pin tool, this joint exhibit
less tensile properties.
5 CONCLUSION
Among the fifteen joints produced in this investi-
gation, the joints produced using the straight square pin
profiled tool at a welding speed of 63 mm/min showed
the best tensile properties.
6 REFERENCES
1 R. S. Mishraa, Z. Y. Ma, Frictions stir welding and processing.
Materials Science and Engineering R, Reports, 50 (2005) 1–2, 1–78
2 R. Nandan, T. DebRoy, H. K. D. H. Bhadeshia, Recent advances in
friction-stir welding process, weldment structure and properties.
Progress in Material Science, 53 (2008) 6, 980–1023
3 Y. Uematsu , K. Tokaji, H. Shibata, Y. Tozaki, T. Ohmune, Fatigue
behavior of friction stir welds without neither welding flash nor flaw
in several aluminium alloys. International Journal of Fatigue, 31
(2009) 10, 1443–1453
4 P. Cavaliere, E. Cerri, A. Squillace, Mechanical response of
2024-7075 aluminium alloys joined by friction stir welding. Journal
of Material Science, 40 (2005) 14, 3669–76
5 T. Saeid, A. Abdollah-zadeh, B. Sazgari, Weldability and mechanical
properties of dissimilar aluminum–copper lap joints made by friction
stir welding. Journal of Alloys and Compounds, 490 (2010) 1–2,
652–655
6 S. Tamancio-Filho, S. Sheikhi, J. F. Dos Santos, C. Balfarini. Prelim-
inary study on the microstructure and mechanical properties of
dissimilar friction stir welds in aircraft aluminium alloys 2024-T351
and 6056-T4. Journal of Material processing Technology, 206 (2008)
1–3, 32–42
7 H. D. Chandler, J. V. Bee, Cyclic strain induced precipitation in a
solution treated aluminum alloy. Acta Metallurgica, 35 (1987) 10,
2503–2510
8 Kevin J. Colligan, Paul J. Konkol, James, J. Fisher, Joseph R.
Pickens, Friction stir welding demonstrated for combat vehicle
construction. Welding Journal, 82 (2003) 3, 1–6
9 A. V. Strombeck, J. F. D. Santos, F. Torster, P. Laureano, M. Kocak.
Fracture toughness behavior of FSW joints on aluminum alloys,
Proceedings of the First International Symposium on Friction Stir
Welding California, USA, 1999, Paper No. S9-P1
10 K. Elangovan, V. Balasubramanian, S. Babu, Predicting tensile
strength of friction stir welded 6061 aluminium alloy joints by
mathematical model. Material and Design, 30 (2009) 1, 188–193
R. PALANIVEL, P. KOSHY MATHEWS: THE TENSILE BEHAVIOUR OF FRICTION-STIR-WELDED ...
626 Materiali in tehnologije / Materials and technology 45 (2011) 6, 623–626
M. @VEGLI^ et al.: SCREEN-PRINTED ELECTRICALLY CONDUCTIVE FUNCTIONALITIES IN PAPER SUBSTRATES
SCREEN-PRINTED ELECTRICALLY CONDUCTIVE
FUNCTIONALITIES IN PAPER SUBSTRATES
ELEKTROPREVODNE OBLIKE, PRIPRAVLJENE S SITOTISKOM
NA PAPIRNIH PODLAGAH
Ma{a @vegli~1, Nina Hauptman1, Marijan Ma~ek2, Marta Klanj{ek Gunde1*
1National Institute of Chemistry, Hajdrihova 19, Ljubljana, Slovenia
2University of Ljubljana, Faculty of Electrical Engineering, Tr`a{ka 25, Ljubljana, Slovenia
marta.k.gunde@ki.si
Prejem rokopisa – received: 2011-05-11; sprejem za objavo – accepted for publication: 2011-07-08
The topological and electrical properties of screen-printed conductive lines, applying silver-based conductive ink are analyzed.
The influence of the substrate, curing time and wet-to-wet overprinting is analyzed. The mechanical profilometer, optical
microscope and SEM images did not give the same information about the topography of the prints. Areas without functional
particles on the line boundaries could be seen on the SEM micrographs, whereas the profilometer and optical microscope could
not support such information, and thus provide wider lines. Careful analysis confirms that the printing parameters influence the
electrical resistivity of the printed products. Overprinting does not influence a great deal on the shape of lines; however, these
lines have a smaller resistivity. The larger resistivity of the prints was obtained on a rough and porous substrate and was smaller
on the smooth one. The influence of curing time was also shown.
Keywords: printed electronics, screen printing, conductive ink, electrical resistivity
Analizirali smo topolo{ke in elektri~ne lastnosti linij, ki so bile natisnjene s sitotiskom prevodne tiskarske barve s srebrovimi
delci. Ugotavljali smo vpliv tiskovne podlage, ~asa termi~nega su{enja in ve~kratnega tiska mokro-na-mokro. Mehanski
profilometer, opti~ni in elektronski mikroskop ne dajejo nujno enakih podatkov o topografiji potiskanih oblik. Obmo~ja brez
funkcionalnih delcev opazimo le na posnetkih elektronskega mikroskopa, opti~ni mikroskop in profilometer pa taka obmo~ja ne
lo~ita od funkcionalnih delov linij. Analiza je pokazala, da parametri tiska vplivajo na elektri~ne lastnosti izdelkov. Tisk
mokro-na-mokro ima zanemarljiv vpliv na tiskane oblike, pa~ pa imajo taki odtisi manj{o elektri~no upornost. Na hrapavi in
porozni podlagi dobimo ve~je upornosti kot na gladki. Na upornost potiskanih linij vpliva tudi ~as segrevanja pri su{enju
odtisov.
Klju~ne besede: tiskana elektronika, sitotisk, prevodna tiskarska barva, elektri~na prevodnost
1 INTRODUCTION
Printed electronics, i.e., fabricating an entire elec-
tronic device by printing, is expected to provide low-cost
electronic systems on common surfaces such as paper,
plastic and textiles. The simplified structures of
electronic devices, printed by a minimal number of
different printing inks, should provide the lowest
possible target price, which was estimated to be below
0.2  per piece 1.
All printed electronic devices require some printed
conductor to replace the metal layers used in conven-
tional electronics. Polymer inks containing electrically
conductive particles are the most common choice for this
purpose in the field of printed electronics. They consist
of a suitable polymer resin with metal particles, which in
most cases are silver, gold, copper, nickel, platinum or
carbon 2–6. These conductive inks have various resisti-
vities; the lowest was obtained with silver particles.
Specific physical properties, such as viscosity, suitable
rheology characteristics and appropriate curing, are
demanded for each printing technology to obtain feasible
prints of acceptable quality 7–9.
Conductive inks are available on the market for
conventional technologies, i.e., screen printing, offset,
and pad-printing. In most cases the producers provide
resistivity data for a layer with a specified thickness
prepared by a recommended application (printing)
method and drying conditions. However, the resistivity
of a shape, printed by a particle-based conductive
printing ink, depends on the internal microstructure of
the printed lines. This structure could be influenced by
several parameters that may affect the functional
properties of the final application 5–7,9.
The objective of our research was to analyze the
influence of the printing parameters on the resistivity of
screen-printed lines using a silver-based conductive ink.
Two flexible substrates were used, the gloss-coated paper
and the clear matt film. The topology of screen-printed
lines was examined with several techniques. The resisti-
vity of the test structures was measured and the influence
of printing parameters was analyzed thoroughly.
2 EXPERIMENTAL
Electrodag PM-470 conductive screen-printable ink
(Acheson Colloiden B.V., Netherlands) was used. It con-
tains finely distributed silver particles in a thermoplastic
resin. Its density is about 2140 kg/m3 and the solid con-
tent 58–62 %. The manufacturer specifies the Broofield
Materiali in tehnologije / Materials and technology 45 (2011) 6, 627–632 627
UDK 621.38:620.1/.2 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 45(6)627(2011)
viscosity of the ink to be 11 000–140 000 mPa s (30 °C,
20 r/min) and the sheet resistance of a 25-μm-thick layer
at a sheet resistance of 0.008–0.015 /. This corre-
sponds to a specific resistivity of 2.0–3.75 · 10–5  cm.
The print form contained suitable structures for
resistivity measurements and several horizontal and
vertical lines (in series of four equally separated strips)
with a width from 5 mm to 0.5 mm. The shape of
narrowest lines (500 μm width on print form) was
evaluated in detail.
The ink was screen printed by applying the SEFAR®
high-modulus monofilament polyester plain weave mesh
43/80Y and a squeegee with a hardness of 75 °Sh. Two
substrates were used, i.e., clear matt film (thermally and
antistatically treated for transfer printing) and
gloss-coated paper. Single and double layer prints
(wet-on-wet) were made. The off-prints were cured at
120 °C for 4, 9 and 13 min. In all cases, dry prints were
obtained.
The thickness and profile of the lines were measured
with a Talysurf profilometer (Rank Taylor Hobson Series
2). The microtexture and profile of the narrow lines were
monitored with a scanning electron microscope SEM –
6060 LV (JEOL, Japan). The shape of the edges, the
degree of wicking and the width of narrowest line were
also evaluated using a Nanometrics optical microscope
(Olympus).
The electrical resistivity of the samples was mea-
sured by applying a four-terminal measurement method
5,6. A constant DC-current source was used to force a
current of ~1 μA through the outer contacts of the struc-
ture. The voltage drop was measured with an electronic
voltmeter (FLUKE 289) on the inner contacts. In this
way the contact resistance was completely eliminated
and the results represent the pure resistance of the layer
between the inner contacts. The specific electrical resis-
tivity  of the printed layer with a thickness d was calcu-
lated using the measured voltage drop Vx and known cur-
rent I according to the equation:
 = ⋅ ⋅
V
I
W
L
d
x
(1)
where W and L denote the width and the length of the
measured strip between the inner contacts.
The adhesion of printed layers on both substrates was
evaluated using the standard cross-cut test, applying the
Byko-cut universal inspection guage (Byk-Gardner
Instruments, Germany). This method evaluates the
coating resistance to separation from the substrate when
a right-angle lattice pattern is cut up to the substrate. The
micrographs of the prepared samples are then rated into
classes with values 0–5, according to ISO 2409:1997.
3 RESULTS AND DISCUSSION
3.1 Print quality
The surfaces of all the printed lines clearly show
silver flakes, which are not completely covered by the
binder (Figure 1). This is the consequence of the very
high solids content in the wet paint.
The dried, printed conductive lines were analyzed
using line-profile measurements and an SEM image
analysis. They are well separated. Surface-profile
measurements show the different surface roughnesses of
the two applied paper substrates and the similar
roughnesses of the lines printed on them (Figure 2).
Such results were obtained on all the prepared samples.
These profiles were applied to determine the average
thickness of the dry printed layers. In general, thinner
layers were measured on the gloss-coated paper and
thicker on the clear matt film. The wet-to-wet over-
M. @VEGLI^ et al.: SCREEN-PRINTED ELECTRICALLY CONDUCTIVE FUNCTIONALITIES IN PAPER SUBSTRATES
628 Materiali in tehnologije / Materials and technology 45 (2011) 6, 627–632
Figure 2: Surface profile of four parallel lines printed on gloss-coated
paper (a) and clear matt film (b) as obtained by the profilometer. The
lines are single-layer prints that were cured for 9 min. The average
thickness is 9.1 μm and 15.5 μm, respectively. It was determined as the
distance between the average lines of the substrate and top of the
printed lines, determined on the marked regions.
Slika 2: Povr{inski profil {tirih vzporednih ~rt, ki so bile natisnjene na
sijajnem premazanem papirju (a) in na hrapavi foliji (b). Meritve so
bile narejene na profilometeru. ^rte so bile natisnjene z enkratnim
prehodom in su{ene 9 min. Povpre~na debelina prikazanih linij je 9.1
μm (a) in 15.5 μm (b). Dolo~ena je kot razdalja med podlago in vrhom
tiskanih linij; polo`aj vsake od teh je dolo~en kot povpre~na lega
obarvanih podro~jih.
Figure 1: SEM micrograph of a typical front surface of printed
conductive line. The silver flakes are not fully covered by the binder
Slika 1: SEM-posnetek povr{ine tiskane prevodne linije. Srebrne
luske niso popolnoma zakrite z vezivom
printed layers are slightly thicker than the single-printed,
but their thickness is much less than doubled. All the
layers become thinner when longer curing was applied.
The influence of curing time and overprinting is small
(Figure 3).
The next important property is the width of the
printed lines, which was evaluated with SEM and optical
micrographs. The edges of the lines are not perfectly
straight. A considerable amount of wicking was detected
on the optical micrographs (Figure 4a). SEM micro-
graphs show that such edges could lack conductive
particles; some regions could also remain without any
electrical functionality (Figure 4b). Therefore, the
contribution of this region to the electrical conductivity
is not the same as it is in the bulk of the printed shape.
M. @VEGLI^ et al.: SCREEN-PRINTED ELECTRICALLY CONDUCTIVE FUNCTIONALITIES IN PAPER SUBSTRATES
Materiali in tehnologije / Materials and technology 45 (2011) 6, 627–632 629
Figure 6: SEM micrograph of analyzed strip printed on gloss-coated
paper (a) and clear matt film (b)
Slika 6: SEM-posnetek analizirane ~rte, natisnjene na sijajnem pre-
mazanem papirju (a) in hrapavi foliji (b)
Figure 4: The edges of a printed line as observed by optical (a) and
SEM micrographs (b). The width of the lines was evaluated from
optical micrographs, taking into account the inner tangents. The
printing squeegee was moved perpendicularly to the printed strip.
Slika 4: Robovi tiskanih ~rt, posneti z opti~nim (a) in elektronskim
mikroskopom (b). [irino ~rt smo dolo~ili med premicama, ki ozna~u-
jeta notranjost robov. Pri tisku se rakelj giblje v smeri od zgoraj
navzdol.
Figure 5: The fully functional width (central part of strips, broken
lines) and the corresponding complete width (wicking area taking into
account, full line) of strips printed by single- (open signs) and double
(solid signs) wet-to-wet overprinting on gloss coated paper (triangles)
and clear matt film (circles) as a function of the curing time.
Slika 5: [irina elektri~no funkcionalnega dela ~rt (osrednji del, ~rtka-
ne ~rte) in celotna {irina (z upo{tevanjem nagubanih robov, polne ~rte)
v odvisnosti od ~asa su{enja. ^rte so bile natisnjene z enkratnim
(prazni znaki) in dvakratnim prehodom raklja (polni znaki) na sijaj-
nem premazanem papirju (trikotniki) in hrapavi foliji (krogi).
Figure 3: The thickness of strips single- (open signs) and double
(solid signs) wet-to-wet overprinted on gloss coated paper (triangles)
and clear matt film (circles) as a function of the curing time
Slika 3: Debelina ~rt natisnjenih z enkratnim (prazni znaki) in dva-
kratnim prehodom raklja (polni znaki) na sijajnem premazanem
papirju (trikotniki) in na hrapavi foliji (krogi) v odvisnosti od ~asa
su{enja
The width of the wicking area was evaluated using two
straight lines, limiting it on both sides of printed strip on
the optical micrograph. Electrically, the fully functional
width of the strips was determined between the inner
tangents, whereas the rest was considered as the amount
of wicking (Figure 4a). An about 10 % narrower
wicking area was obtained on the top of the horizontally
printed strips than on the opposite side. This is the
consequence of the moving direction of the squeegee
during the printing. The width of the central line (i.e., the
fully functional shape) is larger on the gloss-coated
paper and smaller on the clear matt film. It tends to
diminish with curing time; however, the effect is
observable on single-printed samples having a
gloss-coated paper substrate, but could be neglected
elsewhere (Figure 5). The complete width of the strips is
the sum of the fully functional central part and the
wicking area on both sides. It is about the corresponding
width on the print form (i.e., 500 μm), being about 10 %
broader when the shape was double printed on gloss-
coated paper and could be up to about 10 % narrower
elsewhere. In general, narrower lines were obtained on
the gloss-coated paper and wider on the clear matt film
(Figure 5).
A similar analysis was also performed on vertically
printed strips; narrower lines were obtained in this case.
The wicking area strongly depends on the direction of
the printed shape with respect to the moving of the
squeegee. This effect is known within the graphics
industry 8. The width of all the printed shapes is
systematically broader when directed perpendicularly to
the moving direction of the squeegee.
The dependence of the printed strips on the applied
substrate can be attributed to the different roughness,
porosity and the ability of the ink to diffuse into the
substrate. The clear matt film has a very porous and
rough surface, whereas the gloss-coated paper has much
M. @VEGLI^ et al.: SCREEN-PRINTED ELECTRICALLY CONDUCTIVE FUNCTIONALITIES IN PAPER SUBSTRATES
630 Materiali in tehnologije / Materials and technology 45 (2011) 6, 627–632
Figure 8: The cross-cut test of single-layer printed on gloss-coated
paper (a, adhesion ISO class 1) and clear matt film (b, adhesion ISO
class 3)
Slika 8: Metoda kri`nega reza enoslojne plasti, tiskane na sijajnem
premazanem papirju (a, ISO adhezija razreda 1) in hrapavi foliji (b,
ISO adhezija razreda 3)
Figure 9: Specific resistivity of single- (open signs) and double-
printed layers (solid signs) printed on clear matt film (circles) and
gloss-coated paper (triangles)
Slika 9: Specifi~na upornost eno- (prazni znaki) in dvoplastnih linij
(polni znaki), tiskanih na hrapavi foliji (krogi) in na sijajnem prema-
zanem papirju (trikotniki)
Figure 7: SEM micrographs of cross-sections of samples printed on
gloss-coated paper (a) and clear matt film (b)
Slika7: SEM-posnetek prereza vzorcev, natisnjenih na sijajnem pre-
mazanem papirju (a) in hrapavi foliji (b)
smaller pores and a rather smooth surface (Figure 6).
The smooth surface of the gloss-coated paper enables
good orientation of the flakes, which gives thinner layers
than the rough, clear matt film. The porosity of the
substrates, together with the wettability (which was not
evaluated here), influences the width of the printed
strips. They are narrower on the gloss-coated paper and
wider on the surface of the clear matt film. The addi-
tional difference between the prints on both substrates
could be seen on the cross-sections. The layers printed
on clear matt film do not adhere properly to the substrate
(Figure 7b), while no such effect was observed on
gloss-coated paper (Figure 7a). Most likely, this crack-
like feature was created due to the large diffusion of
liquid constituents of the ink through the pores of the
substrate. Because the flakes are rather large, the poly-
mer binder could not permeate sufficiently from the
above positions, and therefore the layer and the substrate
are not entirely merged. No such effects were observed
on the gloss-coated paper (Figure 7). The existence of a
partial separation between the printed layer and the
substrate gives rise to poor adhesion. The effect was
evaluated by cross-cut tests (Figure 8). The gloss-coated
paper has a ISO class 1 (good adhesion, cross-cut area
not significantly greater than 5 % is affected), while the
clear matt film ISO class 3 (poorer adhesion, cross-cut
area between 15 % and 35 % is affected).
3.2 Electrical resistivity
The specific resistivity of all the prepared prints is
shown in Figure 9. Higher values were obtained for the
conductive lines on the clear matt film and smaller on the
gloss-coated paper. In general, single-printed layers have
a higher resistivity and the double-printed, a lower. The
specific resistivity also depends on the curing time: up to
9 min of curing, the resistivity diminishes and then this
effect becomes smaller.
The electrical conductivity of composites having
conductive particles in a dielectric medium (as with our
silver-based printing ink) depends, among other para-
meters, on the concentration of particles and their
orientation within the film 10. It is well known that flaky
particles orient preferably parallel to the substrate 11. The
distribution of silver flakes is influenced by the printing,
overprinting and drying, giving rise to arrange, rearrange
or distribute themselves evenly across the layer. The
flakes are well oriented on the gloss-coated paper and
more random on the clear matt film, according to their
different surface roughness. This effect could explain the
larger resistivity of the strips printed on clear matt film.
During the curing process the volume of liquid compo-
nents diminishes due to the evaporation of the solvents,
due to the spreading of the strips and the penetration into
the substrate; because of that the concentration of the
flakes increases, which diminishes the specific resisti-
vity. The last two processes are intensified on double-
printed layers, which may explain the different specific
resistivity of single- and double-printed strips on the
same substrate: double-printed layers have a higher con-
centration of flakes.
4 CONCLUSIONS
A detailed analysis of screen-printed, silver-based,
conductive lines with a target width of 500 μm for
applications in printed electronics is shown here. The
influence of the substrate, curing time and wet-to-wet
overprinting were considered. While profilometer,
optical microscope and SEM images do not give the
same information about the topography of prints, the
specific resistivity of printed layers has to be determined
very carefully.
Two flexible substrates were applied, one with a
smooth surface and the other with a rough and highly
porous surface. The smooth substrate provides a well-
merged interface between the layer and the substrate,
whereas the porous substrate is not in full contact with
the layer, giving rise to poor adhesion. Thinner and
wider lines were obtained on the smooth surface, but
thicker and narrower lines were seen on the rough one.
All the lines have some wicking area where the content
of the functional particles is, in general, much smaller
than in its central part. Many of such features cannot
contribute to the functional width of the printed line. The
second wet-to-wet overprinted layer does not double the
layer thickness but increases the width of the lines,
especially the area with no proper functionality. These
effects are stronger on the rough and porous substrate.
Longer curing gives somewhat thinner and narrower
strips.
The resistivity of the layers on smooth paper is
smaller than that printed on the rough and porous
surface. Wet-to-wet overprinting gives a smaller
resistivity. It diminishes with a curing time up to 9 min.
The specific resistivity depends on the concentration of
flakes and the microstructure of printed line; it is
influenced by the drying process, the surface roughness
of the substrate and its porosity.
ACKNOWLEDGEMENTS
This research is supported by Slovenian Research
Agency (Project No. L02-1097-0104-08). Ma{a @vegli~
acknowledges the Slovenian Research Agency for young
researchers’ support.
5 REFERENCES
1 Berggren, M., Nilsson, D., Robinson, N. D., Organic materials for
printed electronics, Nature Materials, 2 (2007), 3–5
2 Li, Y., Wong, C. P., Recent advances of conductive adhesives as a
lead-free alternative in electronic packaging: materials, processing,
reliability and applications, Material Science and Engineering, 51
(2006), 1–35
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3 Strûmpler, R., Glatz-Reichenbach, J., Conducting polymer compo-
sites, Journal of Electroceramics, 3 (1999), 329–346
4 Park, B. K., Kim, D., Jeong, S., Moon, J., Kim, J. S., Direct writing
of copper conductive patterns by ink-jet writing, Thin Solid Films,
515 (2007), 7706–7711
5 Klanj{ek Gunde, M., Hauptman, N., Ma~ek M., Electrical properties
of thin epoxy-based polymer layers filled with n-carbon black
particles, Proceedings of SPIE 6882 (2008) 68820M-1-68820M-8.
6 N. Hauptman, M. @vegli~, M. Ma~ek, M. Klanj{ek Gunde, Carbon
based conductive photoresist, J. Mater. Sci., 44 (2009), 4625–4632
7 Pudas, M., Halonen, N., Granat, P., Vähäkangas J., Gravure printing
of conductive polymer inks on flexible substrates, Progress in
Organic Coatings, 54 (2005), 310–316
8 H. Kipphan, Handbook of Print Media, Berlin, Heidelberg: Springer
2001
9 M. Pudas, N. Halonen, P. Granat, J. Vähäkangas, Gravure printing of
conductive polymer inks on flexible substrates, Progress in Organic
Coatings, 54 (2005), 310–316
10 I. Balberg, A comprehensive picture of the electrical phenomena in
carbon black-polymer composites, Carbon, 40 (2002), 139–143
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spectra of metal-effect coatings, Applied Spectroscopy, 57 (2003),
1266–1272
M. @VEGLI^ et al.: SCREEN-PRINTED ELECTRICALLY CONDUCTIVE FUNCTIONALITIES IN PAPER SUBSTRATES
632 Materiali in tehnologije / Materials and technology 45 (2011) 6, 627–632
M. J. FREIRÍA GÁNDARA: RECENT GROWING DEMAND FOR MAGNESIUM IN THE AUTOMOTIVE INDUSTRY
RECENT GROWING DEMAND FOR MAGNESIUM IN
THE AUTOMOTIVE INDUSTRY
RAST POVPRA[EVANJA PO MAGNEZIJU V AVTOMOBILSKI
INDUSTRIJI
María Josefa Freiría Gándara
Xunta de Galicia, Consellería de Educación e Ordenación Universitaria, 15704 Santiago de Compostela, La Coruña, Spain
josefa.freiria@yahoo.es
Prejem rokopisa – received: 2011-04-08; sprejem za objavo – accepted for publication: 2011-06-17
This article summarizes the importance of magnesium and magnesium alloys in the automotive industry. The resources and
properties for magnesium, as well as for magnesium alloys, in manufacturing are concisely treated, taking the SF6 emissions
from magnesium production into account. Moreover, the possibilities of recycling magnesium and magnesium alloys are
considered, ending with the expectations and problems in the wider application of magnesium in motor vehicles.
Keywords: Automotive industry, Mg Alloys, emission SF6, recycling, motor vehicle.
V ~lanku je poudarjen pomen magnezija in magnezijevih zlitin v avtomobilski industriji. Viri in lastnosti magnezija in njegovih
zlitin so na kratko predstavljeni z upo{tevanjem emisije SF6 pri proizvodnji magnezija. Analizirane so tudi mo`nosti recikliranja
magnezija in magnezijevih zlitin. Predstavljena so tudi pri~akovanja in problemi ve~je uporabe magnezija v motornih vozilih.
Klju~ne besede: avtomobilska industrija, Mg-zlitine, emisija SF6, recikliranje, motorna vozila
1 INTRODUCTION
Although a small amount of magnesium has been
used in automobiles for many years, its low density and
the constant search for weight savings are encouraging
subjects to evaluate more potential applications. The ease
with which die castings can be produced makes it a fa-
voured manufacturing route for most applications 1. Cur-
rent applications include seat frames, transmission sys-
tem casings, air-bag housings and lock bodies. Table 1
summarises the benefits of using magnesium die castings
for seat frames.
Table 1: Benefits of using magnesium die castings for seat frames
Tabela 1: Prednosti pri uporabi tla~nih ulitkov magnezija za okvirje
sede`ev
Material choice:
• Lower specific weight than other options
• Better elongation than other die-casting metals
• 20–30 % shorter cycle times than aluminium die casting
• Longer die life (about double) than aluminium die cast-
ing
• ability to produce thinner walls than aluminium die cast-
ing
Economic choice:
• Magnesium price stability
• Investment cost reduction in comparison to current seats
Pure magnesium is about one-third lighter than alu-
minium, and two-thirds lighter than steel. A lighter
weight translates into greater fuel efficiency, making
magnesium-alloy parts very attractive to the auto indus-
try. And these lighter parts come with good deformation
properties, giving the products good dent and impact re-
sistance, as well as fatigue resistance. The alloys also
display good high-speed machinability and good thermal
and electrical conductivity 2.
Magnesium is a silvery-white metal that is princi-
pally used as an alloying element for several non-ferrous
metals (Al, Zn, Pb, etc.). Magnesium is among the light-
est of all the metals and also the sixth most abundant on
earth. The variety of applications of magnesium alloys in
the automotive industry, aerospace engineering, chemical
industry, etc. contribute to a rapid increase in the produc-
tion of magnesium in the world 3.
The production of magnesium in the world increased
from 20 000 t per year in 1937 up to 400 000 t per year
in year 2000, of which China produced 170 000 t of
magnesium in 2000. Magnesium alloys are predicted to
continue to grow in popularity (about 15 % per year in
the automotive industry alone), but the world’s supply of
magnesium, like every other natural resource, will be not
Materiali in tehnologije / Materials and technology 45 (2011) 6, 633–637 633
Table 2: US consumption of primary magnesium in 2007 by use
Tabela 2: Poraba proizvodov iz primarnega magnezija v ZDA v letu
2007
USE kt
Magnesium casting 25.9
Magnesium wrought products 2.7
Alloying element with aluminium (specially in
aluminium cans)
28.9
Iron and steel desulfurization 9.3
Others 5.4
Total 72.3
UDK 629.1-46:669.721 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 45(6)633(2011)
be unlimited forever. The solution is, logically, to recy-
cle. Especially since anywhere between 30 % to 50 % of
the metal handled by die casters ends up as scrap. For the
end of life, a consequential approach is used: recycling
the metal will significantly avoid the need to produce
primary metal. To optimize the recycling of magnesium,
an understanding of the current market of primary mag-
nesium is essential 4. Table 2 is based on a publication of
the US Geological Survey.
2 RESOURCES AND PROPERTIES OF PURE MG
METAL. MANUFACTURING
The abundance of Mg in the Earth is considered to be
the 4th highest metal, following iron, aluminium and sili-
con. The raw ores of Mg are dolomite (MgCO3. CaCO3)
and magnesite (MgCO3), and Mg is the second most
abundant metal in seawater, following sodium. There-
fore, it can be said that magnesium is an almost inex-
haustible resource, and is distributed all over the world.
Magnesium is the lightest of all metals in practical
use, and has a density (1.74 g/cm3) of about 2/3 that of
aluminium and 1/4 that of iron.
On the other hand, magnesium has shortcomings,
such as insufficient strength, elongation and heat resis-
tance, and a corrosion propensity. Its readiness to cor-
rode has been found to be due to trace content of metals
such as iron (Fe), nickel (Ni) and copper (Cu). The prob-
lem of corrosion has to be solved if the purity of the Mg
is improved. However, its electrochemical potential indi-
cates that magnesium will corrode by contact corrosion
whenever it is in contact with any other metal. Therefore,
magnesium is generally surface-treated before it is used.
The methods for manufacturing Mg can roughly be
divided into electrolysis and thermal reduction. The elec-
trolysis method involves extracting magnesium chloride
from Mg ores and then reducing magnesium to its metal-
lic form by electrolysis. Thermal reduction involves ex-
tracting magnesium oxide from Mg ores, adding a reduc-
ing agent such as ferro-silicon to it, and refining the
resulting material by heating it to a high temperature un-
der reduced pressure.
3 PROPERTIES OF MAGNESIUM IMPROVED
WITH ALLOYING
The development of magnesium-alloy products has a
long history that dates from 1945. Research has been
conducted on the manufacture of various products, such
as office goods, agricultural machines and tools, tele-
communications equipment and sporting goods. Mg al-
loys have not yet been used as light structural materials
for aircraft. However, they have been used for the gear-
box housings of helicopters and other aircraft because
they are good vibration dampers, a characteristic that has
also brought them into use in the steering wheel cores of
motor vehicles.
Alloying means altering a pure metal by melting it
and adding other elements to it. This method is applied
to almost all metals. Alloying magnesium improves its
strength, heat resistance and creep resistance (creep is
defined as deformation at a high temperature and under
load). For example, AZ-based Mg alloys are well known
materials produced by adding aluminium (Al) and zinc
(Zn) to pure Mg. The appropriate amounts of additives
may improve the strength, castability, workability, corro-
sion resistance and weldability of these alloys in a
well-balanced way 5.
4 SF6 EMISSIONS FROM MAGNESIUM
PRODUCTION
The use of Mg alloys has recently increased. The sur-
face-treatment techniques used to provide highly corro-
sion resistant Mg alloy products have already advanced
on the same level as the die casting methods of some car-
bon steel plates and aluminium alloy products 6. To put
Mg alloy products to practical use, however, it is neces-
sary to solve the critical problem that magnesium has a
high activity in the presence of oxygen. For Mg, it is es-
sential to prevent the formation of products of reaction
with oxygen in the air. Currently, this is mainly accom-
plished with sulphur hexafloride (SF6) gas, but is a po-
tential greenhouse gas so alternatives that do not use SF6
are now under investigation. The magnesium metal pro-
duction and casting industry uses sulphur hexafluoride
(SF6) as a cover gas to prevent the violent oxidation of
molten magnesium in the presence of air. SF6 is a colour-
less, odourless, non-toxic, and non-flammable gas. The
industry adopted the use of SF6 to replace sulphur diox-
ide (SO2), which is toxic and requires careful handling,
to protect worker safety.
Historically, more than half of SF6 emissions from
the US magnesium industry have come from the primary
magnesium industry, and the magnesium recycling in-
dustry, for the most part, continues to employ sulphur di-
oxide as a cover gas.
In 1999, EPA began the voluntary SF6 Emission Re-
duction Partnership for the Magnesium Industry. SF6 was
introduced to replace SO2, which corrodes casting equip-
ment also. However, safer SO2 handling procedures and
the relatively low cost of SO2 compared to SF6 makes
SO2 more attractive 7.
5 MAGNESIUM TODAY
The magnesium industry is now experiencing a time
of change with the die-casting segment characterized by
rapid growth, both in the European and the North Ameri-
can markets. On the supply side, newcomers have started
production, several projects are under consideration or at
the pilot stage, while other suppliers have withdrawn
from the business 8.
M. J. FREIRÍA GÁNDARA: RECENT GROWING DEMAND FOR MAGNESIUM IN THE AUTOMOTIVE INDUSTRY
634 Materiali in tehnologije / Materials and technology 45 (2011) 6, 633–637
In today’s marketplace, magnesium die-casting
means the production of parts in compliance with the
stringent requirements of high-purity alloy standards.
This is the basis for the successful applications of
die-cast parts for use in corrosive environments without
complicated and cost-adding surface treatments. It is a
paradox that magnesium die-cast alloys are melted in
steel crucibles, when we know that iron is an impurity
with a strong negative influence on corrosion. This can
only be done when the whole production sequence, in-
cluding the processing of ingots, is performed with close
attention to the thermodynamic principles involved in
keeping iron at an acceptable level throughout the pro-
cess.
A number of applications, especially in the automo-
tive field, require the ability to absorb energy without the
formation of a fracture. Typical examples are provided
by steering-wheel cores, instrument panels/cross-car
beams, seat parts and door parts. These are typical exam-
ples where material properties, processing and design in-
teract strongly in determining the properties of the prod-
uct. The safe handling of magnesium requires an
understanding of the basic processes that can cause haz-
ardous events. The most common reactions involved are
the following:
1) Burning/oxidation
2Mg + O2 = 2MgO
2) Rapid evaporation (expansion) of water entrapped by
liquid magnesium
Mg(liq.) + H2O(liq.) = H2O(vap.) + Mg
3) Water reaction/hydrogen explosion
Mg + H2O = MgO + H2
2H2 + O2 = 2H2O
4) Thermite reaction
3Mg + Fe2O3 = 3MgO + 2Fe
5) Silica reaction
2Mg + SiO2 = 2MgO + Si
There is a growing emphasis on environmental as-
pects when materials are selected. Magnesium, like alu-
minium, is an energy-intensive material to produce, and
it is not obvious that use of this material will be benefi-
cial. With a full analysis of the environmental impact of
production, use and recycling, a life-cycle assessment
(LCA) can be performed. Hydro magnesium has issued a
comprehensive life cycle inventory (LCI) for the produc-
tion of magnesium die-casting alloys. This inventory has
been used to make a comparative analysis of the parts
produced in die-cast magnesium alloys and other alloys.
It follows that recyclability is a necessary prerequisite to
release the full environmental benefits of using magne-
sium. Another important factor is to reduce or eliminate
the consumption of the potent greenhouse gas SF6. Re-
cent research has shown that applying a dilute mixture of
SO2 in dry air is a viable method for replacing SF6. SO2
must be added as a 0.5–1.5 % mixture in dry air by using
a well-designed gas distribution system supplied from a
special mixing unit.
The rapid growth of magnesium alloy die-casting has
triggered numerous developments of post-casting treat-
ments. Equipment is readily available for trimming, vi-
bratory finishing, blasting and machining. New, environ-
mentally friendly conversion coatings have been
developed, as well as improved painting and anodizing
processes.
6 MAGNESIUM ALLOYS
Due to their low density (1.8 g/cm3), magnesium al-
loys offer distinct advantages for weight saving in auto-
motive applications 9. It is fair to say that the 1990s have
seen a growth of magnesium applications, among which
are found steering-column assemblies, steering wheels,
instrument panels, seat frames, valve covers and even a
manual transmission-case application. The need for fur-
ther weight reduction is extending the future use of mag-
nesium to critical components such as transmission and
engine parts. Magnesium faces a challenge in meeting
the performance requirements of these components, for
elevated-temperature (150 °C) strength and creep resis-
tance, as well as adequate corrosion resistance. The most
economic use of Mg in the automotive industry at the
moment is in die-cast applications and this is because of
the high productivity of the die-casting process, which
upsets the cost of the magnesium metal.
The 1990s have seen renewed activity in the develop-
ment of elevated-temperature Mg die-casting alloys. In
1994–1997 there have been patent applications on
rare-earth and calcium containing alloys. These alloy
systems have very good creep properties, such as
creep-resistance and bolt-load retention, but contain the
expensive rare-earth additions and their die-castability is
not known. Also, in 1994-95 a cost-effective Mg-Al-Ca
alloy system was developed and a patent application was
filed.
There is presently an ongoing worldwide effort to de-
velop more optimum elevated-temperature Mg alloys.
Work is underway on the development of new alloys
with a potential application in automotive parts requiring
elevated temperature performance, such as engine com-
ponents and automatic transmission cases. The commer-
cial success of these alloys will require Mg producers to
focus their efforts and address issues related to cost and
recyclability. In collaboration with part producers and
end users they will also need to address die-castability
and manufaturability.
7 POSSIBILITIES FOR THE RECYCLING OF
MAGNESIUM AND MAGNESIUM ALLOYS
Simultaneously with the increasing applications and
consumption of magnesium, significant quantities of new
scrap are generated; scrap from production as well as
postconsumer scrap. As a result of this anticipated in-
crease in new scrap generation, companies are planning
new magnesium recycling plants or they are expanding
M. J. FREIRÍA GÁNDARA: RECENT GROWING DEMAND FOR MAGNESIUM IN THE AUTOMOTIVE INDUSTRY
Materiali in tehnologije / Materials and technology 45 (2011) 6, 633–637 635
existing capacity. The principal long-term effect is that
after an automobile is scrapped the magnesium-contain-
ing parts may be removed from the automobile and recy-
cled. These additional magnesium-containing parts
would result in additional quantities of old scrap as a
source of supply. The projected increase in the use of
magnesium in this application has prompted developed
countries such as Germany, Japan, Great Britain, USA
and Canada to install additional magnesium recycling ca-
pacity. In USA in 1998 the recycling efficiency rate for
magnesium was estimated to be 33 %.
New magnesium-based scrap is typically categorized
into one of the four types. Type I is high-grade scrap,
generally materials such as gates, runners and drippings
from die-casting operations that is uncontaminated with
oils. Types II, III and IV are lower-graded materials.
Type II is oil-contaminated scrap, type III is dross from
magnesium-processing operations, and type IV is chips
and fines. Old magnesium-based scrap or postconsumer
scrap consists of such materials as automotive parts, heli-
copter parts, lawnmower decks, used tools and the like.
This scrap is sold to scrap processors. In addition to
magnesium-based scrap, significant quantities of magne-
sium are contained in aluminium alloys that can also be
recycled. In particular, magnesium has a lower specific
heat and a lower melting point than other metals. This
gives the advantage of using less energy in recycling,
with recycled Mg requiring as little as about 4 % of the
energy required to manufacture new material. At present,
however, recycling procedures are still not fully devel-
oped. For example, an experimental investigation of the
processing of non oil-contaminated metal scrap based on
magnesium and its alloys was carried out. Experimental
investigations were conducted on a laboratory scale and
the results were verified on an industrial scale. The in-
vestigations show that: processing of this kind of scrap is
possible with a metal-extraction efficiency rate in range
45–90 %, depending on the quality of the scrap; and for
the purpose of achieving satisfactory techno-economical
effects it is necessary to have suitable processing tech-
nology, which includes the preparation, metallurgical
processing, smelting and refining. An electrolytic pro-
cess to recycle low-grade and post-consumer magnesium
scrap was developed to recover magnesium from magne-
sium oxide, unlike the traditional electrolytic process
that uses magnesium chloride as a feed material 10.
8 THE RECYCLING OF MAGNESIUM
Magnesium scrap is being melted and refined under
strict control. After the removal of oxides and compo-
sitional adjustments, magnesium alloys of at least the
same quality as primary metal are cast into ingots.
The magnesium die-casting industry has grown sig-
nificantly over recent years and this growth is projected
to continue with automotive applications leading the
way. The current consumption of magnesium is about 2
kg per vehicle worldwide, and over the next 20 years, the
automotive industry could use almost 100 kg per vehicle,
or more than 50 times the current demand. This leads to
significant amounts of magnesium scrap. Depending on
the die-casting process used, and the design of the part,
between 80 % to 100 % of the net weight of the casting
can be scrap. To realize this growth, magnesium has to
be economically and environmentally acceptable. All op-
tions for reusing pre- and post-consumer magnesium
scrap in alternative markets, or for recycling, need to be
evaluated. The growing demand for magnesium alloys in
the automobile industry necessitates increasing recycling
capacities. In this respect, the recycling process is today
still basically related to the clean casting returns. How-
ever, there is an ever greater need to recycle other resid-
ual magnesium materials as well. The reuse or recycling
of these residual materials continues to be a problem, in
particular with regard to environmental issues. The
re-melting of painted casting returns results in consider-
able quantities of dioxin. These must be recorded quanti-
tatively and, using an appropriate filter, must be removed
from the waste gases down to a content of 0.1 mg/m3.
This demands a relatively large investment in filter tech-
nology.
Today, the "dross" arising in the die-casting foundries
is also recycled, i.e., the more or less salt-free magne-
sium scrapings that occur when the crucible surface is
skimmed. The chips arising in the production process to-
day are contaminated with either oil or oil-water emul-
sions. Even with long draining times, it is not possible to
reduce the oil content to below 10 %. In this form, the
chips cannot be re-melted using any of the melting tech-
niques applied so far. They burn (to a greater or lesser
extent), producing black oil-smoke clouds in the melting
crucible, resulting in extreme emissions of dioxins, total
C, HCl and Mg oxide, as well as scrapings that have to
be disposed of in an expensive process. This is a very
high-cost method of destroying magnesium chips. A
technique was developed, it consists of pressing the mag-
nesium chips in a hot condition at > 300 °C, which not
only removes the oil thermally, but also causes the chips
to cake in such a way that a re-melting process can be
carried out without any major loss during red heat. Until
now, this technique has only been applied rarely, and the
quantities to be pressed per unit of time are relatively
small.
In addition to protecting the environment, the main
objective, of course, is to manufacture a clean alloy,
which can be reused in the automobile industry and, if
possible, has a high-purity (HP) quality. To achieve this
it requires a type-specific separation and collection of the
casting returns at the die caster’s plant, and a fault-free
handling/transport at the disposing/recycling company. If
these prerequisites are satisfied, the recycling process
will result in an alloy that complies with the standards.
Today we distinguish between four different recycling
categories:
1) "in-cell" recycling
2) "in-house" recycling
3) End of Life Vehicle (ELV) recycling
4) industrial (external) recycling/service.
M. J. FREIRÍA GÁNDARA: RECENT GROWING DEMAND FOR MAGNESIUM IN THE AUTOMOTIVE INDUSTRY
636 Materiali in tehnologije / Materials and technology 45 (2011) 6, 633–637
Depending on the structure of the installation, the an-
nual capacities of the recycling plants vary between
2 000 t per year and 10 000 t per year.
"In-cell" recycling takes place directly at the plant of
the die-caster, who immediately puts the clean, initial
castings into the melting crucible again. "In-house" recy-
cling may become interesting for die-casters with high
throughput quantities. ELV recycling still poses a chal-
lenge. By 2015, automobile manufacturers will be forced
to reuse 95 percent of the materials used in the car and,
what is more, in the same area of applications.
The traditional technology is melting and refining
with salt, also called flux refining. The casting returns
are melted down in an open steel crucible (heated by gas
or electricity) while salt is added.
The Salt Furnace Technology was developed, using
super-heated salt to melt the returns and cleaning the
melt by settling in a multi-chamber furnace. In the one-
furnace version of this technology, ingots are cast di-
rectly from the recycling furnace. There will be a con-
stant increase in the demand for the recycling of residual
magnesium materials. The statutory regulations will be
ever more oriented towards a closed-loop cycle in which
all the residual materials are, if at all possible, reused.
The utilization of larger quantities of "post-consumer
scrap" must be prepared and new applications must be
created for the use of "non-HP alloys".
9 EXPECTATIONS AND PROBLEMS IN THE
WIDER APPLICATION OF Mg TO MOTOR
VEHICLES
Passenger transport accounted for about 60 % of the
total energy consumption in the transport sector, and in
particular the use of private cars contributed significantly
to this consumption. To build a sustainable society in the
future it will be necessary to reduce the weight of the
structural materials used in transport equipment, espe-
cially private cars, both to conserve energy and to mini-
mize global warming 11.Today, magnesium alloys are
recognized alternatives to iron and aluminium for reduc-
ing the weight of structural elements 12. On the other
hand, motor vehicles tend to increase in weight as they
are given additional functions, such as safety devices and
electronic equipment. The challenge for the future is not
only to offset weight increases due to performance en-
hancements, but also to reduce the overall weight of mo-
tor vehicles. Conventional weight reduction technologies
have reduced the weight of motor vehicles by improving
their structural design and thinning steel materials by in-
creasing their strength. For the future, however, it is gen-
erally recognized that drastic changes in structural mate-
rials should be considered 13.
For passenger cars, the general rule is that about 86 %
of their total lifetime energy use (from the time of pro-
duction to the time of disuse or scrapping) is consumed
by carrying their own weight and persons around. In Eu-
rope, the regulation governing CO2 emissions from mo-
tor vehicles has been worked out, setting the standard
that CO2 emission shall not exceed 140 g/km in 2012 and
120 g/km in 2014. According to previous analysis, it will
be necessary to reduce the mass of vehicles by about
10 %. To attain such a great reduction in mass, it will
probably be necessary to replace steel with Mg alloys as
the structural material. For this reason, much attention is
now focused on Mg alloys as structural materials or parts
for motor vehicles.
To work towards a sustainable society, it is absolutely
necessary to develop energy-saving technologies that
contribute to the reduction of CO2 emissions. It is espe-
cially important to reduce the amount of energy that is
consumed simply to enable a vehicle to carry its own
weight around. Therefore, the weight reduction of trans-
port equipment is one of the most important technical
challenges. It is anticipated that activities will be acceler-
ated to develop and commercialize Mg alloys that con-
tribute to the weight reduction of structural materials for
transport equipment.
Although Mg alloys possess a variety of desirable
physical properties, including lightness, they have had a
limited range of applications because they also have per-
formance shortcomings, such as low strength, low heat
resistance and low corrosion resistance. In recent years,
however, advanced basic research on Mg alloys has en-
larged the range of applications 14.
10 REFERENCES
1 Zhi, D., Automotive Engineering, (1991), 1
2 Luo, A. A., Nyberg, E. A., Sadayappan, K., Shi, W. Magnesium front
end research and development: a Canada-China-USA collaboration.
Presented at Magnesium Technology Conference, New Orleans, LA,
USA, March 9–13, 2008
3 Liu, Z., Wang, Y.,Wang, Z., Li, F., Chinese Journal of Material Re-
search, (2000), 5
4 USGS, 2007 Annual Yearbook Magnesium (Advanced Release),
http://minerals.usgs.gov./
5 Wang, Q., Lu, Y., Zeng, X., Ding, W., Zhu, Y., Special Casting &
Nonferrous Alloys, (1999), 1
6 Fu, L., Automobile Technology & Material, (2006), 8
7 Norsk Hydro (2008). Progress to eliminate SF6 as a protective gas in
magnesium diecasting. Hydro Magnesium, Brussels, 2008
8 Shukun, M., Xiuming, W., Jinxiang, X. China’s magnesium industry
development status in 2007. Presented at 65th World Magnesium
Conference, Warsaw. Poland, May 18–20, 2008
9 Report of Investigation on the Technical Trend of Patent Applica-
tions in 2004: Automobile Weight reduction Technologies. Japan
Patent Office, March 2005
10 Gesing, A. J., Dubreuil, A. Recycling of post-consumer Mg scrap .
Presented at 65th Annual World Magnesium Conference, Warsaw,
Poland, May 18–20, 2008
11 Energy consumption Trend in Transport Sector: 2-2-1. Analysis on
private passenger cars’ contribution to the total energy consumption.
Home page provided by the Energy Conservation Center, Japan:
http://www.eccj.or.jp/transportation/2-1-1.html
12 Yan, Z., Hua, R., Automotive Engineering, (1994), 6
13 Nyberg, E.A., Luo, A. A., Sadayappan, K., Shi, W., Advanced Mate-
rials & Process, 166 (2008) 10, 35–37
14 Yan, Z., Automotive Engineering, (1993), 3
M. J. FREIRÍA GÁNDARA: RECENT GROWING DEMAND FOR MAGNESIUM IN THE AUTOMOTIVE INDUSTRY
Materiali in tehnologije / Materials and technology 45 (2011) 6, 633–637 637

L. GOSAR, D. DREV: CONTACT WITH CHLORINATED WATER: SELECTION OF THE APPROPRIATE STEEL
CONTACT WITH CHLORINATED WATER: SELECTION
OF THE APPROPRIATE STEEL
KONTAKT S KLORIRANO VODO: IZBOR USTREZNEGA JEKLA
Leon Gosar1,2, Darko Drev1,3
1Institute for Water of the Republic of Slovenia, Hajdrihova 28, 1000 Ljubljana, Slovenia
2University of Ljubljana, Faculty of Civil and Geodetic Engineering, Chair of Fluid Mechanics with Laboratory, Hajdrihova 28,
1000 Ljubljana, Slovenia
3University of Ljubljana, Faculty of Civil and Geodetic Engineering, Institute of Sanitary Engineering, Hajdrihova 28, 1000 Ljubljana,
Slovenia
leon.gosar@izvrs.si
Prejem rokopisa – received: 2011-02-02; sprejem za objavo – accepted for publication: 2011-08-18
In water-supply systems and public swimming pools, the presence of highly chlorinated water can result in very aggressive
corrosion. When choosing the appropriate type of steel, the extremely corrosive conditions that can occur are often forgotten.
Under these conditions, corrosion-protection layers (zinc layer, polymer colour) can be quickly removed, and stainless-steel
corrosion may occur as well. The high risk of the corrosion of galvanized steel pipes can also be caused by the improper
implementation of disinfection. With aggressive disinfectants, the zinc layer is quickly dissolved, which leads to corrosion of
the steel pipe. Therefore, we must select a particular type of stainless steel, thereby ensuring a much higher corrosion resistance
than normal stainless steel. It is very important that the selection of materials is determined at the design stage of the project. In
the contact of steel elements with swimming-pool water, in most cases extremely aggressive oscillations do not occur under
normal operating conditions, because the content of chlorine and other elements that affect corrosion are mostly low and stable.
However, even in these cases, from time to time, aggressive shocks may occur as a result of the cleaning treatment. Therefore,
with the selection of the appropriate stainless steel the corrosion risk can be prevented. The contribution of the paper is mainly
focused on experiences relating to the appropriate materials selection in the field of sanitary engineering.
Keywords: stainless steel, corrosion, sanitary engineering
Pri vodovodnih sistemih in javnih kopali{~ih se lahko pojavi mo~no klorirana voda, ki je korozijsko zelo agresivna. Ko izbiramo
ustrezno vrsto jekla, pogosto pozabimo na ekstremne korozijske razmere, ki se lahko pojavijo. V njih se lahko zelo hitro
odstranijo protikorozijski za{~itni sloji (cinkov sloj, polimerna barva), lahko pa se pojavi tudi korozija nerjavnega jekla. Veliko
nevarnost za korozijo pocinkanih jeklenih cevi lahko npr. povzro~imo z neustreznim izvajanjem dezinfekcije. Z agresivnimi
dezinfekcijskimi sredstvi hitro raztopino za{~itni sloj cinka in povzro~imo korozijo jeklene cevi. Zato je zelo pomembno, da se
`e v fazi projektiranja odlo~imo, katere materiale bomo izbrali ter na kak{en na~in se bo izvajala dezinfekcija. Pri izbiri jeklenih
elementov, ki so v kontaktu z bazensko vodo, v ve~ini primerov nimamo tako ekstremnih sprememb agresivnih razmer.
Vsebnost klora in drugih sestavin, ki vplivajo na korozijo, je v kopalni vodi ve~inoma vedno nizka in stabilna. Kljub temu pa se
lahko tudi v teh primerih pojavijo ob~asno agresivni {oki, ki lahko nastanejo v fazi ~i{~enja. Zato je treba z ustrezno izbiro
nerjavnega jekla prepre~iti nevarnost korozije. V prispevku je poudarek predvsem na izku{njah pri izbiri ustreznega materiala s
podro~ja sanitarnega in`enirstva.
Klju~ne besede: nerjavno jeklo, korozija, sanitarno in`enirstvo
1 INTRODUCTION
In this paper we focus on facilities, which in addition
to structural strength, also require sanitary adequacy. The
latter requirement is particularly important for the
selection of the stainless steel used in water-supply
systems, public baths, food-processing facilities,
kitchens, etc.1 In these facilities, aggressive corrosive
conditions in some parts of the water-supply systems
may appear. 2 In the case of swimming pools, due to the
presence of chlorinated water, the requirements for
additional corrosion resistance need to be met.3 Planners
and designers often do not pay enough attention to the
operating conditions in such facilities.
A particular case is the construction of a waterslide
structure, where it is crucial for corrosion-resistant
stainless steel to be selected to prevent endangering the
stability of the structure. In addition, the surfaces must
be completely smooth and free of corrosion effects,
which is a sanitary requirement. This problem can be
solved with the appropriate surface protection, 4 which
may be questionable in the junctions,5 if galvanic cells
occur. Conversely, in the case of some steel surfaces
(e.g., in food-processing facilities), contact with the
surface where zinc is gradually being dissolved is
prohibited. 6 Galvanized steel surfaces are resistant to
corrosion by first dissolving the less-noble metals, e.g.,
zinc, thereby protecting the steel against corrosion. This
method of corrosion protection is also used in
galvanized-steel water-system pipes. However, in this
case, corrosion protection is not only important for
maintaining the stability of the installation, but also to
ensure healthy drinking water.
The following types of steel corrosion can occur: 2
– Uniform corrosion: the metal loss is uniform on the
surface.
– Crevice corrosion: due to the specificity of the
electrochemical process in crevices where the
Materiali in tehnologije / Materials and technology 45 (2011) 6, 639–644 639
UDK 669.14.018.8:620.193 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 45(6)639(2011)
changes in pH occur in the medium, the corrosion is
progressing rapidly.
– Pitting corrosion: produced locally in the form of
notches.
– Intergranular corrosion: localized attack along the
grain boundaries, or immediately adjacent to the
grain boundaries, while the bulk of the grains
remains largely unaffected.
– Stress-corrosion cracking: unexpected sudden failure
of normally metals subjected to in a environment.
– Hydrogen brittleness: arises from the destructive
action of absorbed atomic hydrogen or hydrogen
protons in the crystal lattice.
– Erosion corrosion: degradation of the material
surface due to mechanical action, often by impinging
liquid, abrasion by a slurry, particles suspended in
fast-flowing liquid or gas, bubbles or droplets, s etc.
On steel surfaces, rust may occur as a result of
chemical reactions in the steel. The characteristic brown
colour appears in the presence of Fe+2 and Fe+3 iron
compounds. Iron enters the Fe+2 and Fe+3 forms due to
chemical reactions. However, these reactions occur more
rapidly if the steel is not sufficiently "noble", e.g., it does
not contain a sufficient amount of elements that are less
susceptible to corrosion (nickel, chromium, cobalt,
manganese, etc.).
The rate of the corrosion processes is affected by the
steel composition, temperature, atmosphere, and the
substances in contact with the steel construction (sheet
metal). The purpose of corrosion-protection coatings is
to reduce these processes to a minimum. If a connection
with the moisture and the oxygen from the air is
prevented, corrosion will progress very slowly. In this
case, the necessary liquid electrolyte and the atmosphere
that would enable an adequate decay rate of iron are
absent. Since in the transformation of elemental iron in
its compounds electrons are being emitted, this forms a
galvanic cell. For the galvanic-cell formation to occur,
both the electron donor and the recipient must be
present. There exist several chemical reactions, where
elemental iron passes over into its compounds. These
compounds are of a brownish colour and can be seen on
the outside as rust. 2
Fe  Fe2+ + 2e–O2 + 2 H2O + 4e  4 OH
Fe + H2O + ½ O2  Fe(OH)2
Fe · H2O + 1/2 H2O + 1/2 O2  Fe(OH)3
e.g. ½ (Fe2O3 · 3 H2O) ...
Purbaix 7 constructed the following potential/pH
diagram (Figure 1). Based on this diagram, one can
rapidly assess the corrosion resistance of various metals
in water at 25 °C. He defined the possible equilibrium
between the metal and H2O. A simplified Pourbaix
diagram indicates regions of "Immunity", "Corrosion"
and "Passivation" and is a guide to the stability of a
particular metal in a specific environment. Immunity
means that the metal is not attacked, while shows that a
general attack will occur. Passivation occurs when the
metal forms a stable coating of an oxide or other salt on
its surface.
In presence of chlorine ions, a high probability of the
formation of porous or pitting corrosion exists. This
result is the formation of small pits on the surface of the
metal. Pitting corrosion is often difficult to differentiate
from other, similar corrosive processes that may occur,
such as crevice corrosion or the dissolution of zinc (both
involve disparate mechanisms of corrosion). The critical
pitting potential depends on the concentration of Cl-ions,
the inhibitor potential of various anions in the solution,
their concentration (OH–, SO42–, NO3–, CrO42–), as well as
the temperature of the solution. In the case of swim-
ming-pool water, we have a significant content of Cl–
and SO42– ions and a relatively high water temperature.
While this is conducive for the formation of pitting
(porous) corrosion, a relevant microstructure of metals
and their alloys need to be present as well.
In the case of drinking-water systems in the food
industry, stainless steel should be selected. The same
holds for waterslide structures and other facilities in and
around swimming poles. However, it is important to note
that even with the use of stainless steel, corrosion can
still occur if:
– the stainless steel is not of adequate quality for the
given conditions of use,
– the stainless steel forms a galvanic cell,
– there is contact with highly corrosive chemicals,
– the stainless-steel surface was not properly treated
etc.
2 EXPERIMENTS AND RESULTS
We reviewed several examples of corrosion phenom-
ena in steel and stainless steel in the field of sanitary
L. GOSAR, D. DREV: CONTACT WITH CHLORINATED WATER: SELECTION OF THE APPROPRIATE STEEL
640 Materiali in tehnologije / Materials and technology 45 (2011) 6, 639–644
Figure 1: Purbaix diagram of iron 7
Slika 1: Purbaix-ov diagram za `elezo 7
engineering (water-supply network, baths, spas, kitchens,
etc). These were in most cases expert opinions and
related research inquiries into the causes of corrosion in
real-world structures. Due to the sensitive nature of these
studies, we are not able to present specific details about
the individual structures that were the focus of these
investigations. Nevertheless, we are able to present all
the relevant data and facts to support our findings and
conclusions.
2.1 Corrosion in an internal water-supply network
When planning a complex, internal water-supply
network, it is first necessary to determine the properties
of the available water and what the purpose of the local
water-distribution network will be. The designer must
decide which materials to use in order to facilitate
disinfection. If a specific type of disinfection is not
prescribed in advance (e.g., by law), the materials have
to be selected to allow the effective execution of a broad
range of disinfection procedures. When using galvanized
steel pipes and components made of stainless steel,
copper and brass, the designer should take into account
the possibility of galvanic cell formation and the removal
of protective zinc coatings. Zinc layers can easily be
dissolved by repeated disinfectant shocks. When using
plastic pipes, the possibility of heat shocks must be
considered. Heat shocks are often used to disinfect an
internal water-supply system when the presence of
Legionella bacteria is suspected. If heat shocks are
implemented, the water pipes need to withstand
temperatures up to 90 °C. In addition to water pipes,
thermal stability is also expected for the seal components
and other parts of the installed structure.
In the analyzed cases of the internal water-supply
network, we found that multiple chlorine shocks and
large concentrations of disinfectants (Table 1) resulted in
the dissolution of zinc coatings on steel pipes. Corrosion
also occurred inside stainless-steel water tanks (Figure
2).
Table 1: Disinfectants recommended by the National Institute of
Public Health of Slovenia for water supply systems.
Preglednica 1: Dezinfekcijska stredstva za vodovodne sisteme, ki jih
priporo~a IVZ
Disinfectant Quantity /mg/l
Recommended resources for
neutralization
Chlorine in
gaseous form
Cl2
50
(Cl form)
Sulphur dioxide (SO2)
Sodium thiosulphate (Na2S2O3)
Sodium
hypochlorite
NaClO
50
(Cl form)
Sulphur dioxide (SO2)
Sodium thiosulphate (Na2S2O3)
Calcium
hypochlorite
Ca(ClO)2
50
(Cl form)
Sulphur dioxide (SO2)
Sodium thiosulphate (Na2S2O3)
Table 2: Conversion table for stainless steel AISI 304 tags to the
European standard 8
Preglednica 2: Tabela za pretvorbo oznake nerjavnega jekla AISI 304
v evropski standard 8
Standard
(Europe) AISI
Chemical composition / %
C max Cr Ni Mn max
X5 CrNi 18-10 304 0,07 17-19,5 8-10,5 2
In one of the internal water-supply network cases, the
designer prescribed stainless steel AISI 304 (Table 2). In
the lists of stainless steels for use in chlorinated water
conditions, Deutsches Institut für Bautechnik recom-
mends stainless steel with the same chemical compo-
sition as AISI 304, no. 1.4301.8
Extensive corrosion on all the surfaces of the reser-
voir has shown (Figure 2) that it could not result from
conventional chlorine shocks alone, but instead required
the presence of much higher concentrations of chlorine.
This reservoir was constructed of stainless steel X5 CrNi
18-10 as recommended by the Deutsches Institut für
Bautechnik for structures in contact with pool water that
may be highly chlorinated. In the case of the water sup-
ply, disinfection with the recommended agents was em-
ployed to destroy the bacteria of the Legionella group
(Table 1). Since this was not successful, the disinfection
was repeated several times, each time with an increased
concentration of disinfectants. When it was determined
that the disinfectants were not appropriate, oxidizing
agents effective for the removal of biological deposits
and the destruction of Legionella bacteria were intro-
duced into the system. This resulted in a corroded inter-
nal water-supply system and poor quality of the drinking
water. However, by adding an oxidizing disinfectant
based on hydrogen peroxide and small amounts of col-
loidal silver particles, one can indeed achieve microbio-
logical water disinfection. However, this water will also
likely be organoleptically and chemically unsuitable 9
due to the resulting water-pipe corrosion.
2.2 Stainless-steel corrosion in a public swimming
pool
Swimming-pool water is always chlorinated to
prevent the development of adverse microorganisms.
Such water is very aggressive to metals and can easily
L. GOSAR, D. DREV: CONTACT WITH CHLORINATED WATER: SELECTION OF THE APPROPRIATE STEEL
Materiali in tehnologije / Materials and technology 45 (2011) 6, 639–644 641
Figure 2: Corrosion inside a cold water reservoir
Slika 2: Korozija v notranjosti rezervoarja hladne vode
cause corrosion. Swimming-pool water also typically
contains sulphates in addition to chlorine ions. In the
presence of organic matter, chlorine forms chlorinated
hydrocarbons, i.e., trihalomethanes, which are classified
as carcinogenic substances.10 This process can be
prevented by the use of a large amount of Cl2. The
addition of a large quantity of Cl2 is associated with a
shortened contact time that prevents the formation of
trihalomethanes. Residual chlorine can then be removed
from the water by the addition of SO2. This practice
explains why SO4–2 ions were detected in the water.
However, it is important to note, that the concentration of
the SO4–2 in comparison with the chlorine is less
important for corrosion.
Metals with a negative potential are easier to dissolve
(corrode) than metals with a positive potential. It is clear
that iron dissolves when in contact with chlorine. Chlo-
rine is present in bathing water, as well as in various dis-
infectants and cleaning agents. The normal potential for
iron is –0.44 V, while for chlorine it is +1.36 V, an abso-
lute potential difference of 1.80 (Table 3).
One of the inspected cases was that of a waterslide.
In its project documents, stainless-steel AISI 316 to AISI
316Ti was the prescribed material for the construction.
AISI 316 stainless steel contains a chromium (Cr), nickel
(Ni) and molybdenum (Mo) alloy of metals (Tables 4
and 5 13). In addition, it contains small quantities of
phosphorus (P), sulphur (S), as well as some other ele-
ments.
We note that in the table of stainless steels recom-
mended for swimming-pool construction by the
Deutsches Institut für Bautechnik, AISI 316 and AISI
316Ti are not listed. 8 This was overlooked by the
designer, thus the resulting corrosion is not a coincidence
and corrosion appeared on the stainless-steel nuts
(Figure 3), which can endanger the the stability of
structure. Due to the occurrence of corrosion, protective
coatings were applied to the inappropriately chosen
stainless steel. However, this did not ensure adequate
protection; as shown in Figure 4 the protective layer
peeled off the protected surface and consequently the
investor enforced the warranty for the waterslide
structure.
L. GOSAR, D. DREV: CONTACT WITH CHLORINATED WATER: SELECTION OF THE APPROPRIATE STEEL
642 Materiali in tehnologije / Materials and technology 45 (2011) 6, 639–644
Figure 4: Protective coating peeling off of stainless steel structure
Slika 4: Prikaz slabe povr{inske za{~ite nerjave~e jeklene konstrukcije
v kopalnem bazenu
Figure 3: Corroded stainless steel nuts on the waterslide structure
Slika 3: Prikaz korozije na matici izdelani iz nerjavnega jekla na
konstrukciji tobogana
Table 3: Metals typically used in construction 11,12
Preglednica 3: Najbolj poznane konstrukcijske kovine 11,12
Metal Voltage/V Metal
Voltage
/V
Metal,
nonmetal
Voltage
/V
potassium – 2.9 cadmium – 0.40 silver + 0.80
sodium – 2.7 cobalt – 0.29 mercury + 0.86
manganese – 2.34 nickel – 0.25 gold + 1.68
aluminum – 1.28 tin – 0.14 platinum + 1.18
manganese – 1.05 lead – 0.12 sulphur – 0.51
zinc –0.76 antimony + 0.20 hydrogen 0.00
chromium – 0.56 arsenic + 0.30 oxygen + 0.39
iron – 0.44 copper + 0.34 chlorine + 1.36
Table 4: Features of the stainless steel prescribed by the designer
Preglednica 4: Lastnosti nerjavnega jekla, ki ga je predpisal pro-
jektant
W.Nr. DIN AISI JUS
1.4436 X5CrNiMo 17 13 3 316 ^.45706
1.4571 X6CrNiMoTi 17 12 2 316Ti ^.4574
Table 5: Conversion table for stainless steel AISI 316/316Ti tags to
the European standard
Preglednica 5: Tebela za pretvorbo oznak nerjavnega jekla AISI
316/316Ti v evropske standardne oznake
Standards
(Europe)
Chemical composition / %
P
max
S
max
Si
max Mo
Other
elements
X5 CrNiMo
17-12-2 0,045 0,015 1 2-2,5 N 0,11 max
X6 CrNiMoTi
17-12-2 0,045 0,015 1 2-2,5
Ti=5 x Cmin;
0,7 max
3 DISCUSSION
The corrosion of stainless steel can be prevented by
choosing an appropriate type of steel (steel alloys with a
higher content of certain metals – Cr, Ni, Mo), by
chemical treatment (pickling), thermal processing
(annealing) and surface treatment (grinding, polishing).
To avoid the corrosion of stainless steels, it is first
necessary to respect the following rules:
– no use of tools (wrenches, pliers, vices) that were
previously applied for work with non-corrosion-
resistant steel,
– no use of abrasive wheels that were previously used
for grinding or cutting non-corrosion-resistant steel,
– the filings of non-corrosion-resistant steel must not
come into contact with the surface of the stainless
steel,
– no use of cutting tools (saws, files, etc.) that were
previously applied for work with non-corrosion-
resistant steel,
– no use of sanding cloths and brushes that were
previously used for processing non-corrosion-resis-
tant steel.
The rules that define materials used in food
processing and swimming-pool structures do not detail
which materials must be used. Slovenian and EU rules
define which materials can be used in facilities where
they may come in contact with food. It is important to
note that in this case water falls under the definition of
food. Similarly, the Slovenian regulations for technical
measures related to the safety of swimming pools 14 do
not clearly define which materials may be used in the
construction and operation of these facilities. The rules
only simply state (Article 25) that.15
(1) All swimming pools and pool platforms must meet
the requirements of SIST EN 13451, parts 1–9.
(2) Waterslides must be designed and constructed in
accordance with SIST EN 1069-1 and SIST EN
1069-2.
SIST EN 13451 provides general safety requirements
and test methods for swimming pools, while SIST EN
1069 -1 and SIST EN 1069 -2 prescribe safety require-
ments and test methods for waterslides.
4 CONCLUSION
It is evident from the cases presented that in the field
of sanitary engineering the use of appropriate steel types
is of paramount importance. However, even when the
appropriate steel type is used and standard surface
protection is applied, corrosion may still occur. In the
example of the internal water-supply network shown, a
protective layer of zinc had been removed due to the
implementation of disinfection. A need for the repeated
disinfection of the water-supply network arose from the
contamination of the system with Legionella bacteria.
Disinfection was executed by inappropriate methods and
induced a significant corrosion of the water-supply
system, while the contamination problem was not solved.
Legionella bacteria could not be successfully destroyed,
neither by chlorine shocks nor by thermal shocks. When
inducing thermal shocks, it was not possible to reach a
sufficient temperature in all the parts of the system
simultaneously, thus the contamination was merely
transmitted from the contaminated part to the
decontaminated parts of the networks. Effective
disinfection was later achieved by introducing oxidative
disinfectants based on H2O2 and Ag into the system.
However, this resulted in corrosion and lowering of the
chemical and organoleptic quality of water. In the case of
a public swimming pool, stainless steel of insufficient
quality was used and corrosion became evident after
only a few months of use. When the steel was protected
with a plastic coating, this started to flake. The lesson is
that stainless steel should be appropriately chosen, and
protective coatings should be applied before installation.
By selecting appropriate protective coatings,
corrosion can be prevented for the entire lifetime of a
facility. Hot galvanized steel elements can adequately
protect the steel for up to ten years or more, until the
protective layer of zinc is dissolved. Polymeric coatings
are also effective, but it is possible for corrosion to occur
at the edges of the structural elements beneath the top
layer. However, if an appropriate type of stainless steel is
selected ex ante, we can remove the threat of corrosion
altogether as long as the operation of the facility is
performed in accordance with specified rules. The choice
of high-grade stainless steel is associated with higher
investment costs and lower operating costs. In the design
and construction of sanitary engineering facilities, it is
therefore necessary to think comprehensively about
corrosion prevention, as the selection of appropriate steel
types does not only provide structural strength; it can
also ensure healthy drinking or bathing water, while
reducing operational costs.
5 REFERENCES
1 C. Jullien, T. Bénézech, B. Carpentier, V. Lebret, C. Faille, Identifi-
cation of surface characteristics relevant to the hygienic status of
stainless steel for the food industry, Journal of Food Engineering, 56
(2003) 1, 77–87
2 G. Kreye, M. Schütze, Corrosion Handbook, John Wiley&Sons
(1998)
3 A. Sander, B. Berghult, A. Elfström Broo, E. Lind Johansson, T.
Hedberg, Iron corrosion in drinking water distribution systems – The
effect of pH, calcium and hydrogen carbonate, Corrosion Science, 38
(1996), 443–455
4 W. F. Smith, Principles of Materials Science & Engeneering, 2.edt.,
Mc Graw Hill, New York, 1990
5 G. Banerjee, T. K. Pal, N. Bandyopadhyay, D. Bhattacharjee, Effect
of welding conditions on corrosion behaviour of spot welded coated
steel sheets, Corrosion Engineering, Science and Technology, 46
(2011) 1, 64–69
6 D. Jiasheng, C. Sihong, L. Manqi, Y. Ke, History and Current Status
of Antibacterial Materials, Materials Chemistry and Physics, 89
(2005) 1, 38–48
L. GOSAR, D. DREV: CONTACT WITH CHLORINATED WATER: SELECTION OF THE APPROPRIATE STEEL
Materiali in tehnologije / Materials and technology 45 (2011) 6, 639–644 643
7 Pourbaix, M., Atlas of electrochemical equilibria in aqueous solu-
tions. 2d English ed., Houston 1974
8 Anstalt des öffentlichen Rechts, Allgemeine bauaufsichtliche Zu-
lassung, Deutsches Institut für Bautechnik, Z-30.3.6. 2003
9 S. D. Richardson, Disinfection by-products and other emerging
contaminants in drinking water, TrAC Trends in Analytical
Chemistry, 22 (2003) 10, 666–684
10 M. Panyakapo, S. Soontornchai, P. Paopuree, Cancer risk assessment
from exposure to trihalomethanes in tap water and swimming pool
water, Journal of Environmental Sciences, 20 (2008) 3, 372–378
11 P. Leskovar, Tehnologija gradiv, Technology Materials, Univerza v
Ljubljani, Ljubljana 1970
12 P. Leskovar, Gradiva, Del 2, Preizku{anje kovin, litje, va`nej{a
nekovinska gradiva, korozija in povr{inska za{~ita, Testing of metal,
casting, more important non-metallic materials, corrosion and
surface protection, Univerza v Ljubljani, Ljubljana 1986
13 Rostfrei in Schwimmbädern, Merkblatt 831, Edelstahl 2000
14 Pravilnik o tehni~nih ukrepih in zahtevah za varno obratovanje
kopali{~ in za varstvo pred utopitvami na kopali{~ih, Offical journal
of RS No. 88/2003, 56/2006, 84/2007
15 Standards SIST EN 1069-1, SIST EN 1069-2 and SIST EN 13451
L. GOSAR, D. DREV: CONTACT WITH CHLORINATED WATER: SELECTION OF THE APPROPRIATE STEEL
644 Materiali in tehnologije / Materials and technology 45 (2011) 6, 639–644
LETNO KAZALO – INDEX
Letnik / Volume 45 2011
ISSN 1580-2949
© Materiali in tehnologije
IMT Ljubljana, Lepi pot 11, 1000 Ljubljana, Slovenija
M
EHNOLOGIJEIN
AT E R IALI
M A T E R I A L S A N D T E C H N O L O G Y
MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY
VSEBINA / CONTENTS
LETNIK / VOLUME 45, 2011/1, 2, 3, 4, 5, 6
2011/1
Experimental study of some masonry-wall coursework material types under horizontal loads and their comparison
Eksperimentalna raziskava zgradbe nekaterih zidarskih zidov – vodoravna obremenitev in primerjava uporabljenih materialov
M. Kamanli, M. S. Donduren, M. T. Cogurcu, M. Altin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3
Synthesis of aluminium foams by the powder-metallurgy process: compacting of precursors
Sinteza aluminijevih pen po postopku metalurgije prahov: stiskanje prekurzorjev
I. Paulin, B. [u{tar{i~, V. Kevorkijan, S. D. [kapin, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13
A new method for determining the remaining lifetime of coated gas-turbine blades
Nova metoda za izra~un preostale trajnostne dobe lopatic plinskih turbin
L. B. Getsov, P. G. Krukovski, N .V. Mozaiskaja, A. I. Rybnikov, K. A. Tadlja . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21
Reduction of ultra-fine tungsten powder with tungsten (VI)-oxide in a vertical tube reactor
Redukcija ultrafinih prahov volframovega(VI) oksida v reaktorju z vertikalno cevjo
@. Kamberovi}, D. Filipovi}, K. Rai}, M. Tasi}, Z. An|i}, M. Gavrilovski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27
Oxygen diffusion in the non-evaporable getter St 707 during heat treatment
Difuzija kisika v getru St 707 med toplotno obdelavo
S. Avdiaj, B. [etina - Bati~, J. [etina, B. Erjavec . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33
The modeling of auger spectra
Modeliranje augerjevih spektrov
B. Poniku, I. Beli~, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 39
Modelling of the directional solidification of a leaded red brass flange
Modeliranje usmerjenega strjevanja prirobnice iz rde~e svin~eve medenine
V. Grozdani} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47
Characterization of the inclusions in spring steel using light microscopy and scanning electron microscopy
Karakterizacija vklju~kov v vzmetnih jeklih s svetlobno in vrsti~no elektronsko mikroskopijo
A. Bytyqi, N. Puk{i~, M. Jenko, M. Godec . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55
Characterization of the carbides in a Ni-Ti shape-memory alloy wire
Karakterizacija karbidov v @ici zlitine s spominom Ni-Ti
M. Godec, A. Kocijan, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 61
Fracture toughness of the vacuum-heat-treated spring steel 51CrV4
Lomna `ilavost vakuumsko toplotno obdelanega vzmetnega jekla 51CrV4
B. Sen~i~, V. Leskov{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 67
Similarity criteria and effect of lubricant inertia at cold rolling
Merila podobnosti in vpliv vztrajnosti maziva pri hladnem valjanju
D. ]ur~ija, F. Vodopivec, I. Mamuzi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75
In memoriam: Oskar Kürner 1925–2010
M. Gabrov{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 81
2011/2
Material failure of an AISI 316L stainless steel hip prosthesis
Napake materiala v kol~ni protezi iz nerjavnega jekla AISI 316L
M. Godec . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 85
A comparison of the corrosion behaviour of austenitic stainless steels in artificial seawater
Primerjava korozijskih lastnosti avstenitnih nerjavnih jekel v simulirani morski vodi
A. Kocijan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 91
Influence of the foaming precursor’s composition and density on the foaming efficiency, microstructure development and
mechanical properties of aluminium foams
Vpliv sestave in gostote prekurzorjev za penjenje na u~inkovitost penjenja ter razvoj mikrostrukture in mehanskih lastnosti aluminijskih pen
V. Kevorkijan, S. D. [kapin, I. Paulin, B. [u{tar{i~, M. Jenko, M. La`eta. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95
Multi-objective optimization of the cutting forces in turning operations using the Grey-based Taguchi method
Multi namenska optimizacija stru`enja z uporabo Taguchi metode na Grey podlagi
Y. Kazancoglu, U. Esme, M. Bayramo

glu, O. Guven, S. Ozgun . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 105
646 Materiali in tehnologije / Materials and technology 45 (2011) 6
LETNO KAZALO – INDEX
Thermodynamic conditions for the nucleation of boron compounds during the cooling of steel
Termodinami~ni pogoji za nukleacijo borovih spojin pri ohlajanju jekla
Z. Adolf, J. Ba`an, L. Socha . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111
A micro-damage investigation on a low-alloy steel tested using a 7.62-mm AP projectile
Raziskava mikropo{kodb malolegiranega jekla po preizkusu s kroglo AP 7,62 mm
T. Demir, M. Übeyli . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 115
Activation of polymer polyethylene terephthalate (PET) by exposure to CO2 and O2 plasma
Aktivacija polimera polietilentereftalata (PET) s CO2- ali O2-plazmo
A. Vesel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 121
The impact on rigid PVC pipes: a study of the correlation between the length of the crazed zone and the area of the impacted
region
Udar togih PVC-cevi: {tudija korelacije med dol`ino razpokane zone in povr{ino zone udara
C. B. Fokam, M. Chergui, K. Mansouri, M. El Ghorba, M. Mazouzi . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 125
A comparative analysis of theoretical models and experimental research for spray drying
Primerjalna analiza teoreti~nih modelov in eksperimentalna raziskava razpr{ilnega su{enja
D. Tolmac, S. Prvulovic, D. Dimitrijevic, J. Tolmac . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 131
Creep resistance of microstructure of welds of creep resistant steels
Odpornost proti lezenju pri mikrostrukturi zvarov jekel, odpornih proti lezenju
F. Vodopivec, M. Jenko, R. Celin, B. @u`ek, D. A. Skobir . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 139
Effect of the martensite volume fraction on the machining of a dual-phase steel using a milling operation
Vpliv volumenskega dele`a martenzita na obdelavo dvofaznega dualnega jekla z rezkanjem
O. Topçu, M. Übeyli, A. Acir . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 145
Degradation of a Ni-Cr-Fe alloy in a pressurised-water nuclear power plant
Degradacija zlitin Ni-Cr-Fe v tla~novodnih jedrskih elektrarnah
R. Celin, F. Tehovnik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 151
Investigation into the mechanical properties of micro-alloyed as-cast steel
Raziskave mehanskih lastnosti mikrolegiranih jekel
B. Chokkalingam, S. S. M. Nazirudeen, S. S. Ramakrishnan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 159
The effect of electromagnetic stirring on the crystallization of concast billets
Kristalizacija kontinuirno ulitih gredic v elektromagnetnem polju
K. Stransky, F. Kavicka, B. Sekanina, J. Stetina,V. Gontarev, J. Dobrovska. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 163
2011/3
Microwave-assisted non-aqueous synthesis of ZnO nanoparticles
Sinteza nanodelcev ZnO v nevodnem mediju pod vplivom mikrovalov
G. Ambro`i~, Z. Crnjak Orel, M. @igon. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 173
Removal of a thin hydrogenated carbon film by oxygen plasma treatment
Odstranjevanje tanke plasti hidrogeniranega ogljika s kisikovo plazmo
U. Cvelbar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 179
Low temperature destruction of bacteria Bacillus stearothermophilus by weakly ionized oxygen plasma
Nizkotemperaturno uni~evanje bakterij Bacillus stearothermophilus s {ibko ionizirano kisikovo plazmo
M. Mozeti~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 185
Surface characterization of polymers by XPS and SIMS techniques
Analiza povr{ine polimerov z metodama XPS in SIMS
J. Kova~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 191
Modification of PET-polymer surface by nitrogen plasma
Modifikacija povr{ine PET-polimera z du{ikovo plazmo
R. Zaplotnik, M. Kolar, A. Doli{ka, K. Stana-Kleinschek. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 199
Functionalization of AFM tips for use in force spectroscopy between polymers and model surfaces
Funkcionalizacija AFM-konic za uporabo v spektroskopiji sil med polimeri in modelnimi povr{inami
T. Maver, K. Stana - Kleinschek, Z. Per{in, U. Maver . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 205
A novel approach for qualitative determination of residual tin based catalyst in poly(lactic acid) by X-ray photoelectron
spetroscopy
Nov na~in kvalitativne dolo~itve vsebnosti katalizatorja kositra v polilakti~ni kislini z rentgensko fotoelektronsko spektroskopijo
V. Sedlaøík, A. Vesel, P. Kucharczyk, P. Urbánek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 213
Hydrophobization of polymer polystyrene in fluorine plasma
Hidrofobizacija polimera polistiren s fluorovo plazmo
A. Vesel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 217
Materiali in tehnologije / Materials and technology 45 (2011) 6 647
LETNO KAZALO – INDEX
Plasma treatment of biomedical materials
Plazemska obdelava biomedicinskih materialov
I. Junkar, U. Cvelbar, M. Lehocky . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 221
Radiofrequency induced plasma in large-scale plasma reactor
Radiofrekven~no inducirana plazma v reaktorju velikih dimenzij
R. Zaplotnik, A. Vesel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 227
Modification of surface morphology of graphite by oxygen plasma treatment
Sprememba morfologije grafita med obdelavo s kisikovo plazmo
K. Eler{i~, I. Junkar, M. Modic, R. Zaplotnik, A. Vesel, U. Cvelbar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 233
Properties of particleboards made by using an adhesive with added liquefied wood
Lastnosti ivernih plo{~, izdelanih z uporabo lepila z dodanim uteko~injenim lesom
N. ^uk, M. Kunaver, S. Medved . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 241
Poly(styrene-CO-divinylbenzene-CO-2-ethylhexyl)acrylate membranes with interconnected macroporous structure
Poli(stiren-KO-divinilbenzen-KO-2-etilheksil)akrilatne membrane s povezano porozno strukturo
U. Sev{ek, S. Seifried, ^. Stropnik, I. Pulko, P. Krajnc . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 247
Modification of non-woven cellulose for medical applications using non-equlibrium gassious plasma
Modifikacija celuloznih kopren, uporabnih v medicinske namene, z neravnovesno plinsko plazmo
K. Stana - Kleinschek, Z. Per{in, T. Maver . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 253
Use of AFM force spectroscopy for assessment of polymer response to conditions similar to the wound, during healing
Uporaba AFM-spektroskopije sil za spremljanje odziva polimernih molekul na v rani podobna okolja med celjenjem
U. Maver, T. Maver, A. @nidar{i~, Z. Per{in, M. Gaber{~ek, K. Stana-Kleinschek. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 259
Biodegradable polymers from renewable resources: effect of proteinic impurityon polycondensation products of
2-hydroxypropanoic acid
Biorazgradljivi polimeri iz obnovljivih virov: vpliv proteinskih ne~istot na produkte polikondenzacije 2-hidroksipropanojske kisline
I. Poljan{ek, P. Kucharczyk,V. Sedlaøík, V. Ka{párková, A. [alaková, J. Drbohlav . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 265
Sub micrometer and nano ZnO as filler in PMMA materials
Submikrometrski in nano ZnO kot polnilo v PMMA-materialih
A. An`lovar, Z. Crnjak Orel, M. @igon . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 269
Tuning of poly(ethylene terephtalate) (PET) surface properties by oxygen plasma treatment
Prilagoditev lastnosti povr{ine polietilen tereftalata (PET) z obelavo v kisikovi plazmi
A. Doli{ka, M. Kolar. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 275
Probability of recombination and oxidation of O atoms on a-C:H surface
Verjetnost za rekombinacijo in oksidacijo za atome kisika na povr{ini a-C:H
A. Drenik, K. Eler{i~, M. Modic, P. Panjan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 281
Hydrothermal growth of Zn5(OH)6(CO3)2 and its thermal transformation into porous ZnO film used for dye-sensitized solar
cells
Hidrotermalna rast Zn5(OH)6(CO3)2 s termi~no transformacijo v porozno plast ZnO, uporabljeno za elektrokemijske son~ne celice
M. Bitenc, Z. Crnjak Orel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 287
2011/4
High-strength low-alloy (HSLA) steels
Visokotrdna malolegirana (HSLA) konstrukcijska jekla
D. A. Skobir . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 295
TEM investigation of metallic materials – an advanced technique in materials science and metallurgy
Preiskave kovinskih materialov s presevno elektronsko mikroskopijo – moderna tehnika v znanosti o materialih in metalurgiji
D. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 303
Modelling of hot tears in continuously cast steel
Modeliranje vro~ih razpok v kontinuirno litem jeklu
V. Grozdani} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 311
Thermodynamic investigation of the Al-Sb-Zn system
Termodinamska raziskava sistema Al-Sb-Zn
G. Klan~nik, J. Medved. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 317
XPS and SEM of unpolished and polished FeS surface
Rentgenska fotoelektronska spektroskopija in vrsti~na elektronska mikroskopija nepolirane in polirane povr{ine FeS
Dj. Mandrino . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 325
Surface characterization and pickling characteristics of the oxide scale on duplex stainless steel
Povr{inska karakterizacija in lastnosti lu`enja oksidne plasti na dupleksnem nerjavnem jeklu
^. Donik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 329
648 Materiali in tehnologije / Materials and technology 45 (2011) 6
LETNO KAZALO – INDEX
CFD analysis of exothermic reactions in Al-Au nano multi-layered foils
CFD-analiza eksotermnih reakcij v ve~plastnih nanofolijah Al-Au
K. T. Rai}, R. Rudolf, P. Ternik, Z. @uni~, V. Lazi}, D. Stamenkovi}, T. Tanaskovi}, I. An`el . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 335
Microstructure evolution in SAF 2507 super duplex stainless steel
Razvoj mikrostrukture v superdupleksnem nerjavnem jeklu SAF 2507
F. Tehovnik, B. Arzen{ek, B. Arh, D. Skobir, B. Pirnar, B. @u`ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 339
Optimization of the quality of continuously cast steel slabs using the Firefly algorithm
Optimizacija kakovosti kontinuirno lite jeklene plo{~e z uporabo algoritma "Firefly"
T. Mauder, C. Sandera, J. Stetina, M. Seda . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 347
Hot workability of 95MnWCr5 tool steel
Vro~a preoblikovalnost orodnega jekla 95MnWCr5
A. Kri`aj, M. Fazarinc, M. Jenko, P. Fajfar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 351
Lifetime evaluation of a steam pipeline using NDE methods
Ocena preostale trajnostne dobe parovoda z uporabo neporu{itvenih preiskav (NDE)
F. Kafexhiu, J. Vojvodi~ Tuma. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 357
The influence of the chemical composition of steels on the numerical simulation of a continuously cast slab
Vpliv kemi~ne sestave jekel na numeri~no simulacijo kontinuirno lite plo{~e
J. Stetina, F. Kavicka. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 363
Prediction of the mechanical properties of cast Cr-Ni-Mo stainless steels with a two-phase microstructure
Napoved mehanskih lastnosti litih Cr-Ni-Mo nerjavnih jekel z dvofazno mikrostrukturo
M. Male{evi}, J. V. Tuma, B. [u{tar{i~, P. Borkovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 369
Relationship between mechanical strength and Young’s modulus in traditional ceramics
Odvisnost med mehansko trdnostjo in Youngovim modulom pri tradicionalni keramiki
I. [tubòa, A. Trník, P. [ín, R. Sokoláø, I. Medveï . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 375
2011/5
Franc Vodopivec – osemdesetletnik
Laudation in honour of Franc Vodopivec on the occasion of his 80th birthday
M. Jenko. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 381
Cr-V ledeburitic cold-work tool steels
Ledeburitna jekla Cr-V za delo v hladnem
P. Jur~i . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 383
Experimental comparison of resistance spot welding and friction-stir spot welding processes for the EN AW 5005 aluminum
alloy
Eksperimentalna primerjava odpornosti procesov to~kovnega varjenja in to~kovnega tornega varjenja pri aluminijevi zlitini EN AW 5005
M. K. Kulekci, U. Esme, O. Er . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 395
The friction and wear behavior of Cu-Ni3Al composites by dry sliding
Trenje in obraba Cu-Ni3Al kompozitov pri suhem drsenju
M. Demirel, M. Muratoglu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 401
Weldability of metal matrix composite plates by friction stir welding at low welding parameters
Varivost plo{~ kompozita s kovinsko osnovo po vrtilno tornem postopku pri nizkih varilnih parametrih
Y. Bozkurt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 407
Influence of the gas composition on the geometry of laser-welded joints in duplex stainless steel
Vpliv vrste za{~itnega plina na geometrijo zvara pri laserskem varjenju nerjavnega dupleksnega jekla
B. Bauer, A. Topi}, S. Kralj, Z. Ko`uh . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 413
Multiscale modelling of heterogeneous materials
Mikro in makro modeliranje heterogenih materialov
M. Lamut, J. Korelc, T. Rodi~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 421
Genetic programming and soft-annealing productivity
Genetsko programiranje in produktivnost mehkega `arjenja
M. Kova~i~, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 427
Semi-solid gel electrolytes for electrochromic devices
Poltrdni gelski elektroliti za elektrokromne naprave
M. Hajzeri, M. ^olovi}, A. [urca Vuk, U. Posset, B. Orel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 433
Combustible precursor behaviour in the lanthanum chromite formation process
Termi~ne lastnosti reakcijskega gela za pripravo lantanovega kromita
K. Zupan, M. Marin{ek, B. Novosel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 439
Materiali in tehnologije / Materials and technology 45 (2011) 6 649
LETNO KAZALO – INDEX
Nanoscale modification of hard coatings with ion implantation
Nanovelikostna modifikacija trdnih prekritij z ionsko implantacijo
B. [kori}, D. Kaka{, M. Gostimirovi}, A. Mileti} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 447
Influence of the granulation and grain shape of quartz sands on the quality of foundry cores
Vpliv granulacije in oblike zrn kremenovega peska na kakovost livarskih jeder
M. Marin{ek, K. Zupan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 451
Characterization of extremely weakly ionized hydrogen plasma with a double Langmuir probe
Karakterizacija {ibko ionizirane vodikove plazme z dvojno Langmuirjevo sondo
M. Mozeti~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 457
Optical properties of plastically deformed copper: an ellipsometric study
Opti~ne lastnosti plasti~no derformiranega bakra: {tudij elipsometrije
N. Rom~evi}, R. Rudolf, J. Traji}, M. Rom~evi}, B. Had`i}, D. Vasiljevi} – Radovi}, I. An`el . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 463
Relaxation of the residual stresses produced by plastic deformation
Relaksacija zaostalih napetosti zaradi plasti~ne deformacije
N. Tadi}, M. Jeli}, D. Lu~i}, M. Mi{ovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 467
Accelerated corrosion behaviors of Zn, Al and Zn/15Al coatings on a steel surface
Pospe{eno korozijsko obna{anje Zn, Al in Zn/15Al prekritij na povr{ini jekla
A. Gulec, O. Cevher, A. Turk, F. Ustel, F. Yilmaz . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 477
Alloys with modified characteristics
Zlitine z modificiranimi lastnostmi
M. Oru~, M. Rimac, O. Beganovi}, S. Muhamedagi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 483
Evaluation of the microstructural changes in Cr-V ledeburitic tool steels depending on the austenitization temperature
Ocena sprememb mikrostrukture v ledeburitnemn orodnem jeklu Cr-V v odvisnosti od temperature avstenitizacije
P. Bílek, J. Sobotová, P. Jur~i . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 489
In memoriam: Hans Jürgen Grabke
M. Jenko. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 495
2011/6
New discovered paradoxes in theory of balancing chemical reactions
Novoodkriti paradoksi v teoriji uravnote`enja kemijskih reakcij
I. B. Risteski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 503
Characteristics of creep in conditions of long operation
Zna~ilnosti lezenja pri dolgotrajni uporabi
N. A. Katanaha, L. B. Getsov . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 523
A thermodynamic and kinetic study of the solidification and decarburization of malleable cast iron
Termodinami~na in kineti~na analiza strjevanja in razoglji~enja belega litega `eleza
M. Pirnat, P. Mrvar, J. Medved . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 529
Modelling and preparation of core foamed Al panels with accumulative hot-roll bonded precursors
Na~rtovanje in izdelava Al-panelov s sredico iz Al-pen na osnovi ve~stopenjsko toplo valjanih prekurzorjev
V. Kevorkijan, U. Kova~ec, I. Paulin, S. D. [kapin, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 537
Numerical solution of hot shape rolling of steel
Numeri~na re{itev vro~ega valjanja jekla
U. Hanoglu, Siraj-ul-Islam, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 545
Solidification and precipitation behaviour in the AlSi9Cu3 alloy with various Ce additions
Strjevanje in izlo~anje v zlitini ALSI9CU3 pri razli~nih dodatkih Ce
M. Von~ina, S. Kores, P. Mrvar, J. Medved . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 549
Effect of change of carbide particles spacing and distribution on creep rate of martensite creep resistant steels
Vpliv spremembe razdalje med karbidnimi izlo~ki in njihove porazdelitve na hitrost lezenja martenzitnih jekel, odpornih proti lezenju
D. A. Skobir Balanti~, M. Jenko, F. Vodopivec, R. Celin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 555
Stress-strain analysis of an abutment tooth with rest seat prepared in a composite restoration
Napetostno-deformacijska analiza opornega zoba z zapornim sede`em, izdelana s kompozitnim popravilom
Lj. Tiha~ek [oji}, A. M. Lemi}, D. Stamenkovi}, V. Lazi}, R. Rudolf, A. Todorovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 561
Identification and verification of the composite material parameters for the Ladevèze damage model
Identifikacija in verifikacija parametrov kompozitnega materiala za model Ladevèze
V. Kleisner, R. Zem~ík, T. Kroupa. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 567
Evaluation of the strength variation of normal and lightweight self-compacting concrete in full scale walls
Ocena variacije trdnosti normalnega in lahkega vibriranega betona v polnih stenah
M. M. Ranjbar, M. Hosseinali Beygi, I. M. Nikbin, M. Rezvani, A. Barari . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 571
650 Materiali in tehnologije / Materials and technology 45 (2011) 6
LETNO KAZALO – INDEX
The influence of buffer layer on the properties of surface welded joint of high-carbon steel
Vpliv vmesne plasti na lastnosti povr{inskih zvarov jekla z veliko ogljika
O. Popovi}, R. Proki} - Cvetkovi}, A. Sedmak, G. Buyukyildrim, A. Bukvi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 579
Corrosion stability of different bronzes in simulated urban rain
Korozijska stabilnost razli~nih bronov v umetnem kislem de`ju
E. [vara Fabjan, T. Kosec, V. Kuhar, A. Legat . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 585
Morphology and corrosion properties PVD Cr-N coatings deposited on aluminium alloys
Morfologija in korozijske lastnosti CrN PVD-prevlek, nanesenih na aluminijeve zlitine
D. Kek Merl, I. Milo{ev, P. Panjan, F. Zupani~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 593
The effect of electromagnetic stirring on the crystallization of concast billets
Vpliv elektromagnetnega me{anja na kristalizacijo kontinuirno ulitih gredic
F. Kavicka, K. Stransky, B. Sekanina, J. Stetina, V. Gontarev, T. Mauder, M. Masarik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 599
Wear of refractory materials for ceramic filters of different porosity in contact with hot metal
Obraba ognjevzdr`nega materiala kerami~nih filtrov z razli~no poroznostjo v stiku z vro~o kovino
J. Ba`an, L. Socha, L. Martínek, P. Fila, M. Balcar, J. Chmelaø . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 603
The influence of the mineral content of clay from the white bauxite mine on the properties of the sintered product
Vpliv vsebnosti minerala gline iz rudnika belega boksita na lastnosti sintranega proizvoda
M. Krgovi}, I. Bo{kovi}, M. Vuk~evi}, R. Zejak, M. Kne`evi}, R. Mitrovi}, B. Zlati~anin, N. Ja}imovi} . . . . . . . . . . . . . . . . . . . . . . . . 609
Effect of pre-straining on the springback behavior of the AA5754-0 alloy
Vpliv prenapenjanja na povratno elasti~no izravnavo zlitine AA5754-0
S. Toros, M. Alkan, R. Ecmel Ece, F. Ozturk. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 613
Heat treatment and mechanical properties of heavy forgings from A694–F60 steel
Toplotna obdelava in mehanske lastnosti te`kih izkovkov iz jekla A694-F60
M. Balcar, J. Novák, L. Sochor, P. Fila, L. Martínek, J. Ba`an, L. Socha, D. A. Skobir Balanti~, M. Godec . . . . . . . . . . . . . . . . . . . . . . 619
The tensile behaviour of friction-stir- welded dissimilar aluminium alloys
Natezne zna~ilnosti tornih pomi~nih zvarov razli~nih aluminijevih zlitin
R. Palanivel, P. Koshy Mathews. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 623
Screen-printed electrically conductive functionalities in paper substrates
Elektroprevodne oblike, pripravljene s sitotiskom na papirnih podlagah
M. @vegli~, N. Hauptman, M. Ma~ek, M. Klanj{ek Gunde . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 627
Recent growing demand for magnesium in the automotive industry
Rast povpra{evanja po magneziju v avtomobilski industriji
M. J. Freiría Gándara . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 633
Contact with chlorinated water: selection of the appropriate steel
Kontakt s klorirano vodo – izbor ustreznega jekla
L. Gosar, D. Drev . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 639
Materiali in tehnologije / Materials and technology 45 (2011) 6 651
LETNO KAZALO – INDEX
MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY
AVTORSKO KAZALO / AUTHOR INDEX
LETNIK / VOLUME 45, 2011, 1–6, A–@
A
Acir A. 145
Adolf Z. 111
Alkan M. 613
Altin M. 3
Ambro`i~ G. 173
An`el I. 335, 463
An`lovar A. 269
An|i} Z. 27
Arh B. 339
Arzen{ek B. 339
Avdiaj S. 33
B
Balcar M. 603, 619
Barari A. 571
Bauer B. 413
Bayramolu M. 105
Ba`an J. 111, 603, 619
Beganovi} O. 483
Beli~ I. 39
Beygi M. H. 571
Bílek P. 489
Bitenc M. 287
Bo{kovi} I. 609
Borkovi} P. 369
Bozkurt Y. 407
Bukvi} A. 579
Buyukyildrim G. 579
Bytyqi A. 55
C
Celin R. 139, 151, 555
Cevher O. 477
Chergui M. 125
Chmelaø J. 603
Chokkalingam B. 159
Cogurcu M. T. 3
Crnjak Orel Z. 173, 269, 287
Cvelbar U. 179, 221, 233
^
^olovi} M. 433
^uk N. 241
]
]ur~ija D. 75
D
Demir T. 115
Demirel M. 401
Dimitrijevic D. 131
Dobrovska J. 163
Doli{ka A. 199, 275
Donduren M. S. 3
Donik ^. 329
Drbohlav J. 265
Drenik A. 281
Drev D. 639
E
Ece R. E. 613
El Ghorba M. 125
Eler{i~ K. 233, 281
Er O. 395
Erjavec B. 33
Esme U. 105, 395
F
Fajfar P. 351
Fazarinc M. 351
Fila P. 603, 619
Filipovi} D. 27
Fokam C. B. 125
Freiría Gándara M. J. 633
G
Gaber{~ek M. 259
Gavrilovski M. 27
Getsov L. B. 21, 523
Godec M. 55, 61, 85, 619
Gontarev V. 163, 599
Gosar L. 639
Gostimirovi} M. 447
Grozdani} V. 47, 311
Gulec A. 477
Guven O. 105
H
Had`i} B. 463
Hajzeri M. 433
Hanoglu U. 545
Hauptman N. 627
I
Islam S. 545
J
Ja}imovi} N. 609
Jeli} M. 467
Jenko D. 303
Jenko M. 13, 39, 55, 61, 95, 139,
351, 381, 537, 555
Junkar I. 221, 233
Jur~i P. 383, 489
K
Ka{párková V. 265
Kafexhiu F. 357
Kaka{ D. 447
Kamanli M. 3
Kamberovi} @. 27
Katanaha N. A. 523
Kavicka F. 163, 363, 599
Kazancoglu Y. 105
Kek Merl D. 593
Kevorkijan V. 13, 95, 537
Klan~nik G. 317
Klanj{ek Gunde M. 627
Kleisner V. 567
Kne`evi} M. 609
Ko`uh Z. 413
Kocijan A. 61, 91
Kolar M. 199, 275
Korelc J. 421
Kores S. 549
Kosec T. 585
Koshy Mathews P. 623
Kova~ J. 191
Kova~ec U. 537
Kova~i~ M. 427
Krajnc P. 247
Kralj S. 413
Krgovi} M. 609
Kri`aj A. 351
Kroupa T. 567
Krukovski P. G. 21
Kucharczyk P. 213, 265
Kuhar V. 585
Kulekci M. K. 395
Kunaver M. 241
L
Lamut M. 421
652 Materiali in tehnologije / Materials and technology 45 (2011) 6
LETNO KAZALO – INDEX
Lazi} V. 335, 561
La`eta M. 95
Legat A. 585
Lehocky M. 221
Lemi} A. M. 561
Leskov{ek V. 67
Lu~i} D. 467
M
Ma~ek M. 627
Male{evi} M. 369
Mamuzi} I. 75
Mandrino Dj. 325
Mansouri K. 125
Marin{ek M. 439, 451
Martínek L. 603, 619
Masarik M. 599
Mauder T. 347, 599
Maver T. 205, 253, 259
Maver U. 205, 259
Mazouzi M. 125
Medveï I. 375
Medved J. 317, 529, 549
Medved S. 241
Mi{ovi} M. 467
Mileti} A. 447
Milo{ev I. 593
Mitrovi} R. 609
Modic M. 233, 281
Mozaiskaja N .V. 21
Mozeti~ M. 185, 457
Mrvar P. 529, 549
Muhamedagi} S. 483
Muratoglu M. 401
N
Nazirudeen S. S. M. 159
Nikbin I. M. 571
Novák J. 619
Novosel B. 439
O
Orel B. 433
Oru~ M. 483
Ozgun S. 105
Ozturk F. 613
P
Palanivel R. 623
Panjan P. 281, 593
Paulin I. 13, 95, 537
Per{in Z. 205, 253, 259
Pirnar B. 339
Pirnat M. 529
Poljan{ek I. 265
Poniku B. 39
Popovi} O. 579
Posset U. 433
Proki} - Cvetkovi} R. 579
Prvulovic S. 131
Puk{i~ N. 55
Pulko I. 247
R
Rai} K. 27, 335
Ramakrishnan S. S. 159
Ranjbar M. M. 571
Rezvani M. 571
Rimac M. 483
Risteski I. B. 503
Rodi~ T. 421
Rom~evi} M. 463, 463
Rudolf R. 335, 463, 561
Rybnikov A. I. 21
S
Sandera C. 347
Seda M. 347
Sedlaøík V. 213, 265
Sedmak A. 579
Seifried S. 247
Sekanina B. 163, 599
Sen~i~ B. 67
Sev{ek U. 247
Skobir Balanti~ D. A. 139, 295, 339,
555, 619
Sobotová J. 489
Socha L. 111, 603, 619
Sochor L. 619
Sokoláø R. 375
Stamenkovi} D. 335, 561
Stana - Kleinschek K. 199, 205, 253,
259
Stetina J. 163, 347, 363, 599
Stransky K. 163, 599
Stropnik ^. 247
[
[alaková A. 265
[arler B. 427, 545
[etina - Bati~ B. 33
[etina J. 33
[ín P. 375
[kapin S. D. 13, 95, 537
[kori} B. 447
[tubòa I. 375
[u{tar{i~ B. 13, 95, 369
[urca Vuk A. 433
[vara Fabjan E. 585
T
Tadi} N. 467
Tadlja K. A. 21
Tanaskovi} T. 335
Tasi} M. 27
Tehovnik F. 151, 339
Ternik P. 335
Tiha~ek [oji} Lj. 561
Todorovi} A. 561
Tolmac D. 131
Tolmac J. 131
Topçu O. 145
Topi} A. 413
Toros S. 613
Traji} J. 463
Trník A. 375
Turk A. 477
U
Übeyli M. 115, 145
Urbánek P. 213
Ustel F. 477
V
Vasiljevi} – Radovi} D. 463
Vesel A. 121, 213, 217, 227, 233
Vodopivec F. 75, 139, 555
Vojvodi~ Tuma J. 357, 369
Von~ina M. 549
Vuk~evi} M. 609
Z
Zaplotnik R. 199, 227, 233
Zejak R. 609
Zem~ík R. 567
Zlati~anin B. 609
Zupan K. 439, 451
Zupani~ F. 593
@
@igon M. 173, 269
@nidar{i~ A. 259
@u`ek B. 139, 339
@uni~ Z. 335
@vegli~ M. 627
Y
Yilmaz F. 477
LETNO KAZALO – INDEX
Materiali in tehnologije / Materials and technology 45 (2011) 6 653
MATERIALI IN TEHNOLOGIJE / MATERIALS AND
TECHNOLOGY
VSEBINSKO KAZALO / SUBJECT INDEX
LETNIK / VOLUME 45, 2011, 1–6
Kovinski materiali – Metallic materials
Experimental study of some masonry-wall coursework material types under horizontal loads and their comparison
Eksperimentalna raziskava zgradbe nekaterih zidarskih zidov – vodoravna obremenitev in primerjava uporabljenih materialov
M. Kamanli, M. S. Donduren, M. T. Cogurcu, M. Altin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3
Synthesis of aluminium foams by the powder-metallurgy process: compacting of precursors
Sinteza aluminijevih pen po postopku metalurgije prahov: stiskanje prekurzorjev
I. Paulin, B. [u{tar{i~, V. Kevorkijan, S. D. [kapin, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13
A new method for determining the remaining lifetime of coated gas-turbine blades
Nova metoda za izra~un preostale trajnostne dobe lopatic plinskih turbin
L. B. Getsov, P. G. Krukovski, N .V. Mozaiskaja, A. I. Rybnikov, K. A. Tadlja . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21
Reduction of ultra-fine tungsten powder with tungsten (VI)-oxide in a vertical tube reactor
Redukcija ultrafinih prahov volframovega(VI) oksida v reaktorju z vertikalno cevjo
@. Kamberovi}, D. Filipovi}, K. Rai}, M. Tasi}, Z. An|i}, M. Gavrilovski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27
The modeling of auger spectra
Modeliranje augerjevih spektrov
B. Poniku, I. Beli~, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 39
Modelling of the directional solidification of a leaded red brass flange
Modeliranje usmerjenega strjevanja prirobnice iz rde~e svin~eve medenine
V. Grozdani} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47
Characterization of the inclusions in spring steel using light microscopy and scanning electron microscopy
Karakterizacija vklju~kov v vzmetnih jeklih s svetlobno in vrsti~no elektronsko mikroskopijo
A. Bytyqi, N. Puk{i~, M. Jenko, M. Godec . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55
Characterization of the carbides in a Ni-Ti shape-memory alloy wire
Karakterizacija karbidov v @ici zlitine s spominom Ni-Ti
M. Godec, A. Kocijan, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 61
Fracture toughness of the vacuum-heat-treated spring steel 51CrV4
Lomna `ilavost vakuumsko toplotno obdelanega vzmetnega jekla 51CrV4
B. Sen~i~, V. Leskov{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 67
Similarity criteria and effect of lubricant inertia at cold rolling
Merila podobnosti in vpliv vztrajnosti maziva pri hladnem valjanju
D. ]ur~ija, F. Vodopivec, I. Mamuzi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75
Material failure of an AISI 316L stainless steel hip prosthesis
Napake materiala v kol~ni protezi iz nerjavnega jekla AISI 316L
M. Godec . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 85
A comparison of the corrosion behaviour of austenitic stainless steels in artificial seawater
Primerjava korozijskih lastnosti avstenitnih nerjavnih jekel v simulirani morski vodi
A. Kocijan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 91
Influence of the foaming precursor’s composition and density on the foaming efficiency, microstructure development and
mechanical properties of aluminium foams
Vpliv sestave in gostote prekurzorjev za penjenje na u~inkovitost penjenja ter razvoj mikrostrukture in mehanskih lastnosti aluminijskih pen
V. Kevorkijan, S. D. [kapin, I. Paulin, B. [u{tar{i~, M. Jenko, M. La`eta. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95
Multi-objective optimization of the cutting forces in turning operations using the Grey-based Taguchi method
Multi namenska optimizacija stru`enja z uporabo Taguchi metode na Grey podlagi
Y. Kazancoglu, U. Esme, M. Bayramo

glu, O. Guven, S. Ozgun . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 105
Thermodynamic conditions for the nucleation of boron compounds during the cooling of steel
Termodinami~ni pogoji za nukleacijo borovih spojin pri ohlajanju jekla
Z. Adolf, J. Ba`an, L. Socha . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111
LETNO KAZALO – INDEX
654 Materiali in tehnologije / Materials and technology 45 (2011) 6
A micro-damage investigation on a low-alloy steel tested using a 7.62-mm AP projectile
Raziskava mikropo{kodb malolegiranega jekla po preizkusu s kroglo AP 7,62 mm
T. Demir, M. Übeyli . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 115
Creep resistance of microstructure of welds of creep resistant steels
Odpornost proti lezenju pri mikrostrukturi zvarov jekel, odpornih proti lezenju
F. Vodopivec, M. Jenko, R. Celin, B. @u`ek, D. A. Skobir . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 139
Effect of the martensite volume fraction on the machining of a dual-phase steel using a milling operation
Vpliv volumenskega dele`a martenzita na obdelavo dvofaznega dualnega jekla z rezkanjem
O. Topçu, M. Übeyli, A. Acir . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 145
Degradation of a Ni-Cr-Fe alloy in a pressurised-water nuclear power plant
Degradacija zlitin Ni-Cr-Fe v tla~novodnih jedrskih elektrarnah
R. Celin, F. Tehovnik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 151
Investigation into the mechanical properties of micro-alloyed as-cast steel
Raziskave mehanskih lastnosti mikrolegiranih jekel
B. Chokkalingam, S. S. M. Nazirudeen, S. S. Ramakrishnan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 159
The effect of electromagnetic stirring on the crystallization of concast billets
Kristalizacija kontinuirno ulitih gredic v elektromagnetnem polju
K. Stransky, F. Kavicka, B. Sekanina, J. Stetina,V. Gontarev, J. Dobrovska. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 163
High-strength low-alloy (HSLA) steels
Visokotrdna malolegirana (HSLA) konstrukcijska jekla
D. A. Skobir . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 295
TEM investigation of metallic materials – an advanced technique in materials science and metallurgy
Preiskave kovinskih materialov s presevno elektronsko mikroskopijo – moderna tehnika v znanosti o materialih in metalurgiji
D. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 303
Modelling of hot tears in continuously cast steel
Modeliranje vro~ih razpok v kontinuirno litem jeklu
V. Grozdani} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 311
Thermodynamic investigation of the Al-Sb-Zn system
Termodinamska raziskava sistema Al-Sb-Zn
G. Klan~nik, J. Medved. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 317
XPS and SEM of unpolished and polished FeS surface
Rentgenska fotoelektronska spektroskopija in vrsti~na elektronska mikroskopija nepolirane in polirane povr{ine FeS
Dj. Mandrino . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 325
Surface characterization and pickling characteristics of the oxide scale on duplex stainless steel
Povr{inska karakterizacija in lastnosti lu`enja oksidne plasti na dupleksnem nerjavnem jeklu
^. Donik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 329
CFD analysis of exothermic reactions in Al-Au nano multi-layered foils
CFD-analiza eksotermnih reakcij v ve~plastnih nanofolijah Al-Au
K. T. Rai}, R. Rudolf, P. Ternik, Z. @uni~, V. Lazi}, D. Stamenkovi}, T. Tanaskovi}, I. An`el . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 335
Microstructure evolution in SAF 2507 super duplex stainless steel
Razvoj mikrostrukture v superdupleksnem nerjavnem jeklu SAF 2507
F. Tehovnik, B. Arzen{ek, B. Arh, D. Skobir, B. Pirnar, B. @u`ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 339
Optimization of the quality of continuously cast steel slabs using the Firefly algorithm
Optimizacija kakovosti kontinuirno lite jeklene plo{~e z uporabo algoritma "Firefly"
T. Mauder, C. Sandera, J. Stetina, M. Seda . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 347
Hot workability of 95MnWCr5 tool steel
Vro~a preoblikovalnost orodnega jekla 95MnWCr5
A. Kri`aj, M. Fazarinc, M. Jenko, P. Fajfar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 351
Lifetime evaluation of a steam pipeline using NDE methods
Ocena preostale trajnostne dobe parovoda z uporabo neporu{itvenih preiskav (NDE)
F. Kafexhiu, J. Vojvodi~ Tuma. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 357
The influence of the chemical composition of steels on the numerical simulation of a continuously cast slab
Vpliv kemi~ne sestave jekel na numeri~no simulacijo kontinuirno lite plo{~e
J. Stetina, F. Kavicka. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 363
Prediction of the mechanical properties of cast Cr-Ni-Mo stainless steels with a two-phase microstructure
Napoved mehanskih lastnosti litih Cr-Ni-Mo nerjavnih jekel z dvofazno mikrostrukturo
M. Male{evi}, J. V. Tuma, B. [u{tar{i~, P. Borkovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 369
LETNO KAZALO – INDEX
Materiali in tehnologije / Materials and technology 45 (2011) 6 655
Relationship between mechanical strength and Young’s modulus in traditional ceramics
Odvisnost med mehansko trdnostjo in Youngovim modulom pri tradicionalni keramiki
I. [tubòa, A. Trník, P. [ín, R. Sokoláø, I. Medveï . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 375
Cr-V ledeburitic cold-work tool steels
Ledeburitna jekla Cr-V za delo v hladnem
P. Jur~i . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 383
Experimental comparison of resistance spot welding and friction-stir spot welding processes for the EN AW 5005 aluminum
alloy
Eksperimentalna primerjava odpornosti procesov to~kovnega varjenja in to~kovnega tornega varjenja pri aluminijevi zlitini EN AW 5005
M. K. Kulekci, U. Esme, O. Er . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 395
The friction and wear behavior of Cu-Ni3Al composites by dry sliding
Trenje in obraba Cu-Ni3Al kompozitov pri suhem drsenju
M. Demirel, M. Muratoglu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 401
Weldability of metal matrix composite plates by friction stir welding at low welding parameters
Varivost plo{~ kompozita s kovinsko osnovo po vrtilno tornem postopku pri nizkih varilnih parametrih
Y. Bozkurt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 407
Influence of the gas composition on the geometry of laser-welded joints in duplex stainless steel
Vpliv vrste za{~itnega plina na geometrijo zvara pri laserskem varjenju nerjavnega dupleksnega jekla
B. Bauer, A. Topi}, S. Kralj, Z. Ko`uh . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 413
Multiscale modelling of heterogeneous materials
Mikro in makro modeliranje heterogenih materialov
M. Lamut, J. Korelc, T. Rodi~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 421
Genetic programming and soft-annealing productivity
Genetsko programiranje in produktivnost mehkega `arjenja
M. Kova~i~, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 427
Optical properties of plastically deformed copper: an ellipsometric study
Opti~ne lastnosti plasti~no derformiranega bakra: {tudij elipsometrije
N. Rom~evi}, R. Rudolf, J. Traji}, M. Rom~evi}, B. Had`i}, D. Vasiljevi} – Radovi}, I. An`el . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 463
Relaxation of the residual stresses produced by plastic deformation
Relaksacija zaostalih napetosti zaradi plasti~ne deformacije
N. Tadi}, M. Jeli}, D. Lu~i}, M. Mi{ovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 467
Accelerated corrosion behaviors of Zn, Al and Zn/15Al coatings on a steel surface
Pospe{eno korozijsko obna{anje Zn, Al in Zn/15Al prekritij na povr{ini jekla
A. Gulec, O. Cevher, A. Turk, F. Ustel, F. Yilmaz . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 477
Alloys with modified characteristics
Zlitine z modificiranimi lastnostmi
M. Oru~, M. Rimac, O. Beganovi}, S. Muhamedagi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 483
Evaluation of the microstructural changes in Cr-V ledeburitic tool steels depending on the austenitization temperature
Ocena sprememb mikrostrukture v ledeburitnemn orodnem jeklu Cr-V v odvisnosti od temperature avstenitizacije
P. Bílek, J. Sobotová, P. Jur~i . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 489
Characteristics of creep in conditions of long operation
Zna~ilnosti lezenja pri dolgotrajni uporabi
N. A. Katanaha, L. B. Getsov . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 523
A thermodynamic and kinetic study of the solidification and decarburization of malleable cast iron
Termodinami~na in kineti~na analiza strjevanja in razoglji~enja belega litega `eleza
M. Pirnat, P. Mrvar, J. Medved . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 529
Modelling and preparation of core foamed Al panels with accumulative hot-roll bonded precursors
Na~rtovanje in izdelava Al-panelov s sredico iz Al-pen na osnovi ve~stopenjsko toplo valjanih prekurzorjev
V. Kevorkijan, U. Kova~ec, I. Paulin, S. D. [kapin, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 537
Numerical solution of hot shape rolling of steel
Numeri~na re{itev vro~ega valjanja jekla
U. Hanoglu, Siraj-ul-Islam, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 545
Solidification and precipitation behaviour in the AlSi9Cu3 alloy with various Ce additions
Strjevanje in izlo~anje v zlitini ALSI9CU3 pri razli~nih dodatkih Ce
M. Von~ina, S. Kores, P. Mrvar, J. Medved . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 549
Effect of change of carbide particles spacing and distribution on creep rate of martensite creep resistant steels
Vpliv spremembe razdalje med karbidnimi izlo~ki in njihove porazdelitve na hitrost lezenja martenzitnih jekel, odpornih proti lezenju
D. A. Skobir Balanti~, M. Jenko, F. Vodopivec, R. Celin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 555
LETNO KAZALO – INDEX
656 Materiali in tehnologije / Materials and technology 45 (2011) 6
Stress-strain analysis of an abutment tooth with rest seat prepared in a composite restoration
Napetostno-deformacijska analiza opornega zoba z zapornim sede`em, izdelana s kompozitnim popravilom
Lj. Tiha~ek [oji}, A. M. Lemi}, D. Stamenkovi}, V. Lazi}, R. Rudolf, A. Todorovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 561
The influence of buffer layer on the properties of surface welded joint of high-carbon steel
Vpliv vmesne plasti na lastnosti povr{inskih zvarov jekla z veliko ogljika
O. Popovi}, R. Proki} - Cvetkovi}, A. Sedmak, G. Buyukyildrim, A. Bukvi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 579
Corrosion stability of different bronzes in simulated urban rain
Korozijska stabilnost razli~nih bronov v umetnem kislem de`ju
E. [vara Fabjan, T. Kosec, V. Kuhar, A. Legat . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 585
Morphology and corrosion properties PVD Cr-N coatings deposited on aluminium alloys
Morfologija in korozijske lastnosti CrN PVD-prevlek, nanesenih na aluminijeve zlitine
D. Kek Merl, I. Milo{ev, P. Panjan, F. Zupani~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 593
The effect of electromagnetic stirring on the crystallization of concast billets
Vpliv elektromagnetnega me{anja na kristalizacijo kontinuirno ulitih gredic
F. Kavicka, K. Stransky, B. Sekanina, J. Stetina, V. Gontarev, T. Mauder, M. Masarik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 599
Effect of pre-straining on the springback behavior of the AA5754-0 alloy
Vpliv prenapenjanja na povratno elasti~no izravnavo zlitine AA5754-0
S. Toros, M. Alkan, R. Ecmel Ece, F. Ozturk. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 613
Heat treatment and mechanical properties of heavy forgings from A694–F60 steel
Toplotna obdelava in mehanske lastnosti te`kih izkovkov iz jekla A694-F60
M. Balcar, J. Novák, L. Sochor, P. Fila, L. Martínek, J. Ba`an, L. Socha, D. A. Skobir Balanti~, M. Godec . . . . . . . . . . . . . . . . . . . . . . 619
The tensile behaviour of friction-stir- welded dissimilar aluminium alloys
Natezne zna~ilnosti tornih pomi~nih zvarov razli~nih aluminijevih zlitin
R. Palanivel, P. Koshy Mathews. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 623
Recent growing demand for magnesium in the automotive industry
Rast povpra{evanja po magneziju v avtomobilski industriji
M. J. Freiría Gándara . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 633
Contact with chlorinated water: selection of the appropriate steel
Kontakt s klorirano vodo – izbor ustreznega jekla
L. Gosar, D. Drev . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 639
Anorganski materiali – Inorganic materials
Modification of surface morphology of graphite by oxygen plasma treatment
Sprememba morfologije grafita med obdelavo s kisikovo plazmo
K. Eler{i~, I. Junkar, M. Modic, R. Zaplotnik, A. Vesel, U. Cvelbar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 233
Properties of particleboards made by using an adhesive with added liquefied wood
Lastnosti ivernih plo{~, izdelanih z uporabo lepila z dodanim uteko~injenim lesom
N. ^uk, M. Kunaver, S. Medved . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 241
Hydrothermal growth of Zn5(OH)6(CO3)2 and its thermal transformation into porous ZnO film used for dye-sensitized solar
cells
Hidrotermalna rast Zn5(OH)6(CO3)2 s termi~no transformacijo v porozno plast ZnO, uporabljeno za elektrokemijske son~ne celice
M. Bitenc, Z. Crnjak Orel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 287
Semi-solid gel electrolytes for electrochromic devices
Poltrdni gelski elektroliti za elektrokromne naprave
M. Hajzeri, M. ^olovi}, A. [urca Vuk, U. Posset, B. Orel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 433
Combustible precursor behaviour in the lanthanum chromite formation process
Termi~ne lastnosti reakcijskega gela za pripravo lantanovega kromita
K. Zupan, M. Marin{ek, B. Novosel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 439
Influence of the granulation and grain shape of quartz sands on the quality of foundry cores
Vpliv granulacije in oblike zrn kremenovega peska na kakovost livarskih jeder
M. Marin{ek, K. Zupan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 451
Wear of refractory materials for ceramic filters of different porosity in contact with hot metal
Obraba ognjevzdr`nega materiala kerami~nih filtrov z razli~no poroznostjo v stiku z vro~o kovino
J. Ba`an, L. Socha, L. Martínek, P. Fila, M. Balcar, J. Chmelaø . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 603
The influence of the mineral content of clay from the white bauxite mine on the properties of the sintered product
Vpliv vsebnosti minerala gline iz rudnika belega boksita na lastnosti sintranega proizvoda
M. Krgovi}, I. Bo{kovi}, M. Vuk~evi}, R. Zejak, M. Kne`evi}, R. Mitrovi}, B. Zlati~anin, N. Ja}imovi} . . . . . . . . . . . . . . . . . . . . . . . . 609
LETNO KAZALO – INDEX
Materiali in tehnologije / Materials and technology 45 (2011) 6 657
Screen-printed electrically conductive functionalities in paper substrates
Elektroprevodne oblike, pripravljene s sitotiskom na papirnih podlagah
M. @vegli~, N. Hauptman, M. Ma~ek, M. Klanj{ek Gunde . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 627
Polimeri – Polymers
Activation of polymer polyethylene terephthalate (PET) by exposure to CO2 and O2 plasma
Aktivacija polimera polietilentereftalata (PET) s CO2- ali O2-plazmo
A. Vesel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 121
The impact on rigid PVC pipes: a study of the correlation between the length of the crazed zone and the area of the impacted
region
Udar togih PVC-cevi: {tudija korelacije med dol`ino razpokane zone in povr{ino zone udara
C. B. Fokam, M. Chergui, K. Mansouri, M. El Ghorba, M. Mazouzi . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 125
Surface characterization of polymers by XPS and SIMS techniques
Analiza povr{ine polimerov z metodama XPS in SIMS
J. Kova~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 191
Modification of PET-polymer surface by nitrogen plasma
Modifikacija povr{ine PET-polimera z du{ikovo plazmo
R. Zaplotnik, M. Kolar, A. Doli{ka, K. Stana-Kleinschek. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 199
Functionalization of AFM tips for use in force spectroscopy between polymers and model surfaces
Funkcionalizacija AFM-konic za uporabo v spektroskopiji sil med polimeri in modelnimi povr{inami
T. Maver, K. Stana - Kleinschek, Z. Per{in, U. Maver . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 205
A novel approach for qualitative determination of residual tin based catalyst in poly(lactic acid) by X-ray photoelectron
spetroscopy
Nov na~in kvalitativne dolo~itve vsebnosti katalizatorja kositra v polilakti~ni kislini z rentgensko fotoelektronsko spektroskopijo
V. Sedlaøík, A. Vesel, P. Kucharczyk, P. Urbánek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 213
Poly(styrene-CO-divinylbenzene-CO-2-ethylhexyl)acrylate membranes with interconnected macroporous structure
Poli(stiren-KO-divinilbenzen-KO-2-etilheksil)akrilatne membrane s povezano porozno strukturo
U. Sev{ek, S. Seifried, ^. Stropnik, I. Pulko, P. Krajnc . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 247
Use of AFM force spectroscopy for assessment of polymer response to conditions similar to the wound, during healing
Uporaba AFM-spektroskopije sil za spremljanje odziva polimernih molekul na v rani podobna okolja med celjenjem
U. Maver, T. Maver, A. @nidar{i~, Z. Per{in, M. Gaber{~ek, K. Stana-Kleinschek. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 259
Biodegradable polymers from renewable resources: effect of proteinic impurityon polycondensation products of
2-hydroxypropanoic acid
Biorazgradljivi polimeri iz obnovljivih virov: vpliv proteinskih ne~istot na produkte polikondenzacije 2-hidroksipropanojske kisline
I. Poljan{ek, P. Kucharczyk,V. Sedlaøík, V. Ka{párková, A. [alaková, J. Drbohlav . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 265
Identification and verification of the composite material parameters for the Ladevèze damage model
Identifikacija in verifikacija parametrov kompozitnega materiala za model Ladevèze
V. Kleisner, R. Zem~ík, T. Kroupa. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 567
Vakuumska tehnika – Vacuum technique
Oxygen diffusion in the non-evaporable getter St 707 during heat treatment
Difuzija kisika v getru St 707 med toplotno obdelavo
S. Avdiaj, B. [etina - Bati~, J. [etina, B. Erjavec . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33
Removal of a thin hydrogenated carbon film by oxygen plasma treatment
Odstranjevanje tanke plasti hidrogeniranega ogljika s kisikovo plazmo
U. Cvelbar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 179
Low temperature destruction of bacteria Bacillus stearothermophilus by weakly ionized oxygen plasma
Nizkotemperaturno uni~evanje bakterij Bacillus stearothermophilus s {ibko ionizirano kisikovo plazmo
M. Mozeti~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 185
Hydrophobization of polymer polystyrene in fluorine plasma
Hidrofobizacija polimera polistiren s fluorovo plazmo
A. Vesel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 217
Plasma treatment of biomedical materials
Plazemska obdelava biomedicinskih materialov
I. Junkar, U. Cvelbar, M. Lehocky . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 221
Radiofrequency induced plasma in large-scale plasma reactor
Radiofrekven~no inducirana plazma v reaktorju velikih dimenzij
R. Zaplotnik, A. Vesel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 227
LETNO KAZALO – INDEX
658 Materiali in tehnologije / Materials and technology 45 (2011) 6
Modification of non-woven cellulose for medical applications using non-equlibrium gassious plasma
Modifikacija celuloznih kopren, uporabnih v medicinske namene, z neravnovesno plinsko plazmo
K. Stana - Kleinschek, Z. Per{in, T. Maver . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 253
Tuning of poly(ethylene terephtalate) (PET) surface properties by oxygen plasma treatment
Prilagoditev lastnosti povr{ine polietilen tereftalata (PET) z obelavo v kisikovi plazmi
A. Doli{ka, M. Kolar. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 275
Probability of recombination and oxidation of O atoms on a-C:H surface
Verjetnost za rekombinacijo in oksidacijo za atome kisika na povr{ini a-C:H
A. Drenik, K. Eler{i~, M. Modic, P. Panjan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 281
Characterization of extremely weakly ionized hydrogen plasma with a double Langmuir probe
Karakterizacija {ibko ionizirane vodikove plazme z dvojno Langmuirjevo sondo
M. Mozeti~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 457
Kemija – Chemistry
New discovered paradoxes in theory of balancing chemical reactions
Novoodkriti paradoksi v teoriji uravnote`enja kemijskih reakcij
I. B. Risteski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 503
Kemijska tehnologija – Chemical technology
A comparative analysis of theoretical models and experimental research for spray drying
Primerjalna analiza teoreti~nih modelov in eksperimentalna raziskava razpr{ilnega su{enja
D. Tolmac, S. Prvulovic, D. Dimitrijevic, J. Tolmac . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 131
Nanomateriali in nanotehnologije – Nanomaterials and nanotechnology
Microwave-assisted non-aqueous synthesis of ZnO nanoparticles
Sinteza nanodelcev ZnO v nevodnem mediju pod vplivom mikrovalov
G. Ambro`i~, Z. Crnjak Orel, M. @igon. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 173
Sub micrometer and nano ZnO as filler in PMMA materials
Submikrometrski in nano ZnO kot polnilo v PMMA-materialih
A. An`lovar, Z. Crnjak Orel, M. @igon . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 269
Nanoscale modification of hard coatings with ion implantation
Nanovelikostna modifikacija trdnih prekritij z ionsko implantacijo
B. [kori}, D. Kaka{, M. Gostimirovi}, A. Mileti} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 447
Gradbeni material – Materials in civil engineering
Evaluation of the strength variation of normal and lightweight self-compacting concrete in full scale walls
Ocena variacije trdnosti normalnega in lahkega vibriranega betona v polnih stenah
M. M. Ranjbar, M. Hosseinali Beygi, I. M. Nikbin , M. Rezvani, A. Barari . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 571
Osebne biografije – Personal biographies
In memoriam:
Oskar Kürner 1925–2010 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 81
Franc Vodopivec – osemdesetletnik
Laudation in honour of Franc Vodopivec on the occasion of his 80th birthday
M. Jenko. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 381
In memoriam:
Hans Jürgen Grabke . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 495
LETNO KAZALO – INDEX
Materiali in tehnologije / Materials and technology 45 (2011) 6 659