VSEBINA – CONTENTS
PREGLEDNI ^LANEK – REVIEW ARTICLE
Problems and normative evaluation of bond-strength tests for coated reinforcement and concrete
Problemi in normativna ocena preizkusov trdnosti vezi med armaturo s prekritjem in betonom
P. Pokorný, M. Kouøil, J. Stoulil, P. Bou{ka, P. Simon, P. Juránek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 847
IZVIRNI ZNANSTVENI ^LANKI – ORIGINAL SCIENTIFIC ARTICLES
Modeling of occurrence of surface defects of C45 steel with genetic programming
Modeliranje pojava povr{inskih napak pri jeklu C45 z genetskim programiranjem
M. Kova~i~, R. Jager . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 857
Effects of various helically angled grinding wheels on the surface roughness and roundness in grinding cylindrical surfaces
Vpliv razli~nih kotov vija~nice pri brusilnih kolutih na hrapavost povr{ine in okroglost pri bru{enju valjastih povr{in
M. Gavas, M. Kýna, U. Köklü . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 865
Characterization of cast-iron gradient castings
Karakterizacija lito`eleznega gradientnega ulitka
D. Mitrovi}, P. Mrvar, M. Petri~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 871
Comparative mechanical and corrosion studies on magnesium, zinc and iron alloys as biodegradable metals
Primerjalna {tudija mehanskih in korozijskih lastnosti biorazgradljivih zlitin magnezija, cinka in `eleza
D. Vojtìch, J. Kubásek, J. ^apek, I. Pospí{ilová. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 877
Microstructural changes of fine-grained concrete exposed to a sulfate attack
Mikrostrukturne spremembe drobnozrnatega betona, izpostavljenega sulfatu
M. Vy{vaøil, P. Bayer, M. Rovnaníková . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 883
Effect of thermomechanical treatment on the corrosion behaviour of Si- and Al-containing high-Mn austenitic steel with Nb
and Ti micro-additions
Vpliv termomehanske obdelave na korozijsko vedenje manganskega avstenitnega jekla z vsebnostjo Si in Al, mikrolegiranega z Nb in Ti
A. Grajcar, A. Plachciñska, S. Topolska, M. Kciuk . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 889
Surface free energy of hydrophobic coatings of hybrid-fiber-reinforced high-performance concrete
Prosta energija povr{ine hidrofobnih premazov na visokozmogljivem betonu, oja~anem s hibridnimi vlakni
D. Barnat-Hunek, P. Smarzewski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 895
Developing continuous-casting-process control based on advanced mathematical modelling
Uporaba naprednega matemati~nega modeliranja za razvoj kontrole postopka kontinuirnega ulivanja
J. Falkus, K. Mi³kowska-Piszczek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 903
Elastic behaviour of magnesia-chrome refractories at elevated temperatures
Elasti~no vedenje ognjevzdr`nih gradiv magnezija-krom pri povi{anih temperaturah
I. Jastrzêbska, J. Szczerba, J. Szlêzak, E. Œnie¿ek, Z. Pêdzich . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 913
Study on the magnetization-reversal behavior of annealed Sm-Fe-Co-Si-Cu ribbons
[tudij vedenja pri obratu magnetizacije `arjenih trakov Sm-Fe-Co-Si-Cu
M. Doœpial, S. Garus, M. Nabialek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 919
Evaluation of the chitosan-coating effectiveness on a dental titanium alloy in terms of microbial and fibroblastic attachment
and the effect of aging
Ocena u~inkovitosti nanosa hitozana na oprijemanje mikrobov in fibroblastov na dentalni titanovi zlitini ter na pojav staranja
U. T. Kalyoncuoglu, B. Yilmaz, S. Gungor, Z. Evis, P. Uyar, G. Akca, G. Kansu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 925
Structure and properties of the carburised surface layer on 35CrSiMn5-5-4 steel after nanostructurization treatment
Struktura in lastnosti naoglji~ene povr{ine jekla 35CrSiMn5-5-4 po nanostrukturni obdelavi
E. Sko³ek, K. Wasiak, W. A. Œwi¹tnicki. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 933
Optimization of the surface roughness by applying the Taguchi technique for the turning of stainless steel under cooling
conditions
Uporaba Taguchi-jeve metode za optimiranje hrapavosti povr{ine pri stru`enju nerjavnega jekla z ohlajanjem
M. Sarýkaya . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 941
ISSN 1580-2949
UDK 669+666+678+53 MTAEC9, 49(6)845–1018(2015)
MATER.
TEHNOL.
LETNIK
VOLUME 49
[TEV.
NO. 6
STR.
P. 845–1018
LJUBLJANA
SLOVENIJA
NOV.–DEC.
2015
Effect of cryogenic treatment applied to M42 HSS drills on the machinability of Ti-6Al-4V alloy
Vpliv podhlajevanja svedrov M42 HSS na obdelovalnost zlitine Ti-6Al-4V
T. Kývak, U. ªeker . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 949
Load-capacity prediction for the carbon- or glass-fibre-reinforced plastic part of a wrapped pin joint
Napoved nosilnosti plasti~nih delov zati~nega spoja, oja~anega z ogljikovimi ali steklastimi vlakni
J. Krystek, R. Kottner . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 957
A meshless model of electromagnetic braking for the continuous casting of steel
Brezmre`ni model elektromagnetnega zaviranja pri kontinuiranem ulivanju jekla
K. Mramor, R. Vertnik, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 961
Non-singular method of fundamental solutions for three–dimensional isotropic elasticity problems with displacement boundary
conditions
Nesingularna metoda fundamentalnih re{itev za deformacijo tridimenzijskih elasti~nih problemov z deformacijskimi robnimi pogoji
Q. Liu, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 969
Solid-state sintering of (K0.5Na0.5)NbO3 synthesized from an alkali-carbonate-based low-temperature calcined powder
Sintranje v trdnem keramike (K0,5Na0,5)NbO3, sintetizirane iz nizkotemperaturno kalciniranega prahu, pripravljenega na osnovi alkalijskih
karbonatov
M. Feizpour, T. Ebadzadeh, D. Jenko. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 975
Stability of close-cell Al foams depending on the usage of different foaming agents
Stabilnost aluminijevih pen z zaprto poroznostjo glede na uporabo razli~nih penilnih sredstev
I. Paulin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 983
STROKOVNI ^LANEK – PROFESSIONAL ARTICLE
Evolution of the microstructure and magnetic properties of a cobalt-silicon-based alloy in the early stages of mechanical
milling
Razvoj mikrostrukture in magnetnih lastnosti zlitine Co-Si v za~etnem stadiju mehanskega legiranja
W. Rattanasakulthong, C. Sirisathitkul, P. F. Rogl . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 989
IN MEMORIAM
Prof. dr. Milan Trbi`an . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 993
LETNO KAZALO – INDEX
Letnik 49 (2015), 1–6 – Volume 49 (2015), 1–6 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 995
P. POKORNÝ et al.: PROBLEMS AND NORMATIVE EVALUATION OF BOND-STRENGTH TESTS ...
847–856
PROBLEMS AND NORMATIVE EVALUATION OF
BOND-STRENGTH TESTS FOR COATED REINFORCEMENT AND
CONCRETE
PROBLEMI IN NORMATIVNA OCENA PREIZKUSOV TRDNOSTI
VEZI MED ARMATURO S PREKRITJEM IN BETONOM
Petr Pokorný1, Milan Kouøil1, Jan Stoulil1, Petr Bou{ka2, Pavel Simon3,
Pavel Juránek4
1Institute of Chemical Technology in Prague, Department of Metals and Corrosion Engineering, Technická 5, 166 28 Prague 6, Czech
Republic
2Czech Technical University in Prague, Klokner Institute, [olínova 7, 166 08 Prague 6, Czech Republic
3Ing. Vladimír Fi{er, Mlýnská 68, 602 00 Brno, Czech Republic
4Technical and Test Institute for Constructions Prague – Brno Branch, Hnìvkovského 77, 617 00 Brno, Czech Republic
Pokornyt@vscht.cz
Prejem rokopisa – received: 2014-09-13; sprejem za objavo – accepted for publication: 2015-01-05
doi:10.17222/mit.2014.227
This paper focuses on the problems of the bond-strength between concrete and reinforcement and defines the basic factors
affecting the quality of the bond. Two types of coated concrete reinforcement (the zinc- and epoxy-coated) and methods of
testing their bond-strength with concrete are described. The goal of this work is to generalize the results of the bond-strength
tests so that they would consider only the influence of the corrosion of the zinc-coated reinforcement in fresh concrete or, in the
case of the epoxy-coated reinforcement, its probable constriction during the testing. Based on described standards, it is
recommended to use the pull-out test to obtain these generalized results: two Czech standards (Bond-strength test of the
reinforcement cast in prisms, Beam-strength test of the reinforcement in cubes) and a RILEM recommendation.
Keywords: concrete, corrosion of concrete reinforcement, bond-strength, bond-strength tests, hot-dip galvanized reinforcement,
epoxy-coated reinforcement, standards
^lanek obravnava probleme trdnosti vezi med betonom in armaturo ter dolo~a osnovne faktorje, ki vplivajo na kvaliteto vezi.
Opisani sta dve vrsti prekritja armature (prekritje s cinkom in prekritje z epoksi smolo) in kako se preizkusi njihova trdnost vezi
z betonom. Cilj tega dela je bil dobiti rezultate preizkusov trdnosti vezi, ki bi upo{tevali samo vpliv korozije pri armaturi, pokriti
s cinkom v sve`em betonu ali, v primeru armature, pokrite z epoksi smolo, na njeno zo`enje med preizkusom. Na podlagi
opisanih standardov se priporo~a preizkus izpuljenja, da se dobi posplo{ene rezultate: dva ~e{ka standarda (Preizkus trdnosti
vezi betona ulitega v prizme, Preizkus trdnosti betonske kocke) in RILEM-priporo~ilo.
Klju~ne besede: beton, korozija armature v betonu, trdnost vezi, preizkusi trdnosti vezi, vro~e cinkana armatura, armatura z
nanosom epoksi smole, standardi
1 INTRODUCTION
The durability of reinforced-concrete structures is
always limited by the corrosion of carbon-steel rein-
forcement. The volume of corrosion products is signifi-
cantly larger (2–6-times) than the volume of the original
non-corroded reinforcements, thus creating a stress
leading to the formation of cracks in the initial stages of
concrete solidification, eventually resulting in a disinte-
gration of the concrete cover. The initiation of reinforce-
ment corrosion corresponds to the carbonation of the
concrete cover by CO2 (the carbon steel becomes active
due to a pH reduction of the alkaline pore solution)
and/or, more frequently, a local attack by penetrating
chlorides (deicing salt, seawater). A disintegration of the
concrete cover propagates the attack to the other parts of
the reinforcement. A reduction of the reinforcement
diameter by the corrosion poses a significant threat to the
static function of the construction. The whole corroded
reinforcement often needs to be replaced at great
expenses to repair the construction.1–3
The prolongation of the longevity of concrete rein-
forcement basically falls into two categories. The first
one involves a change in the concrete properties to in-
crease the compactness of the concrete cover (the con-
cretes with a low water-cement ratio, a better concrete
densification, final application of various concrete plasti-
cizers or other changes to the concrete composition).
Special cements with a suitable ash substituting the
cement provide another way to increase the compactness
– these concretes can have higher mechanical properties,
a lower inherited porosity and, therefore, lower pene-
tration of water, oxygen and corrosion stimulators.4 A
surface modification with a barrier effect slowing down
the penetration of carbon dioxide and chlorides such as
paint or an organosilane surface modification (the sur-
face becomes hydrophobic) are also not negligible.
The second category focuses on the reinforcement.
Its examples are corrosion inhibitors, a cathodic protec-
tion, the coating of the reinforcement or application of a
Materiali in tehnologije / Materials and technology 49 (2015) 6, 847–856 847
UDK 691:620.1:691.3:620.197 ISSN 1580-2949
Review article/Pregledni ~lanek MTAEC9, 49(6)847(2015)
reinforcement from another material (stainless steel,
composite reinforcement).5,6
A significant advantage of stainless steel, compared
to carbon steel, is its high resistance to a low-pH pore
solution. Its resistance to chloride-containing environ-
ment depends on the composition of the steel – certain
types are prone to a localized corrosion attack. Never-
theless, the critical chloride concentration causing the
activation of steel can be up to 15× higher than that for
carbon steel.7,8 Similarly to carbon steel, the resistance of
stainless steel is limited by the state of its surface. The
scales formed during hot-working or welding have a
strong detrimental effect on its corrosion resistance. It
was discovered that even an increase in the pH from 12.5
to 13.5 improves the corrosion resistance less than the
removal of the surface.9
The surface finish of a steel reinforcement can
enhance the corrosion resistance of the reinforcement in
a concrete environment while the core of the reinforce-
ment maintains all of its necessary mechanical properties
(weldability, tensile and compressive strengths, fatigue
strength, etc.). Currently, the feasibility of hot-dip
galvanized coating and powder-plastic coating is being
discussed. However, these kinds of reinforcement protec-
tion cannot be employed until the bond between the
concrete and these new surface-modified reinforcements
has been thoroughly studied.10
2 CONCRETE-REINFORCEMENT BOND-
STRENGTH
A perfect and permanent bond between all steel-rein-
forcement components and concrete is the basic require-
ment for the static cooperation of both materials. The
quality of the bond depends on their reciprocal cohesion,
the bond durability corresponds to the similarity of their
thermal-expansion coefficients and the corrosion resis-
tance of the reinforcement material. Different thermal-
expansion coefficients cause both materials to behave
differently during temperature changes, thus negatively
affecting their bond.11
The concrete-reinforcement bond is generally a
combination of all the factors affecting the movement of
reinforcement during a transformation of a concrete
structures reinforced by steel. It is thus important for the
reinforcement components to change similarly to the
concrete during loading and prevent their movement.11,12
The total bond-strength (Bs) between concrete and its
reinforcement is a combination of three factors from the
bond-strength formula: adhesion factor fad which in-
cludes the effect of the small surface defects of the
reinforcement, friction factor ff which takes into account
small surface unevenness of the reinforcement causing
friction and, finally, mechanical bonding factor fmech
which includes the effect of the surface geometry (ribs,
imprints, warping, etc.). The bond-strength formula is
written as Equation (1):13
Bs = fad + ff + fmech (1)
The factors mentioned above do not have even effects
on the bond-strength. The mechanical bonding factor has
the greatest effect on the bond-strength. The reason for
this is the fact the above factors also necessarily include
the effects of the mechanical properties of concrete – its
local hardness (included in the friction factor) and com-
pressive strength (included in the mechanical bonding
factor). The adhesion factor is strongly affected by the
concrete porosity – a high porosity decreases the effect
of the physicochemical interactions at the interface
which are always short-range. The bond is formed by
hydrating concrete penetrating the reinforcement-surface
defects, creating a mechanical bond. The mechanical
bond becomes strengthened, to some extent, during the
concrete’s aging due to its constriction around the rein-
forcement.12,13
The tensile forces of the reinforcement must be
transferred to the concrete in the less-stressed areas and
the reinforcement must be well bound. Bond length lb is
defined as the length of the reinforcement inside the
concrete necessary for the reinforcement to crack (during
pulling) instead of being torn out of the concrete. Tan-
gential stress (b) occurs on the surface of the reinforce-
ment – the force (F) is distributed unevenly along the
reinforcement; for the sake of simplicity, the average
value is typically used in structures design. This value is
given by Equation (2) (u is the rod diameter, lb is the
length of the rod set in concrete):
 b
b
=
F
ul
(2)
Considering the bond-strength, we can define the
limit bond length according to Equation (3) (fyk is the
characteristic yield strength of the reinforcement rod, 
is the nominal diameter of the reinforcement rod, fbd is
the design value of the ultimate bond stress):
l
f
f
y
b
k
bd
=
Φ
4
(3)
The key to correctly calculate the sufficient bond
length is the reinforcement-concrete bond-strength. In
mathematical models, the bond-strength is represented
by the design value of the ultimate bond stress fbd which
can be calculated, for ribbed reinforcements, from Equa-
tion (4) (1 is the factor including the aging conditions
and the position of the reinforcement rod during the
casting of concrete, 2 is the factor including the
normative diameter of the reinforcement rod, fctd is the
design tensile strength that should not exceed the value
set for the C60/75 concrete strength class):12
f fbd ctd= 2 25 1 2.   (4)
When assessing the effect of concrete on the strength
of its bond with the reinforcement, it is generally im-
portant for the cement content in the concrete to be high,
since the hydrating cement must adhere well to the rein-
P. POKORNÝ et al.: PROBLEMS AND NORMATIVE EVALUATION OF BOND-STRENGTH TESTS ...
848 Materiali in tehnologije / Materials and technology 49 (2015) 6, 847–856
forcement rods. Concrete must also be well compacted.
Single rods must be covered by concrete on all sides and
the minimum cover thickness must be maintained. As
shown above in Equation (4), the position of the rein-
forcement in concrete also has an effect on its final
strength. The horizontal rods closer to the bottom have a
better bond-strength than those closer to the top surface.
The reason for this is a gradual settlement of the
concrete. This is more significant for the concretes with
a higher mixing-water content.
The surface of a reinforcement plays a major role. It
needs to be rough and clean, suitably degreased and
rust-free. When a load is applied to the reinforcement,
the ribbings stress the concrete in their vicinity, creating
a transverse tensile stress, eventually causing the con-
crete to crack behind the ribbings, relative to the trajec-
tory of the applied load (Figures 1 and 2). A higher
loading causes the shear strength of concrete to be
exceeded and the reinforcement rod to "cut out" of the
concrete (Figure 3).12,14
The bond-strength of concrete is mainly determined
by the mechanical bonding factor; for the ribbed bar
reinforcement, the bond-strength depends on the relative
rib area fR that can be calculated from the rib geometry
using Equation (5), while the area of a single rib FR
needs to be calculated using Equation (6) (d is the nomi-
nal reinforcement-bar diameter, c is the distance between
the centers of two adjacent ribs,  is the angle of the rib
inclination, n is the number of crosswise rows on the rod
perimeter, m is the number of differently angled ribs in a
column, q is the number of crosswise longitudinal ribs
for cold-bent rods, P is the number of threads of the rib
spiral, ak' is the average height of longitudinal ribs, aS,i is
the average height of the i parts of the ribs divided in p
parts with a l length, Figures 4 and 5). The second
addend applies only to cold-bent rods and is neglected if
it exceeds 30 % of the total value of fR:15,16
f
d
m F
P
a
i jj
m
i j
i
n
k
k
q
R
R
ic
= +
=
= =
∑
∑ ∑1
1 11
1 1π
( / ) sin
'
, , ,
(5)
F a li
i
p
R S=
=
∑ ( ), Δ
1
(6)
The desired values of the relative rib surface and their
shapes are given in the design standards. The minimum
P. POKORNÝ et al.: PROBLEMS AND NORMATIVE EVALUATION OF BOND-STRENGTH TESTS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 847–856 849
Figure 3: Idealized model depicting cutting-out of the ribbed
reinforcement right before its pull-off of concrete. Longitudinal crack
spreads through concrete closely above the ribs.
Slika 3: Idealiziran model prikazuje izpuljenje rebraste armature tik
pred njenim izpuljenjem iz betona. Vzdol`ne razpoke se {irijo skozi
beton, tik nad rebri.
Figure 1: Idealized effect (compressive – dotted vectors and tensile –
full length vectors) of ribbed rod reinforcement in the beginning
stages of loading (black vector shows the strain direction)
Slika 1: Idealiziran vpliv (tla~ni – ~rtkasti vektorji in natezni –
polno~rtni vektorji) rebraste palice armature v za~etnem stadiju
obremenitve (~rni vektor ka`e smer deformacije)
Figure 4: Detail of a rib configuration for a calculation of the relative
rib surface fR
Slika 4: Podrobnosti konfiguracije reber za izra~un relativne povr{ine
reber fR
Figure 2: Idealized model of reinforcement movement (black vector),
when the bond is broken (dotted vectors show the compressive forces
on the concrete caused by the ribs)
Slika 2: Idealiziran model premikanja armature (~rni vektor), ko se
povezava prekine (~rtkasti vektorji ka`ejo tla~ne sile na beton, ki jih
povzro~ijo rebra)
relative surface fR,min, related to the nominative bar dia-
meter, is usually known. The nominal evaluation of the
minimum relative rod surface is shown in Table 1.15
Table 1: Evaluation of minimum relative surface of ribs to ensure a
good strength of the bond with bars, foils and also welded nets15
Tabela 1: Ocena minimalnih povr{in reber za zagotovitev dobre
trdnosti vezave s palicami, folijami in tudi z varjenimi mre`ami15
nominal diameter of
reinforcement ds/mm
minimum relative rib surface
(fR,min)
5–6 0.035
6.5–2 0.040
> 12 0.056
3 PROBLEMS OF CONCRETE-REINFORCEMENT
BOND-STRENGTH TESTS
The available standards recommend two experi-
mental set-ups to determine the reinforcement-concrete
bond-strength. The first of them is based on pulling out a
steel-bar reinforcement set from the concrete. The
second set-up is basically a 4-point beaming test – a
determination of the bond-strength for bent concrete
girders, a girder test, a beam test, etc.
Both methods produce quite different results which
make an objective assessment of the concrete-reinforce-
ment bond-strength difficult. Nonetheless, comparing the
data obtained with a repeated measurement using the
same method is also somewhat complicated. For both
methods, the non-objectivity of the measurements come
from the standards themselves. In the case of the pull-out
test of a reinforcement with a bonding component
(indentation, rib, corrugation), the concrete is disinte-
grated during the test (Figure 3). The results are, there-
fore, strongly dependent on the strength of the concrete.
Also, the results of "the beam test" cannot be used to
make a general statement about the bond-strength – the
forced bending momentum creates a compressive stress
in the upper part of the beam and a tensile stress in the
lower part. Any reasonably changed experimental set-up
still cannot provide us with the data disregarding the
mechanical properties of concrete.
These influences can be ignored using a modified
pull-out test with a smooth-surface reinforcement. How-
ever, the reinforcement must be perfectly aligned with
the central axis of the fixed concrete sample so that there
is no compressive stress. The data from these bond-
strength tests for concrete and reinforcement cannot be
used to make general assumptions about the real bond-
strength since they are also influenced by many other
factors. On the other hand, the results of a smooth rein-
forcement-concrete bond-strength test are of no practical
use since the bond is primarily facilitated by the bonding
components which are considered in the static calcula-
tions.15,17
3.1 Bond-strength of a coated reinforcement
Coating provides the reinforcement with a surface
barrier which increases the time until the surface activa-
tion of the steel. A supplementary reinforcement coating
can thus be used to prolong the longevity of iron-con-
crete constructions. However, the strength of the bond
between the coating and the concrete must be correctly
evaluated. A reduced coated reinforcement-concrete
bond-strength, even on a scale of percent units, can make
its practical application much more difficult. The reason
for this is the concern about the static reliability, espe-
cially in the applications with high requirements – con-
structions with very high load-bearing capacities, dyna-
mically stressed constructions. A reduced bond-strength
can be solved by increasing the bond length or adding
bonding profiles. A less effective alternative to this is a
surface modification of the reinforcement (increasing the
surface area of the bonding elements, i.e., ribs, inden-
tations, corrugations, etc.). Increasing the bond length
increases the cost of the construction; the coating
requires additional concrete-reinforcement bond-strength
tests to be performed (using the reinforcement with a
modified surface). Two surface modifications are dis-
cussed: the hot-dip galvanized coating and the epoxy
coating.15
A comparison of the bond-strength results between
the coated and non-coated reinforcements of the same
surface geometry must be done according to the standar-
dized tests. A correct interpretation of the data is also of
great importance. The bond-strength must also be
measured when the reinforcement surface geometry, the
concrete’s chemical composition and other factors
altering the bonding interaction are changed.
3.1.1 Hot-dip galvanized reinforcement
The suitability of a hot-dip galvanized (zinc) coating
for a concrete steel reinforcement is still arguable. This
modification provably has a positive effect on the resis-
tance to chlorides and also on the resistance to carbo-
nated concrete.2,18 However, zinc actively corrodes in
fresh concrete (alkaline, pH often exceeds 13.0) pro-
ducing hydrogen. Hydrogen increases the porosity of the
concrete and reduces the adhesion factor, thus also
reducing the total bond-strength. After the zinc coating
actively corrodes in fresh concrete, the remaining coat-
ing does not always exhibit sufficient quality.17,19 Other
authors verified the initial reduction of the
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850 Materiali in tehnologije / Materials and technology 49 (2015) 6, 847–856
Figure 5: Cross-section of a rib for a calculation of the rib surface FR
Slika 5: Prerez rebra za izra~un povr{ine FR rebra
bond-strength; however, this is later compensated by zinc
corrosion products (Zn(OH)2) resulting from the
concrete carbonation and filling in the pores.20–22
Another group of authors claim that a zinc-coated
reinforcement easily becomes passivated in an environ-
ment with a pH value of 13.3 by forming non-soluble
Ca[Zn(OH)3]2·2H2O; sulphate anions also have a positive
effect on passivation.23–29 Another phenomenon – apart
from the negative effect of zinc on the corrosion resis-
tance of a reinforcement – a detrimental effect of zinc
during the concrete hardening also needs to be men-
tioned. It was proven that zincates slow down the harden-
ing of concrete and, with regard to the water content in
concrete, they can extremely deteriorate the mechanical
properties of concrete.30–32 Poor results for a zinc-coated
ribbed reinforcement are sometimes explained with the
smoothening of the surface by the zinc coating itself. A
lower bond-strength corresponds to a lower relative rib
surface fR, thus reducing the factor of mechanical
bonding in the bond-strength equation Bs. According to
these authors, hot-dip zinc coating can result in the
formation of an uneven coating – thicker at one heel of a
rib (depending on how the rod was removed from the
zinc bath) and very thin at the top of the rib (Figures 6
and 7).13 Others observed a reduced bond-strength for the
hot-dip galvanized reinforcement even after a 28-day
curing of the concrete.33,34
3.1.2 Epoxy-coated reinforcement
The main problem of this type of coating is its
mechanical resistance – it is very fragile and its manipu-
lation is therefore problematic. Coating defects are
formed during the bending, being set in the concrete, and
often also during its fabrication. Another disadvantage is
the necessity of welding prior to coating. A coated rein-
forcement can be linked only by sockets.8 The cracking
of a coated reinforcement stored at a temperature below
10 °C was also observed.35 Corrosion products of steel
forming on the surface of a reinforcement can also
damage the epoxy coating.36 A perfect compact coating
can prolong the longevity of concrete-steel constructions,
but the problem lies in the strength of its bond with the
concrete. Experts agree that an epoxy-coated reinforce-
ment has a decreased strength of the bond with the
concrete (sometimes even by 20–25 %). The recognized
reasons for this include the concrete not adhering well to
the reinforcement and creating only a small number of
physico-chemical interactions, the smoothening of the
ribs done in the way similar to the one used for the zinc-
coated reinforcement, or an elastic deformation of the
epoxy coating during the loading. Epoxy-coated rein-
forcements usually have higher bond lengths and other
anchoring modifications.37–40
4 DETAILED DESCRIPTION AND
CHARACTERIZATION OF THE STANDARDS
The following text sums up the standards and
recommendations for the arrangement, conditions and
evaluation of bond-strength tests (^SN standards,
RILEM recommendations).
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Materiali in tehnologije / Materials and technology 49 (2015) 6, 847–856 851
Figure 6: Model comparing the geometries of ribbed reinforcement
with and without zinc coating
Slika 6: Model primerjave geometrije rebraste armature z nanosom
cinka in brez njega
Figure 8: Schematics of the pull-out-test experimental set-up (con-
crete prism)
Slika 8: Shematski prikaz preizkusa izpuljenja (betonska prizma)
Figure 7: Cross-section model comparing the heights of the rib of
non-coated and coated ribbed reinforcement from Figure 6
Slika 7: Model prereza, ki primerja vi{ino rebra pri rebrasti armaturi z
nanosom in brez njega s slike 6
4.1 ^SN 73 1328 (determination of the concrete-rein-
forcement bond-strength)41
This standard is the basic regulation for evaluating
the bond-strength of the concrete with the components
described above. It deals with both dense and aggregate
concrete – the aggregate can be both dense and porous.
The standard does not evaluate the effect of the rein-
forcement surface on the bond-strength. In this standard,
the bond-strength is defined as the shear strength of the
concrete (in the shear strength m [MPa], m = Pm/(a·o),
a is the bond length, o is the diameter of the reinforce-
ment) when the pull-out of the reinforcement from the
concrete is 0.001 mm to 0.002 mm. The bond-strength is
thus defined as the shear strength on the tensile-stressed
bar perimeter – this value is defined during the project-
ing of iron-concrete constructions.
4.1.1 Determination of the steel-concrete bond-strength
for a bar-shaped reinforcement
Prior to the test, three cubic samples with 20 cm or
15 cm edge diameters, from the same concrete (i.e., an
identical production and treatment procedure), are
manufactured for the concrete-cube strength test. For the
bond-strength test, bars of precise dimensions are manu-
factured using a reinforcement bar. To ensure that this
bar is right in the axis of the concrete sample, it is
inserted in a tube placed in the bottom part. The fresh
concrete must be prevented from entering the tube. The
bond length is the length of the concrete sample reduced
by the length of the anchoring tube. The result is the
arithmetic average value from three measurements (three
parallel samples of one reinforcement type) of the bond-
strength measured after 28 days of curing. The results
that differ by more than 20 % from the average are dis-
carded (Figure 8).
4.1.2 Strength of the bond between the steel and
concrete in beam-stressed girders
Similarly to the previous case, it is recommended to
verify the concrete-cube strength prior to the test. Rec-
tangular cross-section girders are used for this sample.
Their dimensions are chosen to correspond to the bond
lengths of the reinforcements (Figure 9). The test rein-
forcement is set in the pulled part of the girder (i.e., the
part where the bending stress is manifested as the tensile
stress); the sample also contains two auxiliary reinforce-
ment bars in the compressed part. The girder is also
equipped with two closed clamps in the middle of the
tensile-stressed part with an artificial interstice reaching
about half of the girder’s height. There is also a very
narrow interstice in the tensile-stressed part. Both inter-
stices are present to provide a suitable load-distribution
path. The bottom reinforcement is also partially exposed
so that its deformation can be easily measured.
The reinforcements in this set-up are set in tubes to
ensure a precise settlement in the concrete and prevent
their deformation during the loading. There are steel spi-
kes in the heads of the girder with deviation meters for
measuring the movement of the reinforcement towards
the inner part of the girder (Figure 10). The test takes
place after 28 d of curing the concrete in specified con-
ditions. During the test, several parameters are measured:
the bending of the girder, which is measured in the
middle part with an accuracy of 0.001 mm, the deforma-
tion of the middle part of the bottom reinforcement (an
accuracy of 0.0001 mm), the shift of the reinforcement
towards the inner part of the girder (an accuracy of 0.001
mm) and the pulling strength during the first decrease of
the bond-strength (a reinforcement shift of 0.002 mm).
The test is repeated three times with the same reinforce-
ment type and an arithmetic value is calculated. The
strength values should not vary by more than 20 % from
the average value.
4.2 ^SN 73 1333 (testing the bond-strength of pre-
stressing reinforcement in concrete)42
This standard can be used for testing the bond-
strength of prestressing reinforcement with common
compacted or compacted light concrete from an artificial
porous aggregate. This test is used to assess the bond-
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852 Materiali in tehnologije / Materials and technology 49 (2015) 6, 847–856
Figure 9: Schematics of reinforced girder prepared for a beam test
with arrows showing the applied force
Slika 9: Shematski prikaz oja~anega nosilca, pripravljenega za upo-
gibni preizkus, pu{~ici ka`eta uporabljeno silo
Figure 10: Cross-section of the beam-test girder sample: compressed
part (full line) and tensile-stressed part (dotted line)
Slika 10: Prerez nosilca za upogibni preizkus: tla~ni del (polna ~rta)
in natezno obremenjen del (~rtkasta ~rta)
strength of the prestressing reinforcement in prestressed
concrete. On the other hand, this standard cannot be used
for assessing the effect of a construction loading on the
strength of a bond of a prestressed reinforcement with
the concrete (e.g., dynamically loaded constructions). It
is important that this standard considers the mechanical
bond between a prestressed reinforcement and the
modified surface (imprints, ribs, corrugation, etc.) to be
the deciding factor for anchoring the reinforcement in
concrete.
Concrete hinders the movement of the reinforcement
by pressing against its surface inhomegenities. Bonding
is defined as a reliable and safe transfer of the prestress-
ing force from a prestressing reinforcement to concrete.
The initial shift of the prestressing reinforcement is
defined by the standard as a "slip".
Similarly to the previous one, this standard also
recommends two experimental set-ups: the testing of
girders and of cubes. The testing of girders allows us to
assess the bond-strength between a prestressing rein-
forcement and new-type surface modifications. The test-
ing of cubes is viable for a study of the effects of various
factors (surface modification of the reinforcement, com-
position, processing and treatment of concrete) on the
bond-strength.
Three samples are tested after 28 d of the curing. The
composition, the processing and curing process are
precisely defined.
4.2.1 Testing of the bond-strength of girders
This standard evaluates the change in a prestress-
ing-force value before and after a slip of the reinforce-
ment inside a concrete sample. The slip length defining
the bond-strength is defined as 0.001 mm. The girder
geometry is different from the girder in the ^SN 73 1328
standard. The cross-section of the girder is also rectan-
gular; however, it only contains one or two tested pre-
stressing reinforcement(s). The tail pieces of the rein-
forcement are anchored struts so that it cannot slip more
than ×10–4 of its length between the anchoring points.
The girder is usually equipped with detectors measuring
the total longitudinal transformation (at least five on all
the sides along the girder). The positions of dial
deviation meters are similar to the ones defined in ^SN
73 1328.
The relative longitudinal girder transformation is the
average value of all the values from the total longitudinal
transformation of the opposite sides. The transformation
magnitude versus time is plotted alongside the girder for
every test stage. The changes in the prestressing force
alongside the girder and the bond length are estimated
from these plots. In this case, the bond length is con-
sidered to be the distance between the head of the rein-
forcement and the place where the transformation mag-
nitude stops increasing. The reinforcement shift is the
value from a dial deviation meter diminished by the
value of the elastic shortening between the meter and the
head of the girder.
The total value of the bond length is the average
value calculated from the values taken up to six hours
after introducing the prestressing load on both sides of
all three girders (six values). The total value of the
prestressing-reinforcement shift is calculated similarly,
but from the reinforcement components (6–12 values).
Individual values must not differ by more than 20 %
from the average value.
4.2.2 Testing of the bond-strength of cubes
During this test, a non-prestressed reinforcement is
pulled out of a concrete cube. The set-up is similar to the
one for the bond-strength test of the reinforcement and
bars described in ^SN 73 1328. The bond-strength is
calculated from the force necessary to pull the reinforce-
ment out and from the shift of the other end of the rein-
forcement inside the concrete.
The cubes must be produced in such a way that the
reinforcement rod is the cube’s axis. The standard
recommends the use of wooden trapezoidal laths to
prevent the movement of the reinforcement. One lath is
placed diagonally on the bottom of the mold, the other
one is placed on top of it. The tailpieces of the reinforce-
ment rod are placed in the holes of the laths.
The load of the prestressing reinforcement is
measured by deviation with a 0.001 mm accuracy.
The force to pull the reinforcement out of the cube is
increased in 8–12 steps with a short pause (30 s) between
the steps. The force for each step is increased smoothly
and slowly so that each step takes 20 s.
The deviation is first measured before the loading,
then a couple of steps before its maximum value and also
after this value is exceeded. During the test, several
values are monitored: the force during the first shift of
the reinforcement inside the concrete sample For, the
maximum force Fmax and Flim which is the force acting
when the reinforcement is being continuously pulled
inside the sample without the need to increase this force.
The stress of the bonding layer depends on For, Fmax and
Flim and it is calculated with Equation (1), where l is the
Materiali in tehnologije / Materials and technology 49 (2015) 6, 847–856 853
P. POKORNÝ et al.: PROBLEMS AND NORMATIVE EVALUATION OF BOND-STRENGTH TESTS ...
Figure 11: Schematics of a reinforced beam prepared for a beam test
with arrows showing the applied forces and a connecting hinge
according to RILEM RC5
Slika 11: Shematski prikaz oja~anega nosilca, pripravljenega za
upogibni preizkus. Pu{~ici ka`eta uporabljeni sili in povezovalni te~aj,
skladno z RILEM RC5
length of the reinforcement set in the concrete and a is
the nominal perimeter length of the reinforcement:
 x
x
u
F
l
= (7)
The calculated stresses or, max, lim for the bonding
layer are the average values of all the values of at least
three measurements (three cube samples). Individual
values must not vary by more than 30 % from the
average value.
4.3 RC5 (bond test for reinforcement steel 1: beam
test)
RC5 is one of the basic RILEM recommendations
dealing with another modification of the beam test. This
procedure can be used to verify the bond-strength bet-
ween common reinforcement and normal concrete, and
also between prestressed constructions and concrete. An
experimental set-up is depicted in Figure 11. Two sepa-
rate concrete blocks are, at the bottom, connected by a
reinforcement, whose bond-strength is to be measured.
The reinforcement bar is again set into tubes to ensure its
proper alignment and a precise bond length. The top
parts of the blocks are connected by a separating hinge
with a similar purpose as that of the interstices in the
other set-ups. The hinge dimensions, the geometry of the
supporting beams, the diameter of the reinforcement
bars, the length and height of the concrete blocks and
other parameters divide this test into two set-up groups –
types A and B. The inward shift of the reinforcement
inside the block is measured on both sides. The locations
of dial deviation meters are identical to the ones defined
in ^SN 73 1328.
Similarly to the other standards, this recommendation
also requires the use of non-corroded, properly de-
greased test samples. The surface modifications must not
affect the geometry of the tested reinforcement bar. The
standard defines the concrete composition (aggregate and
gravel) in terms of a viable granulometry; the water con-
tent is strictly defined. Again, the compressive strength
of the cubes is verified.
The bond-strength is measured after a 28 d curing in
precisely defined conditions. After setting the test beam
on a mobile ball or triangle support, a force is applied to
the top part of the girder, symmetrically to both blocks.
The force is applied continuously so that the next-step
magnitude of the force is reached in up to 30 s. After
reaching the desired value, the applied force is held at
that value for 120 s. The shift of the bar towards the
inside of the block is measured (an accuracy of 0.001
mm); in addition, a measurement is taken for each step.
The test continues until exceeding the bond-strength
between the concrete and the reinforcement.43
4.4 RC6 (bond test for reinforcement steel 2: pull-out
test)
The RILEM recommendation also provides an alter-
native "pull-out test" described in the ^SN standards. A
non-prestressed steel reinforcement with a minimum of
10 mm in diameter is recommended for a bond-strength
measurement using this method. This test is recom-
mended for the bond-strength testing of the reinforce-
ments with different types of surface geometry. Con-
sidering the statistical evaluation, a minimum of 5
parallel samples is recommended. It is also recom-
mended to set the reinforcement in a cubic concrete
block. Again, the shift of the reinforcement towards the
inside of the block is measured by a suitable deviation
meter placed on the top part of the non-stressed
reinforcement. The bond length of the reinforcement and
the dimensions of the cube are chosen in such a way that
they are proportional to the ratio between the bond
length and the reinforcement-bar diameter. The rein-
forcement bar is, again, precisely inserted in a plastic
tube which is suitably protected against the fresh-con-
crete contamination. The standard defines the prepara-
tion process and the composition of the concrete, its
processing during the curing and the verification of its
cubic compressive strength. The testing continues until a
complete loss of the cohesion between the concrete and
the reinforcement takes place. The result of this test is a
loading curve F = f(l). The loading force must be,
similarly to the beam test (RC5), proportional to the dia-
meter of the reinforcement bar and it must be continually
increased (the rate of the increase must be steady).
The most important result of this experiment is the
stress determined for the bonding layer, calculated from
the loading force after the 28 d curing of the concrete.44
5 CONCLUSION
A good bond-strength between any reinforcement
type and concrete is one of the main prerequisites for a
reliable static function of steel-concrete constructions.
The bond-strength is affected by many factors: the
adhesive forces between concrete and reinforcement, the
friction forces caused by the surface inhomogeneities of
the flat parts of the reinforcement, the surface geometry
(ribs, imprints and corrugations), the composition and
mechanical properties of the used concrete, its process-
ing, curing time and also the position of the reinforce-
ment in the concrete.
To make a valid generalization about the effect of the
corrosion of hot-dip galvanized steel or the constriction
of epoxy coating, it is necessary to use the methods that
minimize the effect of the concrete’s mechanical pro-
perties on the measured bond-strength. The only suitable
method is the pull-out test (the bond-strength test of bars
and cubes). A very important requirement of this test is
that the reinforcement must be placed along the longi-
tudinal axis of the concrete sample so that the pulled-out
854 Materiali in tehnologije / Materials and technology 49 (2015) 6, 847–856
P. POKORNÝ et al.: PROBLEMS AND NORMATIVE EVALUATION OF BOND-STRENGTH TESTS ...
reinforcement bar does not generate pressure forces. This
can be achieved by using the tubes set in the concrete –
the reinforcement rod is inserted in these tubes and a
proper alignment is ensured, unlike when using anchor-
ing agents which are not part of the mold.
Acknowledgement
This work was accomplished with the financial
support from Specific University Research (MSMT No.
20/2014), the Grant Agency of the Czech Republic, reg.
number 14-20856S and the Technological Agency of the
Czech Republic, TA02030164.
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1990
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M. KOVA^I^, R. JAGER: MODELING OF OCCURRENCE OF SURFACE DEFECTS ...
857–863
MODELING OF OCCURRENCE OF SURFACE DEFECTS OF C45
STEEL WITH GENETIC PROGRAMMING
MODELIRANJE POJAVA POVR[INSKIH NAPAK PRI JEKLU C45 Z
GENETSKIM PROGRAMIRANJEM
Miha Kova~i~1, Robert Jager2
1[tore Steel, d. o. o., @elezarska cesta 3, 3220 [tore, Slovenia
2Laboratory for Multiphase Processes, University of Nova Gorica, Vipavska 13, 5000 Nova Gorica, Slovenia
miha.kovacic@store-steel.si
Prejem rokopisa – received: 2013-12-09; sprejem za objavo – accepted for publication: 2015-02-18
doi:10.17222/mit.2013.304
Carbon steel C45 with an increased content of carbon is used for tempering in the automotive industry for highly stressed parts
(axles, shafts), machine parts, screws, drills for wood, axes, knives, hammers and similar. In the present work an attempt of
analyzing the influences of different steelmaking parameters is presented. On the basis of the monitored data about the
casting-temperature changes, the total oxygen, the number of added aluminum rods, the chemical analyses before and after
steelmaking, the added lime, the aluminum-cored wire, the calcium-silicon-cored wire, the sulphur-cored wire, the rolling
dimensions, the casting speed, the opening of the ladle nozzle with oxygen and surface defects (scrap fraction) on rolled bars, a
mathematical model was obtained with the help of the genetic programming method. The results show that the most influential
parameters for the surface-defect occurrence on the C45 steel are the opening of the ladle nozzle with oxygen and aluminum.
On the basis of the results, the steelmaking technology was changed. Instead of the aluminium-killed steelmaking technology
the aluminium-calcium-free (ACF) steelmaking technology was used. The batches from an aluminium-calcium-free steelmaking
period statistically have a significantly lower level of surface defects (scrap fraction). The scrap fraction was reduced from the
average of 68.45 % to 1.92 % – by more than 35 times.
Keywords: mechanical engineering, metallurgy, steel, C45 steel, making steel, casting steel, steel plant, billet surface defects,
genetic programming
Jeklo C45 je ogljikovo jeklo za pobolj{anje s pove~ano koncentracijo ogljika. Uporablja se za obremenjene dele v avtomobilski
industriji (osi, gredi), za dele strojev, vijake, svedre za les, sekire, no`e, kladiva in podobno. V ~lanku je predstavljen poskus
analize vplivov razli~nih parametrov pri postopku izdelave omenjenega jekla. Na podlagi zbranih podatkov o spremembi livne
temperature, aktivnega kisika (kisik v talini), dodanih aluminijevih palic, analiz kemijskih elementov med izdelavo jekla in po
njej, dodanega apna, dodanega aluminija, kalcij-silicija ter `vepla v obliki `ice, dimenzij valjanca, hitrosti litja, pod`iganja in
podatkov o povr{inskih napakah (dele` izmeta) valjanih palic smo izdelali matemati~ni model z genetskim programiranjem.
Rezultati modeliranja genetskega programiranja so pokazali, da sta pod`iganje in aluminij tista parametra, ki najbolj vplivata na
nastajanje napak v jeklu C45 in imata torej najve~ji vpliv na izmet. Na podlagi rezultatov se je spremenila tehnologija izdelave
jekla. Namesto tehnologije z dodatkom aluminija se je uporabila tehnologija aluminium-calcium-free (ACF). [ar`e, izdelane v
obdobju izdelave jekla s to tehnologijo, imajo statisti~no zna~ilno manj povr{inskih napak (izmeta). Izmet se je v povpre~ju
zmanj{al iz 68,45 % na 1,92 % – ve~ kot 35-krat.
Klju~ne besede: strojni{tvo, metalurgija, jeklo, jeklo C45, izdelava jekla, ulivanje jekla, jeklarna, povr{inske napake na gredici,
genetsko programiranje
1 INTRODUCTION
The basic concerns about the development of the
continuous-casting technology are associated with find-
ing the source generating surface defects and taking pro-
per measures to prevent and remedy them where appro-
priate.1–3
A literature review shows that there are three basic
ways of modeling surface defects:
– experimental approach4,
– computational fluid dynamics1–3 and
– artificial-intelligence approach which effectively
combines the above-mentioned approaches.5–10
The authors5,6 propose an interaction between the
numerical heat-transfer model and the artificial-intelli-
gence heuristic-search method, linked to a knowledge
base for the continuous casting. A two-dimensional
heat-transfer model was developed using the finite-
difference method and applied to real continuous-casting
conditions. The heuristic search method, aided by a
knowledge base, explores the technological and me-
tallurgical parameter settings in order to find optimized
cooling conditions, resulting in a defect-free billet pro-
duction with the minimum metallurgical period.
The paper by Tirian et al.7 describes a neural-
network-based strategy for crack prediction aimed at
improving the steel-casting process performance by de-
creasing the number of crack-generated failure cases. A
neural system for estimating the crack-detection proba-
bility was designed, implemented, tested and integrated
into an adaptive control system.
A simulated annealing-optimization algorithm was
used for a multicriteria optimization procedure to deter-
mine appropriate process-parameter values for producing
Materiali in tehnologije / Materials and technology 49 (2015) 6, 857–863 857
UDK 669.186:620.191:004.89 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)857(2015)
quality products in a continuous-casting system.8 A total
of 17 critical-quality conditions were identified; these
had to be satisfied to prevent defect formations during
the casting. An objective function, formulated as a loss
function, was used so that all 17 critical conditions were
satisfied simultaneously.
The article by Sanz-García et al.9 deals with three
successful experiences gained from genetic algorithms
and the finite-element method in order to solve engineer-
ing-optimization problems in connection with the
surface-defect occurrence during continuous casting.
A set of metamodels was developed to satisfy the
necessary process conditions connected to one or more
continuous-casting parameters.10 The values of these
parameters were determined so that all the process
conditions were satisfied simultaneously, ensuring that
the product had the desired quality.
This paper discusses the use of the genetic-program-
ming method for predicting the occurrence of surface
defects on rolled C45 steel, used for highly stressed parts
in the automotive industry, with respect to the influences
of several steelmaking parameters. Genetic programming
is one of the more general and recently developed
approaches of evolutionary algorithms. Similarly to
some other machine-learning approaches such as artifi-
cial neural networks, genetic algorithms, particle-swarm
optimization and gravitational search algorithm (see
examples11–14), genetic programming can be used for
solving a wide spectrum of engineering and other prob-
lems.
The problem is described in Section 2. In the subse-
quent section the effect of the proposed concept and the
results of the defect-occurrence prediction are presented.
A practical implementation of the modeling is presented
in Section 4 and, finally, the main contributions of the
performed research and guidelines for further research
are given in the last section.
2 EXPERIMENTAL BACKGROUND
Steelmaking begins with scrap melting in an elec-
tric-arc furnace. After the melting of the scrap and car-
burizing agents, the carbon carriers, in general, are coke,
anthracite, graphite and slag additives, which regulate
the basicity, viscosity, thermal and electric conductivi-
ties, desulphurization, dephosphorization, neutrality
towards the furnace fireproof linings and capability to
filter non-metallic inclusions.1,2
The melting bath heated up to the tapping tempera-
ture according to further treatment procedures is dis-
charged into the casting ladle after the electric-arc-fur-
nace melting.
The standard steelmaking procedure (in a ladle fur-
nace) for C45 consists of:
1. Melt-temperature measurement. Depending on the
temperature the melt is additionally heated.
2. Visual checking of the slag’s viscosity and amount.
The slag should be properly dense. In the case of not
having enough slag, lime and bauxite should be
added. The argon flux should also be appropriate.
3. Chemical analysis.
4. On the basis of the chemical composition, corrective
alloying is performed and argon stirring is also
consequently increased.
5. If necessary, aluminum and sulfur wires are added.
6. Chemical analysis.
7. On the basis of the chemical composition, fine
corrective alloying is performed (CaSi and aluminum
wires). During the alloying moderate argon stirring is
performed. The melt should also be properly covered
with slag due to a possible melt reoxidation. In the
end a sulfur wire is added.
8. Melt-temperature measurement.
After achieving the proper melt temperature in the
melting bath the billets are continuously cast. The
melting bath flows through the sliding-gate system and
ladle shroud towards the tundish.
During the casting it may be necessary to use the
oxygen lance to cut through the ladle slide gate and pour
the liquid steel through it. In good steelmaking practice,
we want to have a minimum number of such batches
because the oxygen that is blown into the melt causes a
re-oxidation of the melt and impurities may be formed.15
After filling up the tundish, the mould filling system
with tundish stoppers and submerged pouring tubes is
established. Billets with a square section of 180 mm are
cast. After reaching a certain melting-bath level the
potentiometer starts the flattening system which drags a
billet out of the mould. In this way continuous casting is
established. The billet goes through the cooling zone
toward the gas cutters where it is cut and laid off onto the
cooling bed.
The cooled-down billets are reheated, according to
the prescribed temperature, in the continuous-heating
furnace. After the heating the billets are hot rolled in a
strand of rolls. Steel bars are cooled on the cooling bed.
After the cooling they are cut with hot shears to different
lengths.
In line with the customer specifications, after the
rolling the cut bars are inspected on the automatic con-
trol line (Figure 1) equipped with three inspection units:
– anti-mixing control (FÖRSTER MAGNATEST, eddy
current),
– surface-defect control (FÖRSTER CYRCOFLUX,
magnetic-flux leakage) and
– internal-defect control (KRAUTKRAMER ROWA
90/160, ultrasound).
The minimum detectable surface-defect depth and
length are 0.15 mm and 12.5 mm, respectively. Regard-
ing the inspection speed and type of the ultrasonic head,
the minimum detectable inner-defect sizes are 0.8 mm
and 1.2 mm. Due to blind areas and the type of the
M. KOVA^I^, R. JAGER: MODELING OF OCCURRENCE OF SURFACE DEFECTS ...
858 Materiali in tehnologije / Materials and technology 49 (2015) 6, 857–863
internal defects, the control unit detects 85 % of the
material volume.
The data for the analysis was collected on the basis of
34 batches of the C45 steel consecutively inspected
(automatic control line) in [tore Steel Ltd. (Table 1)
from January to December 2010. The data was taken
from the technological documentation of the cast batches
and from the chemical archive. The goal was to get as
wide a range of influential parameters as possible. These
are:
– Active oxygen O2 (× 10–6): It influences the alumi-
num addition and, consequently, the aluminum-oxi-
de content. It is measured with a temperature probe
before the tapping.
– Lime CaO (kg): It is added during the tapping and it
helps create additional slag which is needed also
during the ladle-furnace treatment.
– Calcium carbonate (kg): It is added during the
tapping to reduce the slag density.
– Aluminum bricks (kg): They are added during the
tapping in order to deoxidize the melt (the killed
steel). Their amount depends on the active-oxygen
content.
– The presence of the slag in the electric-arc furnace:
The slag in the electric-arc furnace is the slag that
was unintentionally poured into the ladle during the
tapping. The electric-arc-furnace operator adjusts the
tapping according to the melt quantity and tap-hole
dimensions. The slag in the electric-arc furnace
negatively influences the processing of the melt in
the ladle furnace. Chemical and physical reactions
evolve within a time shift or it can happen that they
do not evolve at all.
– Aluminium wire (m): It is added to the ladle furnace
according to the working instructions. It binds
oxygen into aluminium oxides.
– CaSi wire (m): Calcium is added for melt modifying.
It is proportional with the aluminium content. The
modification is, in fact, the ability of calcium to bind
unwanted inclusions and transport them into the
slag.
– tAl-CaSi (min): This is the time between the additions
of aluminium and CaSi wires.
– Sulphur wire (m): It is added to form MnS inclu-
sions which increase the machinability.
– The contents of Al, Ca, S at the beginning of the
melt’s ladle-furnace treatment (%)
– The final contents of Al, Ca and S (%)
– T (°C): This is the temperature difference between
the actual and prescribed casting temperatures
measured in the tundish.
– ns: This is the number of the strands used during the
casting. A continuous caster has three strands, but it
can happen that the inner nozzle becomes clogged
and the casting is, therefore, performed with fewer
strands. Proportionally, the casting lasts longer and,
consequently, the melt temperature drops influencing
the casting speed.
– The casting speed (m/min): It depends on the steel
chemical composition and the melt temperature.
– The casting-speed variation (m/min): Due to tempe-
rature drops the casting speed must be changed.
– nlno: This is the number of the ladle-nozzle openings
with the oxygen lance.
– The rolled-bar diameter (mm)
M. KOVA^I^, R. JAGER: MODELING OF OCCURRENCE OF SURFACE DEFECTS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 857–863 859
Table 1: Experimental data
Tabela 1: Eksperimentalni podatki
B
at
ch
Electric-arc-furnace
parameters Ladle-furnace parameters Casting parameters
S
cr
ap
(%
)
O
2
×
10
–6
C
aO
(k
g)
C
aC
O
3
(k
g)
A
l
br
ic
ks
(k
g)
sl
ag
A
l
w
ir
e
(m
)
C
aS
i
w
ir
e
(m
)
t (A
l-
C
aS
i)
(m
in
)
S
w
ir
e
(m
)
A
l l
(%
)
S
l
(%
)
C
a l
(%
)
A
l
(%
)
S
(%
)
C
a
(%
)

T
(°
C
)
st
ra
nd
s

v/
(m
/m
in
)
v/
(m
/m
in
)
la
dl
e-
no
zz
le
op
en
in
gs
di
am
et
er
(m
m
)
1 367.8 670 40 65 0 130 80 10 110 0.013 0.02 0.005 0.028 0.027 0.025 40 3 0 1.73 0 37 27.42
1 367.8 670 40 65 0 130 80 10 110 0.013 0.02 0.005 0.028 0.027 0.025 40 3 0 1.73 0 44 33.29
2 357.3 600 60 82 0 90 90 10 105 0.01 0.045 0.003 0.029 0.029 0.035 55 3 0 1.65 0 28 18.91
2 357.3 600 60 82 0 90 90 10 105 0.01 0.045 0.003 0.029 0.029 0.035 55 3 0 1.65 0 33 30.92
3 444.7 600 30 60 0 140 70 29 120 0.007 0.015 0.002 0.03 0.031 0.027 42 3 0 1.65 0 37 33.67
Figure 1: Automatic control line in [tore Steel Ltd.
Slika 1: Samodejna kontrolna linija v [tore Steel, d. o. o.
M. KOVA^I^, R. JAGER: MODELING OF OCCURRENCE OF SURFACE DEFECTS ...
860 Materiali in tehnologije / Materials and technology 49 (2015) 6, 857–863
4 338.5 600 30 63 0 65 80 14 140 0.025 0.034 0.002 0.025 0.034 0.026 39 3 0 1.73 2 28 24.58
4 338.5 600 30 63 0 65 80 14 140 0.025 0.034 0.002 0.025 0.034 0.026 39 3 0 1.73 2 30 49.22
4 338.5 600 30 63 0 65 80 14 140 0.025 0.034 0.002 0.025 0.034 0.026 39 3 0 1.73 2 44 45.74
5 620.7 600 30 60 0 165 87 52 130 0.021 0.026 0.002 0.027 0.03 0.022 44 3 0 1.73 1 28 12.88
5 620.7 600 30 60 0 165 87 52 130 0.021 0.026 0.002 0.027 0.03 0.022 44 3 0 1.73 1 37 17.78
6 210.1 600 30 60 1 125 90 21 135 0.019 0.021 0.004 0.023 0.024 0.03 45 3 0 1.73 1 30 15.69
7 506.5 720 40 70 0 70 96 17 120 0.006 0.02 0.002 0.035 0.02 0.037 39 3 0 1.73 0 33 32.88
8 203.3 600 50 50 0 70 80 30 60 0.021 0.045 0.001 0.034 0.029 0.028 46 3 0.05 1.68 0 44 26.58
9 352.9 650 40 70 0 70 90 45 120 0.022 0.043 0.001 0.03 0.026 0.033 45 3 0 1.73 0 33 35.11
9 352.9 650 40 70 0 70 90 45 120 0.022 0.043 0.001 0.03 0.026 0.033 45 3 0 1.73 0 37 32.16
10 216.7 600 30 55 0 20 85 10 80 0.035 0.038 0.001 0.029 0.026 0.03 50 3 0 1.68 1 37 18.13
11 767.2 600 40 70 0 80 90 45 90 0.022 0.026 0.001 0.034 0.028 0.029 26 2 0 1.83 1 28 29.2
11 767.2 600 40 70 0 80 90 45 90 0.022 0.026 0.001 0.034 0.028 0.029 26 2 0 1.83 1 33 31.88
11 767.2 600 40 70 0 80 90 45 90 0.022 0.026 0.001 0.034 0.028 0.029 26 2 0 1.83 1 37 23.81
12 222.2 600 40 55 0 90 90 80 135 0.019 0.037 0.001 0.028 0.026 0.027 50 3 0 1.68 0 30 58.5
12 222.2 600 40 55 0 90 90 80 135 0.019 0.037 0.001 0.028 0.026 0.027 50 3 0 1.68 0 33 68.33
12 222.2 600 40 55 0 90 90 80 135 0.019 0.037 0.001 0.028 0.026 0.027 50 3 0 1.68 0 44 26.85
13 313.2 640 40 62 0 0 90 / 70 0.037 0.033 0.004 0.032 0.033 0.027 45 3 0.05 1.73 2 37 46.87
14 544.8 600 40 75 0 105 100 40 140 0.015 0.034 0.001 0.029 0.025 0.033 40 3 0 1.73 0 44 37.67
14 544.8 600 40 75 0 105 100 40 140 0.015 0.034 0.001 0.029 0.025 0.033 40 3 0 1.73 0 37 61.15
15 513.6 600 60 45 0 50 75 10 130 0.021 0.03 0.003 0.03 0.024 0.031 53 3 0 1.63 0 44 33.39
15 513.6 600 60 45 0 50 75 10 130 0.021 0.03 0.003 0.03 0.024 0.031 53 3 0 1.63 0 28 53.25
16 560.1 760 70 60 0 75 85 11 173 0.029 0.014 0.001 0.033 0.029 0.034 48 3 0.05 1.68 0 30 37.74
16 560.1 760 70 60 0 75 85 11 173 0.029 0.014 0.001 0.033 0.029 0.034 48 3 0.05 1.68 0 37 50.29
16 560.1 760 70 60 0 75 85 11 173 0.029 0.014 0.001 0.033 0.029 0.034 48 3 0.05 1.68 0 30 5.216
17 163.6 600 70 55 0 70 85 14 120 0.023 0.037 0.001 0.033 0.03 0.027 47 3 0 1.73 0 44 23.76
17 163.6 600 70 55 0 70 85 14 120 0.023 0.037 0.001 0.033 0.03 0.027 47 3 0 1.73 0 37 63.09
18 320.6 600 30 62 0 225 85 60 200 0.007 0.019 0.002 0.028 0.023 0.035 40 3 0 1.73 4 28 4.732
19 162.6 600 30 66 0 0 80 / 70 0.04 0.041 0.002 0.032 0.038 0.036 42 2 0.12 1.85 2 33 4.762
20 317.8 600 30 62 0 90 80 24 130 0.031 0.022 0.004 0.039 0.032 0.038 50 3 0 1.68 1 28 11.27
20 317.8 600 30 62 0 90 80 24 130 0.031 0.022 0.004 0.039 0.032 0.038 50 3 0 1.68 1 30 56.41
21 313 600 40 62 0 100 90 55 135 0.017 0.026 0.002 0.028 0.023 0.031 46 3 0 1.73 3 37 11.08
21 313 600 40 62 0 100 90 55 135 0.017 0.026 0.002 0.028 0.023 0.031 46 3 0 1.73 3 30 14.81
21 313 600 40 62 0 100 90 55 135 0.017 0.026 0.002 0.028 0.023 0.031 46 3 0 1.73 3 37 27.12
22 371.3 600 20 65 0 115 95 18 120 0.017 0.018 0.001 0.031 0.035 0.034 40 3 0 1.73 1 30 38.47
22 371.3 600 20 65 0 115 95 18 120 0.017 0.018 0.001 0.031 0.035 0.034 40 3 0 1.73 1 37 37.12
23 327 600 20 65 0 25 95 15 105 0.032 0.02 0.003 0.028 0.027 0.041 40 3 0 1.73 0 53 68.78
23 327 600 20 65 0 25 95 15 105 0.032 0.02 0.003 0.028 0.027 0.041 40 3 0 1.73 0 37 46.86
24 170.6 600 25 45 0 145 80 38 0 0.017 0.02 0.001 0.029 0.033 0.026 42 3 0 1.73 1 53 18.11
25 630.6 600 30 70 0 125 80 20 133 0.012 0.036 0.005 0.032 0.028 0.026 48 3 0 1.73 1 44 24.85
25 630.6 600 30 70 0 125 80 20 133 0.012 0.036 0.005 0.032 0.028 0.026 48 3 0 1.73 1 37 25.43
26 248 600 40 65 0 110 85 40 75 0.015 0.031 0.002 0.029 0.034 0.025 43 3 0 1.73 0 44 25.38
27 355.7 600 30 70 0 80 85 13 70 0.02 0.026 0.001 0.033 0.03 0.019 51 3 0 1.68 0 33 50
27 355.7 600 30 70 0 80 85 13 70 0.02 0.026 0.001 0.033 0.03 0.019 51 3 0 1.68 0 37 64.21
28 155.8 600 70 60 0 120 85 33 120 0.01 0.03 0.001 0.029 0.035 0.024 42 3 0 1.73 1 33 18.56
28 155.8 600 70 60 0 120 85 33 120 0.01 0.03 0.001 0.029 0.035 0.024 42 3 0 1.73 1 37 35.6
29 205.6 600 40 55 0 110 85 6 107 0.011 0.04 0.002 0.033 0.023 0.03 45 3 0 1.73 0 44 22.81
30 267.1 600 40 60 0 50 75 10 95 0.025 0.033 0.001 0.03 0.027 0.031 46 3 0.05 1.68 0 37 23.84
30 267.1 600 40 60 0 50 75 10 95 0.025 0.033 0.001 0.03 0.027 0.031 46 3 0.05 1.68 0 37 7.297
31 348.4 600 40 60 0 120 75 40 105 0.015 0.023 0.002 0.029 0.024 0.028 41 3 0.05 1.68 1 44 9.655
32 652.6 650 40 80 1 90 80 10 145 0.019 0.024 0.002 0.032 0.023 0.031 51 3 0 1.68 1 30 8.411
32 652.6 650 40 80 1 90 80 10 145 0.019 0.024 0.002 0.032 0.023 0.031 51 3 0 1.68 1 44 35.27
33 184.8 600 30 55 0 150 80 40 115 0.01 0.022 0.002 0.031 0.029 0.032 45 3 0.05 1.68 1 30 16.29
33 184.8 600 30 55 0 150 80 40 115 0.01 0.022 0.002 0.031 0.029 0.032 45 3 0.05 1.68 1 33 23.68
34 455.4 600 30 58 0 120 90 63 110 0.024 0.02 0.002 0.026 0.021 0.027 44 3 0 1.73 1 37 35.61
34 455.4 600 30 58 0 120 90 63 110 0.024 0.02 0.002 0.026 0.021 0.027 44 3 0 1.73 1 44 29.76
34 455.4 600 30 58 0 120 90 63 110 0.024 0.02 0.002 0.026 0.021 0.027 44 3 0 1.73 1 37 35.03
34 455.4 600 30 58 0 120 90 63 110 0.024 0.02 0.002 0.026 0.021 0.027 44 3 0 1.73 1 44 38.83
3 MODELING OF OCCURRENCE OF SURFACE
DEFECTS WITH GENETIC PROGRAMMING
Genetic programming is probably the most general
evolutionary optimization method.15–19 The organisms
that undergo an adaptation are in fact mathematical
expressions (models) for a nozzle-opening prediction
consisting of the available function genes (i.e., the basic
arithmetical functions) and terminal genes (i.e., inde-
pendent input parameters and random floating-point
constants). In our case the models consist of the function
genes of addition (+), subtraction (–), multiplication (×)
and division (/) and terminal genes of active oxygen O2
(O2), lime (CaO), calcium carbonate (CaCO3), aluminum
bricks (Alb), the electric-arc-furnace slag presence
(slag), the aluminium wire (Alw), the CaSi wire (CaSiw),
the time between the additions of aluminium and CaSi
wires (tAl-CaSi), the sulphur wire (Sw), the contents of
Al (All), Ca (Cal) and S (Sl) at the beginning of the
melt’s ladle-furnace treatment, the final contents of Al
(Al), Ca (Ca) and S (S), the temperature difference
between the actual and prescribed casting temperatures
measured in the tundish (dT), the number of strands used
during the casting (ns), the casting speed (vc), the
casting-speed variation (vvc), the number of ladle-nozzle
openings with the oxygen lance (nlno) and the rolled-bar
diameter (d).
Random computer programs of various forms and
lengths are generated by means of selected genes at the
beginning of the simulated evolution. Afterwards, the
varying of the computer programs during several
iterations, known as generations, is performed by means
of genetic operations. For the progress of the population
only the reproduction and crossover are sufficient. After
the completion of the variation in the computer programs
a new generation is obtained that is evaluated and
compared with the experimental data.
The process of changing and evaluating the orga-
nisms is repeated until the termination criterion of the
process is fulfilled. This is the prescribed maximum
number of generations.
For the process of simulated evolutions the following
evolutionary parameters were selected: the population of
organisms: 500, the greatest number of generations: 100,
the reproduction probability of 0.4, the crossover
probability of 0.6, the greatest permissible depth when
creating a population: 6, the greatest permissible depth
after the operation of a crossover of two organisms: 10,
and the smallest permissible depth of organisms when
generating new organisms: 2. Genetic operations of the
reproduction and crossover were used. For the selection
of organisms the tournament method with a tournament
size of 7 was used. For the model fitness, the average
square of deviation from the monitored data was se-
lected. It is defined as:
Δ
Δ
= =
∑ i
i
n
n
2
1
(1)
where n is the size of the monitored data and i is the
square of deviation of a single sample data. The devi-
ation of a single sample data, produced by an individual
organism, is simply:
Δ i i iE G= − (2)
where Ei and Gi are the actual and the predicted scrap
fractions (which depend only on surface defects), res-
pectively.
We developed 100 independent civilizations of ma-
thematical models for the scrap-fraction prediction. Each
civilization has the most succesfull organism – the ma-
thematical model for the scrap-fraction prediction. The
best most succesfull organism from all of the civiliza-
tions is presented here:
(3)
with the average relative deviation of 12.03 %.
The randomly driven process builds the fittest and
most complex models from generation to generation and
uses the ingredients that are most suitable for the expe-
rimental environment adaptation. For curiosity’s sake, an
analysis of the genes (parameters) excluded from the
models is presented in Figure 2.
On the basis of the number of genes excluded from
100 obtained mathematical models we can assume the
influence of the parameters on the scrap fraction. It is
clear from the figure that out of 100 genetically obtained
mathematical models only 12 models do not include the
parameter of aluminium blocks and only 3 out of 100
models do not include the parameter of the number of
ladle-nozzle openings. So, we can deduct that aluminium
blocks and ladle-nozzle openings with an oxygen lance
M. KOVA^I^, R. JAGER: MODELING OF OCCURRENCE OF SURFACE DEFECTS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 857–863 861
are probably the most important parameters influencing
the occurrence of surface defects.
4 PRACTICAL IMPLEMENTATION OF THE
MODELING RESULTS
At the beginning of 2011, the aluminium-calcium-
free (ACF) steelmaking technology was used for the C45
steel instead of the aluminium-killed steelmaking
technology. The data for the analysis was collected on
the basis of 17 consecutively inspected batches of the
C45 steel (automatic control line) in [tore Steel Ltd.
(Table 2).
Table 2: Automatic-control-line results (scrap fractions) for 17 con-
secutively inspected batches (automatic control line) of C45 steel
made with aluminium-calcium-free steelmaking technology
Tabela 2: Rezultati samodejne kontrolne linije (dele` izmeta) za 17
zaporedno pregledanih {ar` jekla C45, izdelanih po tehnologiji alu-
minium-calcium-free
Batch Scrap %
1 0
2 1.036
2 2.102
3 2.042
3 2.809
4 0.642
5 0.877
6 2.889
7 0.902
8 1.731
9 1.512
10 3.218
11 1.570
12 2.345
13 0.520
14 0.700
15 4.757
16 1.086
17 5.695
The data was analyzed using Microcoft Excel. The 
parameter was set at  = 0.05.
The t-test for unequal variances was used to compare
the differences between the two populations. The results
of the comparison of the automatic control results ob-
tained during the aluminium-killed steelmaking period
(Table 1) and aluminium-calcium-free steelmaking pe-
riod (Table 2) are presented in Table 3.
Table 3: Comparison of automatic-control results obtained during alu-
minium-killed steelmaking period (Table 1) and aluminium-calcium-
free steelmaking period (Table 2)
Tabela 3: Primerjava rezultatov samodejne kontrolne linije, pridob-
ljenih v obdobju izdelave jekla na podlagi tehnologije z dodajanjem
aluminija (tabela 1) in med obdobjem uporabe tehnologije alumi-
nium-calcium-free (tabela 2)
Al-killed
steelmaking ACF steelmaking
Mean 68.45302471 1.917526316
Variance 260.6865252 2.146430596
Observations 61 19
Hypothesized mean
difference 0
df 63
t-stat 31.76825617
P(T <= t) one-tail 8.65978E-41
t critical one-tail 1.669402222
P(T <= t) two-tail 1.73196E-40
t critical two-tail 1.998340522
t-test: two-sample assuming unequal variances
The comparison shows that there is a statistically
significant difference between the batches made within
the aluminium-killed steelmaking period and alumi-
nium-calcium-free steelmaking period (p < 0.05). The
batches from the aluminium-calcium-free steelmaking
period statistically have a significantly smaller amount of
surface defects (scrap percenatage).
5 CONCLUSION
The purpose of this paper was to reduce surface
defects (scrap fraction) on the C45 steel. The influences
of the casting-temperature changes, the total oxygen, the
number of added aluminum rods, the chemical analyses
before and after steelmaking, added lime, aluminum-
cored wire, calcium-silicon-cored wire, sulphur-cored
wire, the rolling dimensions, the casting speed, the
opening of the ladle nozzle with oxygen on the surface
defects of rolled bars were analyzed. The data for the
analysis was collected on the basis of 34 consecutively
inspected (automatic control line) batches of the C45
steel in [tore Steel Ltd. The data was taken from the
technological documentation of the cast batches and
from the chemical archive. Afterwards a model for a
scrap-fraction prediction was developed with the
genetic-programming method allowing an evolution of
better and better model variants during the simulation.
There were 100 different models and only the best was
used for the scrap-fraction prediction. The relative ave-
rage deviation of the actual scrap fraction from the pre-
dicted scrap fraction was 12.03 %. Also, the frequencies
of the genes excluded from the best 100 mathematical
M. KOVA^I^, R. JAGER: MODELING OF OCCURRENCE OF SURFACE DEFECTS ...
862 Materiali in tehnologije / Materials and technology 49 (2015) 6, 857–863
Figure 2: Frequency of the genes excluded from the best 100 mathe-
matical models for scrap fraction
Slika 2: Frekvenca izlo~enih genov na podlagi najbolj{ih 100 mate-
mati~nih modelov za dele` izmeta
models for the scrap fraction were analyzed. The results
show that the most influential parameters for the surface
defects occurring on the C45 steel are the opening of the
ladle nozzle with the oxygen lance and aluminum
(blocks). According to the results, the steelmaking tech-
nology was changed. At the end of 2009 and at the
beginning of 2010 the aluminium-calcium-free (ACF)
steelmaking technology was used with the C45 steel
instead of the aluminium-killed steelmaking technology.
The t-test for unequal variances was used to compare the
differences between the automatic-control results
obtained during the aluminium-killed steelmaking period
and the aluminium-calcium-free steelmaking period. The
batches from the aluminium-calcium-free steelmaking
period statistically had a significantly smaller amount of
surface defects (scrap percenatage). The scrap fraction
was reduced from the average of 68.45 % to 1.92 % – by
more than 35 times. In future the implemented procedure
can be applied to several steel grades.
6 REFERENCES
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Materiali in tehnologije / Materials and technology 49 (2015) 6, 857–863 863

M. GAVAS et al.: EFFECTS OF VARIOUS HELICALLY ANGLED GRINDING WHEELS ...
865–870
EFFECTS OF VARIOUS HELICALLY ANGLED GRINDING
WHEELS ON THE SURFACE ROUGHNESS AND ROUNDNESS IN
GRINDING CYLINDRICAL SURFACES
VPLIV RAZLI^NIH KOTOV VIJA^NICE PRI BRUSILNIH
KOLUTIH NA HRAPAVOST POVR[INE IN OKROGLOST PRI
BRU[ENJU VALJASTIH POVR[IN
Muammer Gavas1, Muammer Kýna2, Uður Köklü3
1Department of Manufacturing Engineering, Dumlupinar University, Kütahya, Turkey
2Istanbul Arel University, Istanbul, Turkey
3Department of Mechanical Engineering, Karamanoglu Mehmetbey University, Karaman, Turkey
ugurkoklu@gmail.com
Prejem rokopisa – received: 2014-04-07; sprejem za objavo – accepted for publication: 2015-01-05
doi:10.17222/mit.2014.065
Grinding is generally used in the final step of machining metallic materials to achieve the necessary surface quality and
dimensions. Grinding wheels with flat surfaces are commonly used in the process of grinding. However, due to the fact that
there is a great deal of contact length (corresponding to the grinding-wheel width) between the grinding wheel and the
workpiece, effective cooling during the grinding process may not be possible and, consequently, the heat in the deformation
region is increased. Due to these reasons, some undesired results such as an unqualified surface and a roundness error take
place. Various profiles of the grinding wheel were, therefore, proposed to improve the surface quality and decrease the
roundness error by modifying the grinding wheel and developing various methods. In this study, AISI 1050, AISI 4140 and AISI
7131 steel materials were subjected to the cylindrical-grinding process using wheels helically grooved at 15 °, 30 ° and 45 ° and
the obtained results such as the average surface roughness and roundness errors were compared with the results of the
flat-surface grinding wheels. The experimental results show that the surface roughness and roundness error are reduced when
using a helically grooved grinding wheel and, thus, the quality of the machined parts is improved.
Keywords: cylindrical grinding, helically grooved grinding wheel, roundness error, surface roughness
Bru{enje se navadno uporablja kot kon~na stopnja obdelave kovinskih materialov za zagotovitev kvalitete povr{ine in mer. Pri
postopku bru{enja se navadno uporabljajo brusni koluti z ravno povr{ino. Vendar pa je zaradi velike kontakne dol`ine (odvisno
od {irine brusilnega koluta) med brusilnim kolutom in obdelovancem ote`eno u~inkovito ohlajanje, zato se podro~je deformacije
ogreva. Lahko se pojavijo ne`eleni rezultati, kot so neustrezna povr{ina in napaka okroglosti. Za izbolj{anje kvalitete povr{ine
in zmanj{anje napak okroglosti so predlagani modificirani brusilni koluti z razli~nimi profili in razli~ne metode. V tej {tudiji so
bila jekla AISI 1050, AISI 4140 in AISI 7131 okroglo bru{ena s koluti z vija~nim utorom 15 °, 30 ° in 45 °. Dobljeni rezultati,
kot sta povpre~na hrapavost povr{ine in napaka okroglosti, so bili primerjani s tistimi, dobljenimi z brusilnimi koluti z ravno
povr{ino. Rezultati ka`ejo, da se povr{inska hrapavost in napaka okroglosti zmanj{ujeta pri uporabi brusilnih kolutov z vija~nim
utorom, torej se kvaliteta strojnih delov izbolj{a.
Klju~ne besede: okroglo bru{enje, brusni kolut z vija~nim utorom, napaka okroglosti, hrapavost povr{ine
1 INTRODUCTION
A better surface quality and higher efficiency are the
prerequisites for today’s machining industry in order for
it to be more competitive since modern manufacturing
processes require shorter production time and higher-
precision components.1 Compared with the other mate-
rial-removal processes such as turning, milling and
boring, the grinding process is more complex and more
difficult to control.2 Grinding is a finishing process,
broadly used in the manufacturing of the components
requiring fine tolerances, a good surface finish and a
higher dimensional and geometrical accuracy.3,4 In spite
of all the good results of this finishing process, there are
some phenomena that affect the results. These are the
chattering and vibration of the machine and workpiece,
burning, unacceptable changes in the surface layer, and
microcracks and burns that cause surface defects, in-
creasing the surface roughness and other defects. These
defects are caused by clamping the workpiece, the course
and magnitude of the grinding-wheel wear, and the
stiffness of the whole machine-tool/workpiece fixture
system.5,6
The effects of a discontinuous workpiece material on
the grinding performance were investigated by resear-
chers.7,8 To control and improve the grinding perfor-
mance we would have to manufacture engineered grind-
ing wheels with a desirable topography to optimize the
metal-cutting process. One example of such a design is a
grooved wheel.9 Several studies confirmed that inter-
mittent grinding not only decreases the grinding force,
specific energy, surface burn, waviness and temperature
but also optimizes the material-removal rate.10–12 Fan and
Miller12 developed a force model for grinding with seg-
mental wheels. Both experimental and analytical results
Materiali in tehnologije / Materials and technology 49 (2015) 6, 865–870 865
UDK 621.92:621.7.01:669.14 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)865(2015)
show that the average grinding force decreases and the
peak force increases due to segmental wheels, as
compared to conventional wheels. Larger spaces between
the segments further reduce the average force and in-
crease the surface roughness and peak force. Kim et al.13,
Jin and Meng14 constructed discontinuous grinding
wheels (DGWs) with multi-porous grooves. Their study
illustrated that DGWs significantly improved the grind-
ing performance and surface roughness.
Shaji and Radhakrishnan15 used slotted wheels with
graphite integrated into the slots. Three such wheels
were developed with a varying number of the slots for
lubricant sandwiching. The results showed that the
surface roughness and residual stress were lower in the
case of the graphite-slotted wheels. The results of the ex-
periments with helically grooved wheels grinding four
different materials were reported by Gavas et al.5 Their
study illustrated that the ground-surface roughness de-
creased for some materials in comparison to the conven-
tional grinding. The roundness slightly increased for
brass and AISI 1010, but did not change for the AISI
1040 and AISI 2080 steels. Zhang16 conducted helical
scan grinding (HSG) on brittle materials such as cera-
mics and glass, and ductile materials such as steel. The
experimental results showed that HSG not only improved
the ground-surface finish, but also decreased the
adhesion of the workpiece material to the cutting grits in
the grinding of tough materials such as stainless steel
SUS304, and the fracture area in the grinding of brittle
materials such as ceramics.
Zhang and Uematsu17 analytically studied this topic
to find the difference in the surface-generation mecha-
nism between HSG and traverse grinding, and they
proposed models for predicting the surface roughness in
HSG. Both the experimental results and analysis show
that the ground-surface roughness decreases with the
helix angle and reaches the limiting surface-roughness
value at the critical helix angle, which is dependent on
the speed ratio. The HSG method is more effective in im-
proving the ground-surface roughness for large and/or
coarse wheels than for small and/or fine wheels. Recent-
ly, Nguyen and Zhang18 investigated the performance of
a segmented-grinding-wheel system and succeeded to
maintain the sharpness of the active cutting edges while
minimizing the ploughing and rubbing deformations of
ground workpieces.
Using the intermittent grinding wheels, the characte-
ristics of the grinding force, temperature, surface rough-
ness and geometric error were evaluated by Kwak and
Ha.19 With the results of the experiments with conven-
tional and intermittent wheels, it was proven that the
intermittent wheel appeared to be superior in various
aspects such as the grinding force, the temperature and
the geometric error, showing little deterioration of the
surface roughness. Tawakoli and Azarhoushang20 investi-
gated the feasibility of intermittent grinding with a seg-
mented wheel, using two ceramic-matrix composite ma-
terials. The grinding forces, surface roughness, surface
profile, elastic deformation and tool wear were compared
when grinding the ceramic-matrix composites with
segmented and normal diamond wheels. The finer sur-
face roughness obtained with the conventional grinding,
compared to the intermittent grinding with a T-tool
wheel, was due to a higher number of active cutting
edges and more rubbing in the process.
In this paper, the results of the grinding experiments
with helically grooved wheels used on three materials
are presented and compared with those obtained with the
conventional method. The grinding experiments were
conducted on the AISI 1050, AISI 4140 and AISI 7131
steel materials, with three helically grooved grinding
wheels. The surface roughness and roundness were
studied as the performance criteria and better results
were achieved when using the wheels helically grooved
at 15 ° and 30 ° than a flat-surface wheel.
2 EXPERIMENTAL PROCEDURE
2.1 Preparation of helically grooved grinding wheels
The grinding tests were performed on a horizontal
spindle-type cylindrical-grinding machine with four
aluminum-oxide grinding wheels. All the wheels used in
the experiments had the same specifications. However,
one of the grinding wheels had a flat surface. This type is
called the flat-surface grinding wheel (FSGW). The
other grinding wheels with different helix angles were
manufactured for this study. These types are called
helically grooved grinding wheels with angles of 15 °
(HGGW 15 °), 30 ° (HGGW 30 °), and 45 ° (HGGW 45 °).
Figure 1 shows schematic drawings of the dimensions of
M. GAVAS et al.: EFFECTS OF VARIOUS HELICALLY ANGLED GRINDING WHEELS ...
866 Materiali in tehnologije / Materials and technology 49 (2015) 6, 865–870
Figure 1: Dimensions of the grinding wheels: a) FSGW, b) HGGW
15 °, c) HGGW 30 ° and d) HGGW 45 ° (dimensions in milimeters),
e) isometric perspective of HGGW 15 °
Slika 1: Dimenzije brusnih kolutov: a) FSGW, b) HGGW 15 °, c)
HGGW 30 ° in d) HGGW 45 ° (dimenzije v milimetrih), e) izome-
tri~na perspektiva HGGW 15 °
the conventional grinding wheel and helically grooved
grinding wheels (HGGWs).
The helical grooves at the angles of 15 °, 30 °, and 45 °
on the circumference of the grinding wheels were cut in
with a universal milling machine using a cut-off disc in
the angle grinder. The HGGWs were made by cutting
radial grooves with a helical profile. These grinding
wheels consisted of 24 equal grooves on the circumfe-
rence of a grinding wheel. The width and depth of each
groove was 2.6 mm and 3 mm, respectively. The forming
of the helical grooves on a grinding wheel with a cut-off
disc is shown in Figure 2. FSGW and HGGWs with
different helix angles are shown in Figure 3.21
2.2 Materials, grinding parameters and the measure-
ment procedure
The materials used in this study were AISI 1050,
AISI 4140, and AISI 7131; the chemical compositions
and hardness values of the materials are listed in Table 1
and the dimensions of the workpieces are shown in Fig-
ure 4. Grinding-test bar specimens with a diameter 39
mm and length 170 mm were prepared by turning them
directly from the as-received materials.
The grinding conditions and dimensions of the spe-
cimens were the same for both methods. In the grinding
experiments, a horizontal spindle-type cylindrical-grind-
ing machine with aluminum-oxide grinding wheels with
dimensions of 400 mm × 40 mm × 127 mm and a con-
stant wheel speed of 1570 r/min was used. Furthermore,
these elements were common. 60-K-6-V indicates a
wheel grain size of 60, hardness K, structure 6 and a
vitrified bond. The selection was based on a wide indu-
strial application of grinding wheels. Before each grind-
ing experiment, the grinding wheel was dressed using a
single-point diamond dresser to produce a sharp, clean
wheel surface. During the grinding process, a water-
soluble metalworking fluid diluted to 1 : 5 was supplied
to the grinding zone, and the coolant flow rate from the
outlet was 8 L/min. During all the experiments, the
M. GAVAS et al.: EFFECTS OF VARIOUS HELICALLY ANGLED GRINDING WHEELS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 865–870 867
Figure 4: Dimensions and measurement points for the surface roughness and roundness of a grinding workpiece used in the experiments
Slika 4: Dimenzije in to~ke merjenja hrapavosti povr{ine in okroglosti na obdelovancu, uporabljenem pri preizkusih
Table 1: Chemical compositions (w/%) and hardness of the materials used in the experiment
Tabela 1: Kemijske sestave (w/%) in trdote materialov, uporabljenih pri preizkusih
Materials C Si Mn P S Cr Mo Hardness (HRB)
AISI 1050 0.47–0.55 0.15–0.35 0.60–0.90 0.04 0.05 – – 96
AISI 4140 0.35–0.44 0.15–0.40 0.60–0.90 0.035 0.035 0.80–1.10 – 99
AISI 7131 0.14–0.19 0.15–0.40 1.00–1.30 0.035 0.035 0.80–1.10 0.15–0.25 115
Figure 3: Photograph of FSGW and HGGWs
Slika 3: Posnetek FSGW in HGGW
Figure 2: Forming of helical grooves on a grinding wheel with a cut-
off disc
Slika 2: Izdelava vija~nega utora na brusnem kolutu z rezalno
plo{~ico
grinding length (170 mm), workpiece diameter (39 mm),
depth of cut (10 μm) and feed rate (1.57 mm/r) were
used. In addition, each work material was ground for 5
min before measuring the surface roughness and round-
ness error. The experimental set-up is shown in Figure 5.
The measurement points for the surface roughness
and waviness of the work material are shown in Figure
4. Each test specimen was measured on 12 different
points (A1, A2, A3), (B1, B2, B3), (C1, C2, C3), and
(D1, D2, D3) for the surface roughness. For the round-
ness, each test specimen was measured on three points
(1, 2, and 3). In this study, the surface roughness (Ra, the
arithmetic average) of each machined workpiece was
measured using a surface roughness tester (Mahr Per-
then) with a 4 mm cut-off length. In addition, a recorder
that transfers the obtained values onto a graphic was also
used. The roundness was measured with a round test
instrument (Mitutoyo RA-114) in the same locations.
3 RESULTS AND DISCUSSION
The experimental results for the surface roughness of
three steels at the same grinding conditions using both
FSGW and HGGWs are shown in Figure 6. Each sur-
face-roughness value was obtained by averaging 12
measurements. One of these 12 measurements was used
to represent the surface roughness of the ground part. It
is obvious that both the material type and the grinding
method have an effect on the surface roughness. The
highest surface roughness was obtained for the AISI
7131 steel ground with FSGW. The effect of the material
type on the surface roughness is often attributed to the
difference in the hardness of different materials.5
Generally, the surface roughness obtained with
HGGWs is better than that obtained with FSGW. It is
apparent that HGGW 45 ° gives better surface finishes of
the three steels than FSGW. As the grooves on the cir-
cumference of a grinding wheel decrease the grinding
force, temperature15 and contact length between the
workpiece and the grinding wheel, and enough cutting
fluid is being delivered to the wheel-workpiece inter-
face,13 it is, consequently, expected that HGGWs are
better in terms of the ground-surface quality.21,22
Figures 6 and 7 show a comparison of the roundness
of the ground workpieces produced with FSGW and
HGGWs under the same grinding conditions. It is ob-
vious that both the material type and the grinding me-
thod have an effect on the roundness. Generally, the
roundness obtained with HGGWs is better than that ob-
tained with FSGW. When the roundness values obtained
M. GAVAS et al.: EFFECTS OF VARIOUS HELICALLY ANGLED GRINDING WHEELS ...
868 Materiali in tehnologije / Materials and technology 49 (2015) 6, 865–870
Figure 5: Experimental set-up
Slika 5: Eksperimentalni sestav
Figure 7: Profiles of the ground surfaces produced with FSGW and
HGGWs
Slika 7: Profil bru{enih povr{in pri FSGW in HGGW
Figure 6: Ground surface roughness and roundness for FSGW and
HGGWs
Slika 6: Bru{ena povr{inska hrapavost in okroglost pri FSGW in
HGGW
with FSGW and HGGWs are compared, it is clearly seen
that the roundness error increased for the AISI 7131
steel, while the AISI 1050 and AISI 4140 steels behave
in an identical manner during the grinding in view of the
roundness error. For comparison, the roundness values of
the three steels ground in the FSGW and HGGW pro-
cesses are shown in Figure 7. Among the HGGWs,
HGGW 30 ° seemed to generate the lowest roundness.
Consequently, the grinding with helically grooved
wheels (HGGW 15 ° and HGGW 30 °) increases the
roundness quality, while the AISI 7131 steel showed
conflicting behavior, namely, the roundness increased for
the this material.
The AISI 7131 steel and HGGW 45 ° showed unex-
pected behavior; namely, the surface roughness and
roundness increased for this material. This is explained
with the segmentation of the grinding wheel, reducing
the number of static and kinematic cutting edges and,
hence, the rubbing regime, which is one of the important
mechanisms for improving the surface roughness.
Additionally, the uncut-chip thickness increases with the
decrease in the kinematic cutting edges. Tawakoli and
Azarhoushang20 suggested the following relation (Equa-
tion (1)) involving Rt (the distance between the highest
peak and the deepest valley of the profile of the total
evaluation length or, in other words, the total height of
the profile) and the uncut-chip thickness hcu:
R
h
a
v
v C r dt
∞ ≈
⎛
⎝
⎜
⎞
⎠
⎟cu
e
ft
c kin s
4 3
1 3
2 3
1/
/
/
(1)
where hcu is the uncut-chip thickness, ae is the depth of
cut, vft is the feed speed, vc is the cutting speed, ds is the
wheel diameter, r is the grain-cutting-point shape factor,
and Ckin is the kinematic cutting-edge density.
4 CONCLUSIONS
In this study, grinding operations of a flat-surface
wheel and helically grooved wheels were performed on
three different steels (AISI 1050, AISI 4140 and AISI
7131) under the same grinding conditions, except for the
grinding-wheel profile. The surface roughness and
roundness obtained from these processes were demon-
strated. Both the material type and the grinding method
have an effect on the surface roughness and roundness.
The experimental results show that, generally, the sur-
face roughness and roundness obtained with a helically
grooved grinding wheel are better than in the case of
conventional grinding. Additionally, among the helically
grooved wheels, HGGW 30 ° seemed to generate a lower
roundness than the other two HGGWs and FSGW.
Consequently, grinding with helically grooved wheels
(HGGW 15 ° and HGGW 30 °) increased the surface
roughness and roundness, while the AISI 7131 steel
showed conflicting behavior, namely, the roundness of
this material was increased.
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870 Materiali in tehnologije / Materials and technology 49 (2015) 6, 865–870
D. MITROVI] et al.: CHARACTERIZATION OF CAST-IRON GRADIENT CASTINGS
871–875
CHARACTERIZATION OF CAST-IRON GRADIENT CASTINGS
KARAKTERIZACIJA LITO@ELEZNEGA GRADIENTNEGA ULITKA
Danijel Mitrovi}1, Primo` Mrvar2, Mitja Petri~2
1Livar d.d., Slovenska cesta 43, 1295 Ivan~na Gorica, Slovenia
2Department of Materials and Metallurgy, Faculty of Natural Sciences and Engineering, University of Ljubljana, A{ker~eva 12,
1000 Ljubljana, Slovenia
Danijel.Mitrovic@livar.si, Primoz.Mrvar@omm.ntf.uni-lj.si, Mitja.Petric@omm.ntf.uni-lj.si
Prejem rokopisa – received: 2014-07-07; sprejem za objavo – accepted for publication: 2014-12-08
doi:10.17222/mit.2014.102
This work deals with the topic of composed castings, also called gradient castings. The characterization of the microstructures
and the subsequently monitored mechanical properties of the composite castings are discussed. The production technology is a
combination of the horizontal centrifugal casting of alloyed white cast iron followed by an intermediate layer of flake grey cast
iron and gravity casting of the core, which occurs in the third sequence. The core was made of spheroidal grey cast iron. The
research was focused on the intermediate layer and a metallographic analysis of the intermediate layer, which is crucial for the
defects and the lifetime of rolls.
The TCW program was used for the thermodynamic calculations of equilibrium phases in order to prove that the phases present
in the microstructures were determined with light and scanning electron microscopy (SEM). Dilatometry in the solid state was
done for all three layers to study the behavior of composed rolls during the thermal loading. The densities of microstructural
constituents were calculated with the TAPP 2.2 software to explain the distribution of the phases in the intermediate layer.
Additionally, the linear hardness and the tensile strength at room and higher temperatures were measured.
Keywords: composite casting, centrifugal casting, gravity casting, intermediate layer, microstructure
V predstavljenem delu smo opredelili in karakterizirali mikrostrukturne sestavine ter spremljali mehanske lastnosti
kompozitnega ulitka. Tehnologija izdelave je kombinacija horizontalnega in centrifugalnega litja legirane bele litine in kasneje
sive litine (dve sekvenci) ter gravitacijskega litja jedra, ki se pojavi v tretji sekvenci. Jedro je izdelano iz sive litine s kroglastim
grafitom. Raziskava je bila usmerjena na vmesno plast, ki je ulita iz sive litine z lamelnim grafitom in je klju~nega pomena za
pojav napak na valjih oziroma za trajnostno dobo valjev.
Uporabili smo naslednje preiskovalne metode: program TCW za termodinami~ni izra~un faznih ravnote`nih faz v povezavi s
svetlobno in elektronsko mikroskopijo, dilatometrijo v trdnem stanju za ugotavljanje razlik pri kr~enju in {irjenju vseh plasti,
izra~un gostot mikrostrukturnih komponent s programom TAPP 2.2 za dolo~evanje razporeditve karbidov med strjevanjem,
linearne meritve trdot ter natezne trdnosti pri sobni in povi{anih temperaturah.
Klju~ne besede: kompozitni ulitek, centrifugalno litje, gravitacijsko litje, vmesna plast, mikrostruktura
1 INTRODUCTION
Centrifugal pouring technology is a casting process,
where metal can be poured and solidified in a rotating
permanent mold under the influence of the centrifugal
force.1 The direction of solidification in the centrifugal
process differs from that in the sand casting. Due to a
rapid transfer of heat to the permanent mold, crystalli-
zation starts on the outer surface of the casting and pro-
gresses towards the inside. The result is a fine-grained
surface crust. Further solidification towards the interior
takes place with the growth of dendrite crystals.2 Den-
drites can have an equiaxed or columnar shape, depend-
ing on the thermal gradient during the solidification.3
A cast roll can be produced as a gradient casting
composed of a hard steel shell as the working layer and a
tough core. The microstructure of the working layer of a
roll after the solidification depends on the chemical
composition of the alloy. The main microstructural
constituent is austenite () that solidified in the form of
dendrites, followed by a eutectic, which consists of
austenite and carbides.4 The type of the carbides is
determined by the chemical composition. Usually the
alloys for rolls are alloyed with Cr, Ni, W, Mo and V.5,6
For a chromium-white-cast-iron working layer, it is
desirable to have as little as possible of the retained
austenite in the matrix and no pearlite phase in the
microstructure.
In the as-cast state, the matrix contains a substantial
proportion of the residual austenite (30–60 %) that has to
decompose with a single or multistage heat treatment in
order to achieve the required microstructure, containing
small and evenly distributed M23C6 type carbides in the
-metallic matrix.7 Due to their high hardness and uni-
form distribution in the matrix, secondary carbides are of
a great importance for the wear resistance. The target
mechanical properties are obtained with the heat treat-
ment, where a casting is heated to the austenitizing
temperature and control cooled to room temperature.
Such a treatment allows a good control over the segre-
gation of the secondary carbides in a temperature range
from 800 °C to 1050 °C.
Rolls can be produced by forging or by casting. Cast
rolls are often produced as gradient castings made of
more than one material. The outer layer of a casting, or
Materiali in tehnologije / Materials and technology 49 (2015) 6, 871–875 871
UDK 621.74:536.7:669.017.3 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)871(2015)
the working layer, made with horizontal centrifugal
casting, is followed by a centrifugally cast intermediate
layer and the final layer is the gravity-cast core made of
spheroidal cast iron. The working layer is often made of
chromium white cast iron to achieve the hardness and
wear resistance. The intermediate layer is a sort of grey
cast iron, creating a good bond between the working
layer and the core; and the core is made of nodular cast
iron to obtain ductility during the casting.8,9 The bonding
between the working layer and the core is important
since poor bonding can cause bond-related spalls.10
The aim of the paper is to describe the formation and
types of the carbides in the intermediate layer and the
bonding between the working layer and the core. The
bonding is crucial for the defects and lifetime of a roll.
Also, the mechanical properties across the cross-section
are determined and the mechanism of a possible crack
formation is explained.
2 EXPERIMENTAL WORK
The analyses were carried out on the samples taken
from the working layer, the intermediate layer and the
core. Light microscopy, scanning electron microscopy
(SEM) with an EDS analysis, tensile tests at various tem-
peratures, hardness measurements and a dilatometric
analysis were carried out. The densities of the solidified
phases were calculated with the TAPP 2.2 program and
the thermodynamic-phase equilibrium calculations were
performed with the Thermo-Calc software.
The analyses were carried out on the samples taken
from the working layer, the intermediate layer and the
core, at a depth of approximately 80 mm into the roll.
Metallographic samples were taken from all three layers
and metallographically prepared by grinding, polishing
and etching with 2 % Nital. The samples for the tensile
tests were taken from the working layer in the
longitudinal direction of a roll. The test pieces had round
cross-sections with a diameter of 10 mm. The test pieces
for dilatometric analyses were taken from all three layers
of a roll as indicated in Figure 1. The test pieces had
dimensions of 5 mm × 5 mm × 20 mm. The dilatometric
analyses were carried out at a heating rate of 20 K/s, up
to 1100 °C, at which the samples were held for 5 min
and then cooled at the same rate to 400 °C.
The microstructure was analyzed with light micro-
scopy using an Olympus BX61 microscope and with
scanning electron microscopy using a JEOL JSM–5600
electron microscope equipped with EDS. The tensile
tests at elevated temperatures were carried out according
to the EN ISO 6892 standard, using an Instron 1255
machine at room temperature and at (300, 400, 500, 600,
700 and 800) °C. Rockwell-hardness measurements were
performed using an Instron Tukon 2100 B machine and a
dilatometric analysis performed by a BÄHR DIL 801
instrument, on the area from the surface of a roll up to 80
mm into the depth of the roll, every 2–5 mm. The den-
sities of the solidified microstructural phases were calcu-
lated with the TAPP 2.2 program and the thermo-
dynamic-phase equilibrium calculations were performed
with the Thermo-Calc software.
Chemical compositions of all three layers are pre-
sented in Table 1.
3 RESULTS AND DISCUSSION
Figure 2 presents the boundary area of the working
layer, the intermediate layer and the core. The micro-
structure of the working layer consists of austenite and
carbides since the alloy is rich in chromium and molyb-
denum. The intermediate layer, which merges with the
core, is highly rich in M7C3 carbides. During the first
stage of the solidification of the intermediate layer, the
formation of primary austenite crystals occurs in the
liquid melt, which, due to the centrifugal force and the
density higher than the rest of the melt, starts to move in
the direction of the working layer. The melt of the
intermediate layer re-melts and merges with a thin layer
D. MITROVI] et al.: CHARACTERIZATION OF CAST-IRON GRADIENT CASTINGS
872 Materiali in tehnologije / Materials and technology 49 (2015) 6, 871–875
Table 1: Chemical analysis of a gradient casting (w/%)
Tabela 1: Kemijska analiza gradientnega ulitka (w/%)
C Si Mn P S Cr Ni Mo Mg Cu Sn Al V Ti W Co Fe
Working layer 2.799 0.703 0.965 0.030 0.037 16.681 1.433 1.154 0.003 0.081 0.000 0.0000 0.296 0.000 0.000 0.000 75.818
Intermediate layer 3.118 1.034 0.342 0.026 0.010 0.129 0.247 0.029 0.000 0.039 0.005 0.0013 0.012 0.006 0.009 0.018 94.965
Core 3.002 2.734 0.366 0.031 0.007 0.131 0.247 0.029 0.099 0.039 0.007 0.0229 0.013 0.007 0.009 0.017 93.198
Figure 1: Presentation of the sampling for a dilatometric analysis
(sample 1 – working layer, sample 2 – working layer, sample 3 – inter-
mediate layer, sample 4 – core)
Slika 1: Predstavitev vzor~enja za dilatometrsko analizo (vzorec 1 –
delovna plast, vzorec 2 – delovna plast, vzorec 3 – vmesna plast,
vzorec 4 – jedro)
of the working layer and some carbide-promoting ele-
ments, especially chromium, dissolve in the intermediate
melt, causing the formation of carbide sand due to a
lower density in comparison with , deposited at the
interface of the intermediate layer and the core. A large
amount of carbides can be clearly seen on Figure 2. A
sufficient solidification interval and lower cooling rates
cause an intermediate layer stratified in this way, which
is an undesirable microstructure. The carbides that are
not evenly dispersed in the metal matrix represent the
brittle layer of the casting.
Figure 3 shows the isopleth phase diagram of the
working-layer material. The solidification starts with the
solidification of austenite, followed by a eutectic reac-
tion with the solidification of the M7C3 type carbides
between 1260 °C and 1230 °C. The precipitation of the
M23C6 type carbides takes place at 820 °C.
Figure 4a shows a SEM microphotograph of the
working layer, where the large particles are the carbides
that solidified during the eutectic reactions. The EDS
analysis shows that this phase includes the M7C3 type
carbides. Smaller particles in the matrix are the M23C6
type carbides that precipitated in the solid state from the
solid solution of austenite. According to Vitry et al.6 the
M7C3 and M23C6 type carbides should be chromium
carbides, but from the EDS analyses it is clear that these
carbides are mainly chromium and iron carbides with
some small amounts of V and Mo. The types of carbides
are the same as predicted by the thermodynamic calcula-
tions from Figure 3. It is clear that practically the whole
austenite was transformed into martensite during the heat
treatment. The intermediate layer also contains carbides,
determined with the EDS analysis to be of the M7C3
type, as seen on Figure 4b. There is a small amount of
smaller particles of secondary carbides M23C6 distributed
D. MITROVI] et al.: CHARACTERIZATION OF CAST-IRON GRADIENT CASTINGS
Materiali in tehnologije / Materials and technology 49 (2015) 6, 871–875 873
Figure 2: Macrostructure of the working layer, the intermediate layer
and the core
Slika 2: Makroposnetek delovne plasti, vmesne plasti in jedra
Figure 4: Microstructures of all three layers: a) SEM microphoto-
graph of working layer, b) SEM microphotograph of intermediate
layer and c) light microphotograph of core
Slika 4: Mikrostrukture vseh treh plasti: a) SEM-mikroskopija delov-
ne plasti, b) SEM-mikroskopija vmesne plasti ter c) svetlobna mikro-
skopija jedra
Figure 3: Isopleth phase diagram of the working layer with marked
chemical composition
Slika 3: Izopletni fazni diagram delovne plasti z ozna~eno kemijsko
sestavo
in the martensite matrix; this amount is much lower since
the concentration of the carbide-promoting elements is
much lower than in the working layer. Figure 4c shows a
light microphotograph of the core in the polished state,
where graphite can be seen in the iron matrix. Graphite
is expected to be in the nodule-like form, but here this is
not the case, probably due to an insufficient Mg-treat-
ment and a burn-off of Mg during the long solidification
time of the core.
Figure 5 presents EDS spectrums of the analyzed mi-
crostructural constituents marked on Figures 4a and 4b.
With the help of the TAPP 2.2 program, the densities
of the eutectic M7C3 carbides and austenite at the refe-
rence solidification temperature were calculated. The
density of carbides M7C3 at the temperature of precipi-
tation is 6.738 kg/dm3; austenite has a density of a 6.99
kg/dm3 at the temperature of precipitation. Given the
relative differences in the density between the austenite,
the carbides and the melt in the solidification stage of the
intermediate layer, it seems that the stratification of both
microstructural ingredients takes place. In the first stage,
austenite dendrites appear and they are pushed by the
centrifugal forces and a higher density in the direction
towards the working layer. When the temperature of the
remaining melt falls within the scope of the eutectic
solidification, leading to the development of a nucleation
and growth of the carbides they are pushed in the direc-
tion of the interface between the intermediate layer and
the core since the carbides have a lower density than the
austenite. The difference in the density between the
austenite and carbides is only 0.3 kg/dm3, but it is much
more significant at a centrifugal force of 120 G leading
to an inhomogeneous microstructure development.
Rockwell-hardness measurements were carried out in
the area between the surface of the casting and the depth
of 80 mm. Figure 6a presents the hardness of the work-
ing layer, which is around 61 HRC up to the depth of
60 mm where the intermediate layer starts. In this layer
the hardness starts to descend and continues to descend
in the core too where it drops to only 32 HRC.
The tensile tests (Figure 6b) of the working layer at
different temperatures show that the tensile strength is
lowered by about 10 % at 400 °C and, at higher tempera-
tures, it declines even faster and reaches only 200 MPa at
700 °C. This means that the working layer has a ten-
dency to from cracks during the cooling of the casting,
D. MITROVI] et al.: CHARACTERIZATION OF CAST-IRON GRADIENT CASTINGS
874 Materiali in tehnologije / Materials and technology 49 (2015) 6, 871–875
Figure 6: Mechanical properties: a) hardness and b) tensile strength at
different temperatures
Slika 6: Mehanske lastnosti: a) trdota ter b) natezna trdnost pri raz-
li~nih temperaturah
Figure 5: EDS spectrums of phases: a) M7C3 carbide, spot 1 on
Figure 4b, b) martensite, spot 3 on Figure 4b, c) carbides M23C6,
spot 6 on Figure 4a and d) martensite with carbide M23C6, spot 4 on
Figure 4b
Slika 5: EDS-spektri faz: a) M7C3 karbida, obmo~je 1 na sliki 4b, b)
martenzit, obmo~je 3 na sliki 4b, c) karbid M23C6, obmo~je 6 na sliki
4a ter d) martenzit s karbidi M23C6, obmo~je 3 na sliki 4b
since the core has a much higher temperature, causing
the tensile stresses in the working layer.
Figure 7 presents a dilatometric analysis of four sam-
ples, two from the working layer, one from the inter-
mediate layer and one from the core. We can see that the
working layer has the lowest dilatation in the tempera-
ture range from room temperature to up to 1100 °C. The
intermediate layer has a slightly higher dilatation at the
highest temperature, but the core has the highest dilata-
tion. These differences in dilatation cause high internal
stresses during the cooling of the gradient casting. It is
clear that when the casting is being cooled from the sur-
face side, the working layer is shrinking faster than the
core, which causes the tensile stresses in the working
layer and these might lead to a crack formation. A simi-
lar situation takes place during the heat treatment, where
the whole casting is heated to the austenitizing tempe-
rature and the core expands more than the working layer,
causing tensile stresses again. On Figure 7 we can see
that the quantitative difference in dilatation of the sam-
ples is 0.055 mm at 1000 °C which is not an insignificant
value.
4 CONCLUSIONS
The solidification of three different layers was deter-
mined in the present work. It is clear that the interme-
diate layer re-melts the working layer and some
carbide-promoting elements dissolve in the intermediate
liquid layer, causing the formation of the M7C3 type car-
bides during the solidification. The working layer con-
sists also of secondary carbides of the M23C6 type. This
finding is confirmed by a thermodynamic calculation,
which shows that the solidification of austenite is
followed by the solidification of the M7C3 type carbide
eutectic and a further precipitation of the M23C6 carbides.
The M7C3 type carbides have a lower density than
austenite; as a result, at a slow solidification rate of the
intermediate layer, the formed carbides are pushed by
high centrifugal forces into the inner part of the layer. In
this way, an inhomogeneous microstructure is obtained,
which is inappropriate for the lifetime of the roll.
The tensile strength of the working layer is not
changed until the temperature reaches 500 °C, and then it
is rapidly lowered, which can lead to a casting failure if
such working temperatures occur also during the
lifetime.
The hardness of the gradient-casting cross-section is
on a decrease towards the intermediate layer, which is
the result of a lower concentration of primary and secon-
dary carbides.
A dilatometric analysis showed big differences in the
linear thermal-expansion coefficients of different layers.
The difference in dilatation between the working layer
and the core is 0.055 mm at 1000 °C. Such differences in
the contraction or expansion during the cooling and
heating of the solidification process, heat treatment or
during the working cycles of a roll can cause stresses
that may exceed the ultimate tensile strength of the
working layer, initiating a crack that may result in a roll
failure.
5 REFERENCES
1 Manual for centrifugal casting of rolls, Valji [tore, Valji [tore, d. o.
o., [tore 1986
2 J. P. Breyer, G. Walmag, Metallurgy of High Chromium-Molybde-
num White Iron and Steel Rolls, Rolls for the Metalworking
Industries, Iron and Steel Society, Warendale 2002, 29–40
3 W. Wolczynski, E. Guzik, W. Wajda, D. Jedrzejczyk, B. Kania, M.
Kostrzewa, CET in Solidifying Roll – Thermal Gradient Field Analy-
sis, Archives of Metallurgy and Materials, 57 (2012) 1, 105–117,
doi:10.2478/v10172-011-0159-9
4 M. Yamamoto, I. Narita, H. Miyahara, Fractal Analysis of Solidifica-
tion Microstructure of High Carbon High Alloy Cast Roll Manu-
factured by Centrifugal Casting, Tetsu To Hagane – Journal of the
Iron and Steel Institute of Japan, 99 (2013) 2, 72–79, doi:10.2355/
tetsutohagane.99.72
5 M. Kang, Y. Suh, Y. J. Oh, Y. K. Lee, The effects of vanadium on the
microstructure and wear resistance of centrifugally cast Ni-hard rolls,
Journal of Alloys and Compounds, 609 (2014), 25–32, doi:10.1016/
j.jallcom.2014.04.184
6 V. Vitry, S. Nardone, J. P. Breyer, M. Sinnaeve, F. Delaunois, Micro-
structure of two centrifugal cast high speed steels for hot strip mills
applications, Materials and Design, 34 (2012), 372–378,
doi:10.1016/j.matdes.2011.07.041
7 F. Gologranc, Preoblikovanje, Vol. 1, Fakulteta za strojni{tvo, Ljub-
ljana 1991
8 G. Rivera, P. R. Calvillo, R. Boeri, Y. Houbaert, J. Sikora, Exami-
nation of the solidification macrostructure of spheroidal and flake
graphite cast irons using DAAS and ESBD, Materials Characteri-
zation, 59 (2008) 9, 1342–1348, doi:10.1016/j.matchar.2007.11.009
9 Y. Bai, Y. Luan, N. Song, X. Kang, D. Li, Y. Li, Chemical Com-
positions, Microstructure and Mechanical Properties of Roll Core
used Ductile Iron in Centrifugal Casting Composite Rolls, Journal of
Materials Science and Technology, 28 (2012) 9, 853–858,
doi:10.1016/S1005-0302(12)60142-X
10 Roll Failures Manual, Hot Mill Cast Work Rolls, 1st ed., CAEF – The
European Foundry Association - Roll Section, Duesseldorf 2002, p.
10
D. MITROVI] et al.: CHARACTERIZATION OF CAST-IRON GRADIENT CASTINGS
Materiali in tehnologije / Materials and technology 49 (2015) 6, 871–875 875
Figure 7: Dilatation curves of the samples at different layers
Slika 7: Dilatacijske krivulje vzorcev razli~nih plasti

D. VOJTÌCH et al.: COMPARATIVE MECHANICAL AND CORROSION STUDIES ON MAGNESIUM ...
877–882
COMPARATIVE MECHANICAL AND CORROSION STUDIES ON
MAGNESIUM, ZINC AND IRON ALLOYS AS BIODEGRADABLE
METALS
PRIMERJALNA [TUDIJA MEHANSKIH IN KOROZIJSKIH
LASTNOSTI BIORAZGRADLJIVIH ZLITIN MAGNEZIJA, CINKA
IN @ELEZA
Dalibor Vojtìch, Jiøí Kubásek, Jaroslav ^apek, Iva Pospí{ilová
Department of Metals and Corrosion Engineering, Institute of Chemical Technology, Technická 5, 166 28 Prague 6, Czech Republic
Dalibor.Vojtech@vscht.cz
Prejem rokopisa – received: 2014-07-30; sprejem za objavo – accepted for publication: 2014-12-11
doi:10.17222/mit.2014.129
In this paper, selected magnesium, zinc and iron biodegradable alloys were studied as prospective biomaterials for temporary
medical implants like stents and fixation devices for fractured bones. Mechanical properties of the alloys were characterized
with hardness and tensile tests. In-vitro corrosion behavior was studied using immersion tests in a simulated physiological
solution (SPS, 9 g/L NaCl) to roughly estimate the in-vivo biodegradation rates of implants. It was found that the Mg and Zn
alloys were limited by a tensile strength of 370 MPa, while the tensile strength of the Fe alloys achieved 530 MPa. The main
advantage of the Mg alloys is that their Young’s modulus of elasticity is similar to that of the human bone. However, the
corrosion tests revealed that the Mg-based alloys showed the highest corrosion rates in the SPS, ranging between 0.6 mm and
4.0 mm per year, which is above the tolerable degradation rates of implants. The corrosion rates of the Zn alloys were between
0.3 mm and 0.6 mm per year and the slowest corrosion rates of approximately 0.2 mm per year were observed for the Fe alloys.
The results indicate that all three kinds of alloys meet the mechanical requirements for the load-bearing implants. From the
corrosion-behavior point of view, the Zn- and Fe-based “slowly corroding” alloys appear as promising alternatives to the
Mg-based alloys.
Keywords: biodegradable metal, magnesium, zinc, iron, mechanical properties, corrosion
^lanek obravnava {tudij izbranih magnezijevih, cinkovih in `elezovih potencialnih biorazgradljivih zlitin kot obetajo~ih
biomaterialov za za~asne medicinske vsadke, kot so opornice in pripomo~ki za utrjevanje zlomljenih kosti. Mehanske lastnosti
zlitin so bile dolo~ene z merjenjem trdote in z nateznimi preizkusi. Korozijski preizkusi in vitro so bili izvr{eni s pomakanjem v
simulirano fiziolo{ko raztopino (SPS, 9 g/L NaCl) za grobo oceno in vivo hitrosti biorazgradnje implantatov. Ugotovljeno je, da
so zlitine Mg in Zn omejene z natezno trdnostjo 370 MPa, medtem ko je natezna trdnost Fe-zlitin dosegla 530 MPa. Glavna
prednost Mg-zlitin je v podobnem Young-ovem modulu elasti~nosti, ki je podoben kot pri ~love{ki kosti. Vendar pa so
korozijski preizkusi pokazali pri zlitinah na osnovi Mg najve~je korozijske hitrosti v SPS, med 0,6 mm in 4,0 mm na leto, kar je
nad dopustno hitrostjo degradacije implantata. Korozijske hitrosti Zn zlitin so bile med 0,3 mm in 0,6 mm na leto, najni`je
hitrosti korozije, okrog 0,2 mm na leto, pa so bile opa`ene pri Fe-zlitinah. Rezultati so pokazali, da vse tri vrste zlitin ustrezajo
mehanskim zahtevam za obremenjene implantate. S stali{~a korozijskega vedenja so zlitine na osnovi Zn in Fe “zlitine s
po~asno korozijo” in se ka`ejo kot obetajo~e nadomestilo za zlitine na osnovi Mg.
Klju~ne besede: biorazgradljive kovine, magnezij, cink, `elezo, mehanske lastnosti, korozija
1 INTRODUCTION
Metallic biomaterials have been used in bone and
joint replacements, fractured-bone fixation devices,
stents, dental implants, etc., for a long time. The
advantage of metals over polymers or ceramics is in
higher strength and toughness. In addition, metals can be
simply processed with the established technologies like
casting, forming, powder metallurgy, machining. The
most important metallic biomaterials in the current use
are stainless steels (SUS 316L), titanium alloys (Ti,
Ti-6Al-4V, Ti-6Al-7Nb), cobalt alloys (Co-Cr-Mo),
superelastic Ni-Ti, noble-metal alloys (Au, Pd, dental
amalgams – Hg-Ag-Cu-Sn).1 All these kinds of materials
show a high corrosion resistance to human-body fluids
due to the noble nature and/or spontaneous passivation
and are, therefore, considered as bio-inert materials.2
Besides the bio-inert materials, biodegradable mate-
rials have attracted a great attention. The term biodegra-
dability means that a material progressively corrodes and
degrades in the body environment.3,4 Products of this
degradation are not toxic, allergic or carcinogenic and
they are readily excreted by the human body.3 Biode-
gradable materials can be used for the implants whose
functions in the human body are only temporary, like
fixation devices (screws, plates) for fractured bones and
stents.3–5 When using inert biomaterials in bone-fixation
devices, a second surgery is often necessary to remove
them after the healing process of the bone has
completed. In contrast, a biodegradable material slowly
degrades in the human body and is progressively re-
placed by the growing tissue. No second surgery is
needed which significantly reduces the inconvenience to
Materiali in tehnologije / Materials and technology 49 (2015) 6, 877–882 877
UDK 669.721.5:669.55:620.193:620.17 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)877(2015)
the patient, morbidity and health cost. Among the bio-
degradable materials, polymeric materials (for example
poly-lactic acid – PLA) are commonly used at present,
but their disadvantages are a low mechanical strength
and hardness.5 For this reason, research and development
activities all over the world are also focused on metallic
biodegradable alloys with a higher strength, hardness
and toughness as compared to the polymers.
Biodegradable alloys should have a good biocompati-
bility with the human body tissues. This basic require-
ment limits a number of possible candidates to three
metals, magnesium, zinc and iron.3
Magnesium is generally considered as a relatively
non-toxic metal. It is essential for proper biological
functions of the human body. Its recommended daily
value is about 400 mg.6 Magnesium supports the growth
of the bone tissue, heart functions, the neurologic
system, etc.7–10 Overdoses of magnesium are unlikely to
occur because the metal absorption is efficiently con-
trolled by the metabolism and excess amounts are ex-
creted by the kidneys.10 There are many studies covering
the mechanical, corrosion, in-vitro and in-vivo biocom-
patibility of the Mg alloys.11,12 On the basis of these
studies, magnesium alloys are generally believed to show
a good combination of mechanical performance and bio-
compatibility depending on the actual alloying elements
present. However, the main drawbacks of most of the
investigated biodegradable Mg alloys are excessive
in-vivo corrosion rates.4–10
Iron is also an essential element for proper biological
functions, mainly for the transfer of oxygen by blood.13
The recommended daily value of Fe is about 10 mg.6
Regarding the biocompatibility of iron-based alloys,
there are a number of reported results,14–16 but they are
often controversial. In order to explain the discrepancies
between biocompatibility tests, more in-vitro and in-vivo
experiments are needed.
Zinc supports the immune system, the proper func-
tions of taste, smell, etc.11,17 Like Mg, Zn is also a com-
ponent of many food supplements, therefore, it is consi-
dered as relatively non-toxic. Its recommended daily
value is about 40 mg, but short-term overdoses of up to
100 mg do not cause significant health problems18. Zinc
has been considered as a prospective biodegradable im-
plant material only for a relatively short time;19 therefore,
the in-vitro and in-vivo biocompatibility tests with zinc
alloys are very limited. However, the tests reported
recently indicate a good biocompatibility of Zn.20
In the present work, magnesium, zinc, iron and their
alloys are studied with respect to their mechanical and
corrosion properties. Appropriate alloying of Mg, Zn and
Fe can positively modify their mechanical, corrosion and
physical properties, which are important for potential
medical applications. In the available literature many
alloying elements are proposed for these purposes,3–19 but
in this study Mg-RE (RE = rare earth metals, Gd, Nd,
Y), Zn-Mg and Fe-Mn based alloys were selected, be-
cause all these alloying systems are generally considered
as relatively safe and acceptable for a potential medical
use.3
2 EXPERIMENTS
Chemical compositions of the investigated Mg, Zn
and Fe alloys are summarized in Table 1. Magnesium
and iron alloys were prepared by melting pure metals in
a vacuum-induction furnace under argon. Zinc alloys
were prepared by melting pure metals in air. The alloys
were cast into cast-iron metal molds (Mg, Zn) or sand
molds (Fe) to prepare ingots of 20 mm in diameter and
150 mm in length. Parts of the as-cast ingots of Mg and
Zn alloys were hot extruded at a temperature of 400 °C,
an extrusion ratio of 10 : 1 and a rate of 5 mm/min to
prepare rods of 6 mm in diameter. Ingots of iron alloys
were hot forged at 850 °C into rods of 6 mm in diameter.
Table 1: Designations and chemical compositions of the studied
alloys (w/%)
Tabela 1: Oznake in kemijska sestava preiskovanih zlitin (w/%)
Alloy
designation
Element (w/%)
Mg Zn Fe Gd Nd Y Mn
Mg 	 99.8 - 0.02 - - - -
Mg-3Gd bal. - 0.01 2.7 - - -
Mg-3Gd-1Y bal. - 0.01 2.6 - 0.8 -
Mg-3Nd-4Y bal. - - - 2.8 4.2 -
Zn - 	 99.8 - - - - -
Zn-1Mg 0.9 bal. - - - - -
Zn-2Mg 1.6 bal. - - - - -
Fe - - 	 99.7 - - - -
Fe-30Mn - - bal. - - - 30.5
The microstructures of the alloys were examined
using light (LM) and scanning electron microscopy with
energy dispersive spectrometry (SEM + EDS) and X-ray
diffraction (XRD). Mechanical properties were charac-
terized with the Vickers-hardness (HV 5) and tensile
testing. The tensile tests were carried out on a LabTest
5.250SP1-VM universal loading machine at a deforma-
tion rate of 1 mm/min.
In the human body any implant is exposed to fluids
containing complicated water solutions of inorganic salts
(chlorides, phosphates, etc.), organic compounds (gluco-
se, amino acids, etc.) and biological matter (proteins,
cells, etc.). In this study, the biological environment was
simulated with a simple NaCl solution, in which the
concentration of chlorides is similar to that in the blood
plasma. This simulated physiological solution (SPS)
contained 9 g/L NaCl, 7 × 10–6 of dissolved oxygen and
its initial pH was 6.2 due to dissolved CO2. The corro-
sion behavior was characterized with in-vitro immersion
tests in the SPS. The alloy samples were immersed in the
SPS for 168 h at 37 °C. Afterwards, the corrosion pro-
ducts were chemically removed and the corrosion rates
were then calculated, in mm/year, using the weight
losses measured with a balance.
D. VOJTÌCH et al.: COMPARATIVE MECHANICAL AND CORROSION STUDIES ON MAGNESIUM ...
878 Materiali in tehnologije / Materials and technology 49 (2015) 6, 877–882
3 RESULTS AND DISCUSSION
3.1 Mechanical properties
The Vickers hardness (HV 5), the ultimate tensile
strength (UTS) and the elongation (E) are summarized in
Figure 1. As expected, pure Mg and Zn show the lowest
hardness and strength levels that do not exceed 50 HV 5
and 130 MPa, respectively (Figures 1a and 1b). In con-
trast, pure iron has higher hardness and tensile-strength
values of approximately 100 HV 5 and 300 MPa in the
as-forged state. Figure 1 also demonstrates that both the
hardness and the strength of all three groups of alloys
increase with the increasing concentrations of the alloy-
ing elements due to the solid-solution strengthening and
hardening and due to the influence of the intermetallic
phases present in the structures. Moreover, positive
influences of hot-extrusion or hot-forging operations on
the hardness, the strength and the elongation are ob-
served. The reason for the influence is the fact that hot-
forming steps cause an elimination of casting defects,
dynamic recrystallization and structural refinement, as is
illustrated for the Zn-2Mg alloy in Figure 2. The as-cast
Zn-2Mg alloy (Figure 2a) is composed of primary Zn
dendrites (light) and an interdendritic network of a Zn +
Mg2Zn11 eutectic mixture (dark). A detailed view of the
eutectic is seen in the insert in Figure 2a. Hot extrusion
(Figure 2b) partially breaks down the continuous
eutectic network and the structure becomes oriented
parallel to the extrusion direction. The dynamic recrys-
tallization occurring in the Zn grains (light) results in the
formation of equi-axed and refined Zn grains. The aver-
age grain size in these regions is 15 μm, which is more
than a three-fold reduction in comparison with the
primary dendrites in the as-cast alloy (50 μm).
Figure 1 shows that among the Mg-based alloys, the
highest hardness (114 HV 5) and strength (290 MPa) are
measured for the hot-extruded Mg-3Nd-4Y alloy due to
the presence of the recrystallized and fine-grained struc-
ture. This alloy also shows a good elongation of 13 %
(Figure 1c). Magnesium additions to zinc lead to signi-
ficant hardening and strengthening of the Zn-Mg alloys.
The hot-extruded Zn-2Mg alloy shows the highest hard-
ness (95 HV 5) and tensile strength (367 MPa). But the
elongation of this alloy is only 6 %. On the other hand,
the Zn-1Mg alloy exhibits a slightly lower tensile
strength (301 MPa) but a considerably higher plasticity
(elongation of 13 %). As it was expected, the hot-forged
Fe-30Mn alloy exhibits the highest hardness (175 HV 5)
and strength (530 MPa) among all the studied alloys.
The reason is that manganese remains dissolved in -Fe,
stabilizing the austenitic structure to the room tem-
perature (as proved with XRD) and causing significant
solid-solution hardening and strengthening. Moreover,
this material also shows an acceptable elongation of
15 %.
It is important to compare the mechanical characte-
ristics of novel biodegradable alloys shown in Figure 1
with those of today´s commercial biodegradable
polymers, for example, the poly-lactic acid (PLA). It is
known that the tensile strength of the PLA does not
exceed 60 MPa.1 Therefore, all three groups of metallic
biodegradable materials show significantly higher
strength levels than the PLA which is of a great import-
ance for load-bearing implants like fixation screws, nails
or plates. The advantage of the Mg alloys over the Zn
and Fe alloys is a low density (
 3 g/cm3) similar to that
of the bone (
 2 g/cm3) and also a low Young´s modulus
(
 50 GPa). A low modulus is desirable for a proper
D. VOJTÌCH et al.: COMPARATIVE MECHANICAL AND CORROSION STUDIES ON MAGNESIUM ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 877–882 879
Figure 1: a) Vickers hardness (HV 5), b) ultimate tensile strength
(UTS) and c) elongation (E) of the investigated alloys (C – as cast, E –
as hot extruded, F – as hot forged)
Slika 1: a) Trdota po Vickersu (HV 5), b) natezna trdnost (UTS) in c)
raztezek (E) preiskovanih zlitin (C – ulito stanje, E – vro~e ekstru-
dirano, F – vro~e kovano)
transfer of mechanical loading between the implant and
the bone and for a proper healing process of the bone.
The Zn alloys show a strength similar to the Mg alloys
but higher density and modulus of elasticity. The Fe
alloys are characterized by the highest strength, exhibit-
ing also high density and modulus. As demonstrated in
the following section, the advantages of Zn and Fe over
Mg are in their lower corrosion rates.
3.2 Corrosion behavior
Figure 3 summarizes the corrosion rates of the alloys
in the SPS.
It is observed that, of all the materials, pure Mg
corrodes at the highest rate (4 mm per year). The fast
corrosion is caused by the presence of impurities, mainly
Fe, in magnesium (Table 1). It is known that more noble
metallic impurities like Fe, Ni, Cu and Co strongly acce-
lerate the corrosion of Mg by forming cathodic sites and
micro-galvanic cells with the Mg matrix.11,12 All the
Mg-RE alloys studied show slower corrosion rates in the
SPS than Mg. The Mg-3Gd alloy corrodes at the lowest
rate (0.6 mm per year). The rare-earth metals reduce the
corrosion rate by forming RE-Fe intermetallic phases,
which decrease the galvanic effects between the Mg
matrix and cathodic impurities. Figure 3 also indicates
that the corrosion rates slightly increase with an increase
in the total RE concentration. This may be due to higher
volume fractions of the intermetallic phases and the
resulting galvanic effects. In comparison with the Mg
alloys, the Fe and Zn alloys exhibit significantly lower
corrosion rates ranging between 0.2 mm and 0.6 mm per
year. The differences between the three groups of alloys
are related to different mechanisms of the corrosion pro-
cess:
Magnesium is the least noble of the three metals
studied as its standard potential is –2.4 V (vs. SHE).2 It
thus shows a high tendency to dissolve in water solu-
tions. Magnesium corrosion includes anodic metal disso-
lution (Equation (1)) and cathodic water decomposition
(Equation (2)) to form gaseous hydrogen and to alkalize
the solution.11,12 The corrosion rate of a magnesium
implant that is too fast is thus undesirable because both
corrosion products have adverse effects on the biocom-
patibility by negatively influencing the tissue adherence
and healing.21 In the alkaline solutions, surface corrosion
products may form, but these products are broken down
in the presence of Cl- anions in the SPS:
Mg  Mg2+ + 2e– (1)
2H2O + 2e
–  2OH– + H2 (2)
It is important that the corrosion of the Mg alloys is
not controlled by the access of oxidizing species (for
example, dissolved oxygen) to the metal. For this reason,
it may proceed relatively rapidly even in neutral-water
solutions (Figure 3). The chloride ions present in the
SPS generally accelerate the corrosion process.
Iron is the most noble of all the three metals inve-
stigated as its standard potential is –0.4 V (vs. SHE).2
Therefore, its tendency to dissolve is the lowest. The
corrosion process includes anodic dissolution and, in
contrast to Mg, cathodic reduction of dissolved oxygen
D. VOJTÌCH et al.: COMPARATIVE MECHANICAL AND CORROSION STUDIES ON MAGNESIUM ...
880 Materiali in tehnologije / Materials and technology 49 (2015) 6, 877–882
Figure 3: Corrosion rates of the alloys in the SPS measured with
immersion tests
Slika 3: Korozijske hitrosti zlitin, potopljenih v SPS
Figure 2: Microstructures of the Zn-2Mg alloy: a) as cast, b) as hot
extruded (LM, SEM)
Slika 2: Mikrostruktura Zn-2Mg-zlitine: a) ulito stanje, b) vro~e eks-
trudirano (SM, SEM)
(Equations (3) and (4)). The low corrosion rates of the Fe
alloys in the SPS (Figure 3) can be attributed to the fact
that the corrosion of Fe needs dissolved oxygen whose
concentration in the SPS is low. In addition, the corro-
sion of Fe in neutral solutions is accompanied by the
formation of more or less protective corrosion products
on the surface. Another positive feature of iron is that its
corrosion in neutral solutions does not produce gaseous
hydrogen, which would negatively influence the tissue
healing around the implant:
Fe  Fe2+ + 2e– (3)
O2 + 2H2O + 4e
–  4OH– (4)
Zinc nobility is between those of Mg and Fe. The
standard potential is –0.8 V (vs. SHE).2 The corrosion of
zinc shows some similarity with iron (Equations (5) and
(6)) because it is controlled by dissolved oxygen in the
neutral SPS. Therefore, the corrosion process is relati-
vely slow (Figure 3). Zinc is also easily covered with
protective the passive films of corrosion products in
neutral and slightly alkaline solutions. Like in the case of
iron, hydrogen gas is not generally formed during the
corrosion of zinc in neural solutions. All the above
features are good prerequisites for low corrosion rates of
zinc implants in human body fluids:
Zn  Zn2+ + 2e– (5)
O2 + 2H2O + 4e
–  4OH– (6)
Regarding the corrosion process of a biodegradable
implant in vivo, it is important to know which corrosion
rate is acceptable for a particular application. Too fast an
in-vivo degradation of an implant is undesirable because
such an implant would degrade before the completion of
the healing process. In the case of fixation devices of
fractured bones, it is necessary that the implants mecha-
nically fix the bones for a certain minimum period de-
pending on the implant type, design, location, the
surrounding tissue, etc. The requirement may be, for
example, that a fixation screw must keep 95 % of its
original load-bearing capability for at least six weeks
after the implantation.22 In other words, the corrosion of
the screw should not reduce its cross-section by more
than 5 %, providing the maximum acceptable corrosion
rate of 0.4 mm per year. Figure 3 indicates that the Fe
and Zn alloys meet this requirement and that some
Mg-RE alloys approach the acceptable corrosion rate. It
should be noted that real in-vivo environments contain,
in addition to chlorides, like the ones contained by the
SPS used in this study, also various organic and inor-
ganic compounds which, in contrast to chlorides, retard
the corrosion process by forming protective surface
films. For this reason, an in-vivo corrosion would be
probably slower than the corrosion in the simple SPS and
some Mg-RE alloys would thus also fall into the
acceptable range.23
4 CONCLUSIONS
In this study, biodegradable Mg, Fe and Zn alloys
were compared with regard to mechanical and corrosion
properties. It can be concluded that all of these materials
show significantly higher strengths than the commercial
biodegradable material (PLA). They are thus promising
materials for highly loaded implants. In the case of Mg,
there are still concerns regarding high corrosion rates,
hydrogen-gas release and local alkalization. The main
positive feature of the Zn and Fe alloys is their slow
corrosion due to the absence of the gaseous-hydrogen
release. In addition, the Fe alloys show the highest
strength among all the studied materials.
Acknowledgements
The authors wish to thank the Czech Science
Foundation for the financial support of this research
(project no. P108/12/G043).
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882 Materiali in tehnologije / Materials and technology 49 (2015) 6, 877–882
M. VY[VAØIL et al.: MICROSTRUCTURAL CHANGES OF FINE-GRAINED CONCRETE ...
883–888
MICROSTRUCTURAL CHANGES OF FINE-GRAINED CONCRETE
EXPOSED TO A SULFATE ATTACK
MIKROSTRUKTURNE SPREMEMBE DROBNOZRNATEGA
BETONA, IZPOSTAVLJENEGA SULFATU
Martin Vy{vaøil, Patrik Bayer, Markéta Rovnaníková
Brno University of Technology, Faculty of Civil Engineering, Institute of Chemistry, @i`kova 17, 602 00 Brno, Czech Republic
vysvaril.m@fce.vutbr.cz
Prejem rokopisa – received: 2014-07-30; sprejem za objavo – accepted for publication: 2015-01-05
doi:10.17222/mit.2014.138
Sulfate attack is one of the major threats to durability of concrete constructions and it becomes a major destructor of sewage-
collection systems where concrete sewer pipes are exposed to sulfates. The most frequent biodeterioration in sewage pipes is
caused by biogenic sulfuric-acid corrosion. During this attack, the pH of the surfaces of concrete sewer pipes is reduced and
chemical reactions lead to the cracking and scaling of the concrete material, accelerated by the sewage flow. This paper is
focused on the sulfate attack on fine-grained concrete where the effect of a 0.5 % sulfuric-acid solution on the microstructural
changes of various types of concrete after a treatment for a period of 6 months was investigated with mercury intrusion porosi-
metry and scanning electron microscopy. It was found that the total porosity of most samples was decreased after the sulfate
attack, indicated by the products of the sulfate corrosion filling in the pores of the concrete. The smallest changes in the micro-
structure were observed in the samples made from sulfate-resisting cements. The formation of the locations rich in sulfur, iron
and aluminum during the sulfate attack of the concrete was determined by mapping the chemical-element distribution.
Keywords: sulfate attack, porosity, microstructure, ettringite, gypsum
Izpostavitev sulfatom je ena od glavnih gro`enj za obstojnost betonskih struktur in postaja glavni uni~evalec v sistemih zbiranja
odplak, kjer so betonske cevi izpostavljene sulfatom. Najpogostej{a biorazgradnja betonskih cevi za zbiranje odplak je povzro-
~ena z biogensko korozijo z `vepleno kislino. Med tem napadom se pH povr{ine zmanj{a in kemijske reakcije, pospe{ene s
tokom odplak, povzro~ijo pokanje in lu{~enje materiala iz betona. ^lanek obravnava vpliv izpostavitve sulfatov na drobnozrnat
beton, kjer je bil preiskovan u~inek raztopine `veplene kisline 0,5 % na mikrostrukturne spremembe razli~nih vrst betonov po
izpostavitvi 6 mesecev. Uporabljena je bila porozimetrija z vdorom `ivega srebra in vrsti~na elektronska mikroskopija. Ugotov-
ljeno je bilo, da je bila skupna poroznost vzorcev zmanj{ana po u~inkovanju sulfata, kar ka`e, da se pore v proizvodih napolnijo
s produkti sulfatne korozije. Najmanj{e spremembe mikrostrukture so bile opa`ene pri vzorcih izdelanih betonov, odpornih proti
sulfatom. V betonu, izpostavljenem sulfatom, je bil nastanek podro~ij, bogatih z `veplom, `elezom in aluminijem dolo~en z
razporeditvijo kemijskih elementov.
Klju~ne besede: izpostavitev sulfatom, poroznost, mikrostruktura, etringit, sadra
1 INTRODUCTION
Durability of concrete has become a very relevant
issue in construction projects over the last decades. Sul-
fate attack is one of the major threats to the durability of
concrete constructions and it is a major destructor in
sewage-collection systems where concrete sewer pipes
are exposed to sulfates. The sulfates originate from the
waste water as well as from the biogenic activity of
bacteria – microbiologically induced concrete corrosion
(MICC).1 The most frequent biodeterioration in sewage
pipes is caused by biogenic sulfuric-acid corrosion.1,2
During this attack, the pH of the surfaces of concrete
sewer pipes is reduced and chemical reactions lead to the
cracking and scaling of the concrete material accelerated
by the sewage flow. The rate of the loss of a concrete
material can be 3–6 mm per year, thereby the breaking-
out of steel reinforcement can take place, resulting in the
subsequent corrosion of the steel reinforcement.3 From
this aspect, a sulfate attack in the sewage system is very
dangerous, especially in the areas with waste water rich
in sulfates or H2S.
The formation of ettringite (AFt) from calcium sul-
fate (gypsum) and C3A via monosulfate (AFm), accord-
ing to Equation (1), is the main chemical reaction of the
sulfate attack on concrete.4 Gypsum is the primary
product of the chemical sulfate attack on concrete
(formed by the reaction of sulfate anion with calcium
hydroxide):
CSH C A 10H C ASH C AS H2 3 4 12
2CSH H
6 3 32
2+ + → ⎯ →⎯⎯⎯⎯+16
gypsum monosulfate ettringite (1)
Ettringite has a relative low density (1.75 g cm–3) in
comparison to, e.g., the C-S-H phase (2.0 g cm–3) and a
considerably larger volume than the initial compounds,
therefore its formation provides a potential stress in the
hardened cement paste.5 The theoretical volume increase
varies depending on the source of the available alumi-
num. The sources of reactive aluminum are AFm phases
(monosulfate, monocarbonate) and calcium aluminate
originating from the clinker phases (C3A, C4AF). The
ettringite formation can be reduced by lowering the C3A
content in the cement. According to the European stan-
Materiali in tehnologije / Materials and technology 49 (2015) 6, 883–888 883
UDK 691.32:620.193:620.192.47 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)883(2015)
dards EN 196 and EN 197, sulfate-resistant cements are
limited in the C3A content to < 3 % and the Al2O3 con-
tent to < 5 %.
Systematic studies and reviews were done to evaluate
the deterioration processes of the secondary-ettringite
formation due to an external sulfate attack on hydrated
cement paste, mortar and concrete.4,6,7 The results
showed that the amounts of C3A (0–8 %) in the cement
systems do not necessarily protect them from a sulfate
deterioration. In contrast, the w/c ratio of the investigated
samples had a major impact on the failure time of the
samples during 40 years of exposure under real condi-
tions. The permeability (porosity) of concrete samples
has a major influence on the deterioration of the samples.
It is generally agreed that the alumino-ferrite phase
(C4AF) seems to be less important with regard to the
secondary-ettringite formation during a sulfate attack
due to its lower reaction kinetics.8,9 The formation of
secondary ettringite from C4AF in the presence of
gypsum (sulfate) and the formation of aluminum- (AH3)
and iron-hydroxide (FH3) are given in Equation (2):
3C AF CSH xH C AS H 2(A, F)H4 2 6 3 32 3+ + → +12 4 (2)
Besides ettringite, gypsum can also form during a
sulfate attack, especially in highly concentrated sulfate
solutions.10,11 The influence of a gypsum formation on
the performance of cement pastes, mortars and concretes
was studied in detail by many authors.6,12–14 It was
suggested that the secondary-gypsum formation is
related to the amount of alite (C3S) in the cement system
due to the possibility of a portlandite formation. The
transformation of portlandite into gypsum can cause an
expansion, usually later, during the sulfate exposure, and
the softening of the near-surface regions due to the gyp-
sum formation was also observed. The softening was
attributed to the decalcification of the C-S-H phase.
At lower temperatures, thaumasite (CaSiO3·CaSO4·
CaCO3·15H2O; C3SSCH15) is formed in addition to
ettringite as a result of the reaction between C-S-H and
SO42–, CO2 or CO32–, and water. The AFt phase is
structurally similar to ettringite, with Al3+ substituted by
Si4+. Thaumasite is more stable at lower temperatures
since silicon tends to adopt the octahedral co-ordination.
However, thaumasite is formed also at temperatures
around 20 °C and above.15,16 Once thaumasite is formed,
it remains stable up to 30 °C.17 The formation of thauma-
site always needs a source of carbonate which can be
supplied from the limestone contained in the cement
itself. The damage due to the thaumasite formation was
investigated by various authors.18–21
Thus, the interactions of sulfate ions with the cement
matrix result in a disruption of the concrete and a signi-
ficant loss of the mechanical strength and mass, and it
leads to a reduction of the service life of concrete com-
posites.
This paper is focused on the sulfate attack on fine-
grained concrete, investigating the effect of 0.5 %
sulfuric acid (simulating MICC) on the microstructural
changes of various types of concrete after a treatment for
the period of six months. The changes in the microstruc-
ture were determined with mercury intrusion porosi-
metry and scanning electron microscope. The aim of this
study was to compare the resistances of various types of
concrete against the sulfate attack, in view of the micro-
structural changes that may result in a disintegration and
the subsequent loss of the concrete material.
2 MATERIALS AND METHODS
Fine-grained concrete specimens (40 mm × 40 mm ×
160 mm) were prepared with a water-to-binder ratio of
0.45 and three fractions of quartz sand according to
Czech standard CSN 72 1200, with designations PG 1
(< 0.5 mm), PG 2 (0.5–1 mm) and PG 3 (1–2.5 mm), in
the weight-to-binder ratio of 1 : 1 : 1 : 1. Seven concrete
mixtures with different compositions of the binder were
prepared: PC – Portland cement CEM I 42.5 R (100 %);
SRP – sulfate-resisting Portland cement (100 %); SRS –
sulfate-resisting slag cement CEM III/B 32.5 N-SV
(100 %); MK – metakaolin (20 %), Portland cement
(80 %); GL – ground limestone (20 %), Portland cement
(80 %); GBFS – granulated blast-furnace slag (20 %),
Portland cement (80 %) and FA – low-calcium fly ash
(20 %), Portland cement (80 %). The mixtures with the
supplementary cementing materials (metakaolin, lime-
stone, slag and fly ash) corresponded to Portland blended
cements according to European standard EN 197-1. The
chemical compositions and physical properties of the
initial materials are given in Table 1. The specimens
were unmolded 24 h after the casting under laboratory
conditions (t = (22 ± 2) °C, R. H. was (55 ± 5) %) and
M. VY[VAØIL et al.: MICROSTRUCTURAL CHANGES OF FINE-GRAINED CONCRETE ...
884 Materiali in tehnologije / Materials and technology 49 (2015) 6, 883–888
Table 1: Chemical compositions and physical properties of initial ma-
terials in mass fractions, w/%
Tabela 1: Kemijska sestava in fizikalne lastnosti izhodnih materialov
v masnih dele`ih, w/%
Materials PC SRP SRS MK GL GBFS FA
C
he
m
ic
al
co
m
po
si
ti
on
CaO 61.48 60.92 47.53 0.20 54.73 34.81 3.60
SiO2 21.26 21.88 32.84 58.70 0.81 39.78 53.70
Al2O3 5.08 3.65 6.01 38.50 0.32 8.17 24.62
Fe2O3 3.64 4.36 1.54 0.72 0.10 1.54 7.91
SO3 2.42 2.39 2.30 0.02 0.05 0.46 0.96
MgO 0.86 3.15 7.21 0.38 0.38 13.22 1.67
Na2O 0.12 0.41 0.32 0.04 – – 0.85
K2O 0.91 0.86 0.65 0.85 – – 2.62
MnO 0.07 0.11 0.03 – 0.01 0.89 –
TiO2 0.29 0.34 0.41 0.49 – 0.24 1.03
Cr2O3 – – – – – 0.14 –
P2O5 – – – – – 0.03 2.25
L.O.I. 4.17 1.97 0,82 1.67 43.99 1.48 2.82
P
hy
si
ca
l
pr
op
er
ti
es SSA
(m2 kg-1) 360 685 504 13060 390 384 340
SG
(kg m-3) 3120 2650 2950 1070 2700 2810 2300
Note: "–" ... not tested
placed into a water bath for another 27 d. Afterwards, the
specimens were air-dried for 24 h and then the pore
structures of the samples were studied to determine the
total porosity and the pore-size distribution with high-
pressure mercury intrusion porosimetry using a Micro-
meritics PoreSizer 9310 porosimeter, and the microstruc-
tures of samples were observed with a scanning electron
microscope, MIRA3 (TESCAN), equipped with an EDX
probe.
Subsequently, the test samples were covered with a
protective coat to prevent drying and they were placed
into a solution of 0.5 % H2SO4 for a period of six
months. The concentration of the sulfuric acid was
chosen in accordance with the literature.22 The solution
level was maintained at a height of 5 mm and the solu-
tions were weekly renewed. After six months, the sam-
ples were slit lengthwise and the changes in their micro-
structures were studied with high-pressure mercury
intrusion porosimetry and scanning electron microscope.
Attention was paid mainly to the lower parts of the
samples, near the H2SO4 solution.
3 RESULTS AND DISCUSSION
3.1 Visual appearance
After being in 0.5 % H2SO4 for six months, a visible
degradation of the submerged part occurred on all the
investigated samples and the creation of a rust layer on
the periphery of each sample, just above the level of the
acid, took place. Figure 1 illustrates the effect of the
0.5 % sulfuric acid on the SRP-concrete sample. A de-
tailed view of the immersed part of the sample is shown
in Figure 2. There are well recognizable exposed aggre-
gates and white precipitates of gypsum (CaSO42H2O) on
the surface.
3.2 Changes in the porosity
It was found that all the concrete samples had almost
identical pore distributions before the sulfate attack. The
samples mainly contained pores with a diameter of
0.1 μm (Figure 3). The largest porosity was observed for
the FA concrete and the lowest for the SRS concrete. The
results of the determination of the cumulative pore vol-
umes of the samples exposed to the sulfate attack for six
months are also in Figure 3. The pore-size distribution
remained approximately the same, but the samples had a
slightly higher amount of the pores with a diameter
below 0.1 μm. This means that larger pores were filled in
to a certain degree by the products of the reactions of the
M. VY[VAØIL et al.: MICROSTRUCTURAL CHANGES OF FINE-GRAINED CONCRETE ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 883–888 885
Figure 3: Pore-size distribution of the investigated concrete samples
before and after a 6-month (6M) exposure to H2SO4 0.5 %
Slika 3: Razporeditev velikosti por v preiskovanih vzorcih betona
pred 6-mese~no (6M) izpostavitvijo H2SO4 0,5 % in po njej
Figure 2: Immersed part of SRP concrete sample after being in H2SO4
0.5 % for 6 months
Slika 2: Potopljeni del SRP-betonskega vzorca po namakanju 6 me-
secev v H2SO4 0,5 %
Figure 1: Lower part of SRP concrete sample after being in H2SO4
0.5 % for 6 months
Slika 1: Spodnji del SRP-betonskega vzorca po namakanju 6 mesecev
v H2SO4 0,5 %
sulfate attack. The total porosities of the investigated
concrete samples are depicted in Figure 4. The reduction
in the total porosity of most samples occurred after the
sulfate attack, which indicates that the pores were filled
in by the products of the sulfate corrosion of the con-
crete.
The most significant decrease was observed for the
PC and FA concretes. Conversely, for the SRS and GL
concretes, an increase in the total porosity was found. In
the case of the SRS concrete, it is apparently due to the
low porosity before the sulfate attack, prohibiting a
smooth crystallization of the sulfate-corrosion products
in the pores and a disruption of the microstructures of the
samples. The increase in the total porosity of the GL
concrete is in accordance with the conclusions of E. F.
M. VY[VAØIL et al.: MICROSTRUCTURAL CHANGES OF FINE-GRAINED CONCRETE ...
886 Materiali in tehnologije / Materials and technology 49 (2015) 6, 883–888
Figure 6: SEM image of gypsum identified in the structure of sub-
merged part of PC-concrete sample after a 6-month sulfate attack
Slika 6: SEM-posnetek sadre, dobljene v potopljenem delu PC-beton-
skega vzorca po 6-mese~ni izpostavitvi sulfatu
Figure 7: SEM image of microstructure of PC-concrete sample after a
6-month sulfate attack
Slika 7: SEM-posnetek mikrostrukture vzorca PC-betona po 6-me-
se~ni izpostavitvi sulfatu
Figure 5: Distribution of sulfur (red), aluminum (green) and iron (blue) on the surfaces of lower parts of lengthwise slices of the selected types of
concrete samples obtained with EDX SEM
Slika 5: EDX-SEM-razporeditev `vepla (rde~e), aluminija (zeleno) in `eleza (modro) na povr{ini spodnjega dela vzdol`nega prereza izbranih
vrst betona
Figure 4: Total porosity of the concrete samples before and after a
6-month exposure to 0.5 % H2SO4
Slika 4: Skupna poroznost betonskih vzorcev pred 6-mese~no izpo-
stavitvijo H2SO4 0,5 % in po njej
Irassar et al.12 He found that a concrete containing a
limestone filler is more susceptible to a sulfate attack and
less durable than the corresponding plain mortar, as
indicated by its larger expansion, greater surface
deterioration, deeper transition zone of the attack, larger
deposition of gypsum and higher degree of the CH
depletion.
3.3 Changes in the microstructure
By mapping the chemical-element distribution on the
surface of the lower part of the lengthwise slice of each
samples, the presence of the locations rich in sulfur, iron
and aluminum was determined. In the lowest parts of the
samples, sulfur prevailed in the gypsum form. Using
EDX REM, an aluminum-rich zone, identified as hydra-
ted Al(OH)3, was located just above the gypsum. The
presence of Al(OH)3 is presumably caused by the forma-
tion of ettringite from C4AF in the presence of gypsum
(Equation (2)). Thus, it can be concluded that even
during a six-month sulfate attack, ettringite forms from
C4AF. In the case of the samples made from the
materials with a high iron content (FA, PC, SRP), the
zone rich in iron was present between the sulfur and alu-
minum locations. The iron compound was not sufficient-
ly established; it is just known that it is a silicate without
aluminum. The identification of this compound is still in
process. Selected EDX REM images of the mapped che-
mical-element distributions are shown in Figure 5. The
PC-, SRP- and FA-concrete samples had very similar
distributions of chemical elements, and so did SRS and
GBFS. The MK sample did not have a narrow zone rich
in aluminum because of a high aluminum content in the
used metakaolin.
In the submerged parts of all the investigated sam-
ples, typical prismatic gypsum crystals were identified
with EDX SEM (Figure 6). The pores of the PC-con-
crete sample were largely filled with the crystalline
M. VY[VAØIL et al.: MICROSTRUCTURAL CHANGES OF FINE-GRAINED CONCRETE ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 883–888 887
Figure 11: SEM image of microstructure of FA-concrete sample after
a 6-month sulfate attack
Slika 11: SEM-posnetek mikrostrukture vzorca FA-betona po 6-me-
se~ni izpostavitvi sulfatu
Figure 9: SEM image of microstructure of GL-concrete sample after a
6-month sulfate attack
Slika 9: SEM-posnetek mikrostrukture vzorca GL-betona po 6-me-
se~ni izpostavitvi sulfatu
Figure 10: SEM image of microstructure of GBFS-concrete sample
after a 6-month sulfate attack – detail of a partially reacted slag grain
Slika 10: SEM-posnetek mikrostrukture vzorca GBFS-betona po
6-mese~ni izpostavitvi sulfatu – detajl delno zreagiranega zrna `lindre
Figure 8: SEM image of microstructure of MK-concrete sample after
a 6-month sulfate attack
Slika 8: SEM-posnetek mikrostrukture vzorca MK-betona po 6-me-
se~ni izpostavitvi sulfatu
products of the sulfate attack, particularly with the
typical needle crystals of ettringite (Figure 7), confirm-
ing the reduction in the total porosity. Of all the investi-
gated concrete samples, the SRP- and SRS-concrete
samples showed the smallest changes in the microstruc-
tures as just small amounts of ettringite were identified.
In the microstructures of the MK- and GL-concrete
samples, a lot of portlandite in typical hexagonal crystals
was identified on the surfaces of the pores (Figures 8
and 9), in contrast to the microstructures of these sam-
ples before the sulfate attack. An increased amount of
portlandite in the pores was observed in all the studied
samples, but for the MK and GL samples, this amount
was enormous. This increased presence of portlandite
was probably caused by the capillary action of the solu-
tion while the samples were kept in H2SO4. A direct
contact of the cement matrix with the solution of sulfuric
acid leads to a sulfate attack and the formation of gyp-
sum. The non-aggressive solution penetrates to the
higher parts of a sample, dissolves the surrounding
Ca(OH)2 and, during the drying of the sample, its
crystallization occurs in the pores. For this reason, in the
lowest parts of the samples, only gypsum as the primary
product of the sulfate attack was identified. In addition to
gypsum, the higher parts of the samples also contained
ettringite and other by-products of the sulfate attack. In
the case of the GBFS-concrete sample, crystals of port-
landite were concentrated in the crevices around the
partially reacted slag grains (Figure 10). Ettringite was
identified on the surfaces of the pores. The FA-concrete
sample contained a lot of damaged particles of fly ash
and needle crystals of ettringite (Figure 11).
4 CONCLUSIONS
Microstructural changes in various types of fine-
grained concrete after a six-month sulfate attack were
investigated with mercury intrusion porosimetry and
scanning electron microscopy. It was found that the total
porosity of most samples was decreased after the sulfate
attack, indicated by the products of the sulfate corrosion
filling in the pores of the concrete. The most significant
decrease was observed for the PC and FA concretes. The
smallest changes in the microstructure occurred in the
samples made from the sulfate-resisting cements. There-
by, the suitability of their use in the case of a sulfate
attack was confirmed. The formation of the locations
rich in sulfur, iron and aluminum during the sulfate
attack on the concrete was determined by mapping the
chemical-element distribution. Due to the presence of
Al(OH)3, which was identified in the aluminum-rich
location, it can be concluded that, even during a six-
month sulfate attack, ettringite is formed from C4AF. The
fate of the iron originating from C4AF is still not clear.
Acknowledgement
This outcome was achieved with the financial support
of the Czech Science Foundation (grant no. 13-22899P)
and the European Union "Operational Programme Re-
search and Development for Innovations", No. CZ.1.05/
2.1.00/03.0097 (AdMaS).
5 REFERENCES
1 A. Neville, Cem. Concr. Res., 24 (2004), 1275–1296, doi:10.1016/
j.cemconres.2004.04.004
2 H. Yuan, P. Dangla, P. Chatellier, T. Chaussadent, Cem. Concr. Res.,
53 (2013), 267–277, doi:10.1016/j.cemconres.2013.08.002
3 D. Stein, Instandhaltung von Kanalisationen, 3rd ed., Ernst, Berlin
1999, 141
4 J. Skalny, J. Marchand, I. Odler, Sulfate attack on concrete, 1st ed.,
Spon Press, London 2002
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of cement-based systems, 1st ed., E & FN Spon, London 1997,
177–184
6 P. J. Monteiro, K. E. Kurtis, Cem. Concr. Res., 33 (2003), 987–993,
doi:10.1016/S0008-8846(02)01097-9
7 R. P. Khatari, V. Sirivivatnanon, Cem. Concr. Com., 27 (1997),
1179–1189, doi:10.1016/S0008-8846(97)00119-1
8 H. F. W. Taylor, Cement Chemistry, 1st ed., Thomas Telford, London
1997, doi:10.1680/cc25929
9 B. Lothenbach, E. Wieland, Waste Management, 26 (2006), 706–719
doi:10.1016/j.wasman.2006.01.023
10 R. Gollop, H. F. W. Taylor, Cem. Concr. Res., 26 (1996), 1013–1028,
doi:10.1016/0008-8846(96)00089-0
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doi:10.1016/0008-8846(92)90033-R
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(2003), 31–41, doi:10.1016/S0008-8846(02)00914-6
13 J. Marchand, E. Samson, Y. Maltais, J. J. Beaudoin, Cem. Concr.
Com., 24 (2002), 317–329, doi:10.1016/S0958-9465(01)00083-X
14 M. Santhanam, M. D. Cohen, J. Olek, Cem. Concr. Res., 31 (2001),
845–851, doi:10.1016/S0008-8846(01)00510-5
15 S. Diamond, Cem. Concr. Com., 25 (2003), 1161–1164, doi:10.1016/
S0958-9465(03)00138-0
16 M. Collepardi, Cem. Concr. Com., 21 (1999), 147–154, doi:10.1016/
S0958-9465(98)00044-4
17 D. Macphee, S. J. Barnett, Cem. Concr. Res., 34 (2004), 1591–1598,
doi:10.1016/j.cemconres.2004.02.022
18 N. J. Crammond, Cem. Concr. Res., 15 (1985), 1039–1050,
doi:10.1016/0008-8846(85)90095-X
19 G. R. Gouda, D. M. Roy, A. Sarkar, Cem. Concr. Res., 5 (1975),
519–522, doi:10.1016/0008-8846(75)90026-5
20 M. Romer, Cem. Concr. Com., 25 (2003), 1173–1176, doi:10.1016/
S0958-9465(03)00155-0
21 M. Romer, L. Holzer, M. Pfiffner, Cem. Concr. Com., 25 (2003),
1111–1117, doi:10.1016/S0958-9465(03)00141-0
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Gemertb, W. Verstraetec, Cem. Concr. Res., 34 (2004), 2223–2236,
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M. VY[VAØIL et al.: MICROSTRUCTURAL CHANGES OF FINE-GRAINED CONCRETE ...
888 Materiali in tehnologije / Materials and technology 49 (2015) 6, 883–888
A. GRAJCAR et al.: EFFECT OF THERMOMECHANICAL TREATMENT ON THE CORROSION BEHAVIOUR ...
889–894
EFFECT OF THERMOMECHANICAL TREATMENT ON THE
CORROSION BEHAVIOUR OF Si- AND Al-CONTAINING
HIGH-Mn AUSTENITIC STEEL WITH Nb AND Ti
MICRO-ADDITIONS
VPLIV TERMOMEHANSKE OBDELAVE NA KOROZIJSKO
VEDENJE MANGANSKEGA AVSTENITNEGA JEKLA Z
VSEBNOSTJO Si IN Al, MIKROLEGIRANEGA Z Nb IN Ti
Adam Grajcar, Aleksandra Plachciñska, Santina Topolska, Monika Kciuk
Silesian University of Technology, Institute of Engineering Materials and Biomaterials, Konarskiego Street 18a, 44-100 Gliwice, Poland
adam.grajcar@polsl.pl
Prejem rokopisa – received: 2014-07-31; sprejem za objavo – accepted for publication: 2014-12-19
doi:10.17222/mit.2014.148
The corrosion behavior of the 27Mn-4Si-2Al type austenitic steel micro-alloyed with Nb and Ti was evaluated in acidic 0.1 M
H2SO4 and chloride 3.5 % NaCl environments using potentiodynamic polarization tests. The corrosion properties of
solution-treated specimens were compared to thermomechanically processed specimens. In the acidic solution, the steel
exhibited a lower corrosion resistance than in the chloride solution, independently of the heat treatment applied. SEM and light
micrographs confirmed that the corrosion attack in the acidic solution was higher when compared to the chloride solution. The
steel showed evidence of pitting and uniform corrosion in both the acidic and chloride solutions. The corrosion resistance of
supersaturated specimens in both 0.1 M H2SO4 and 3.5 % NaCl media was lower when compared to the thermomechanically
treated specimens. It was found that the corrosion behavior of the examined high-Mn steel depends on the passivation tendency
of the alloying elements (Mn, Al) and the grain size.
Keywords: high-Mn steel, austenitic steel, corrosion resistance, passivity, thermomechanical treatment, potentiodynamic
polarization test
Korozijsko vedenje avstenitnega jekla 27Mn-4Si-2Al, mikrolegiranega z Nb in Ti je bilo ocenjeno v kislem 0,1 M H2SO4 in v
kloridnem okolju 3,5 % NaCl, s potenciodinami~nim polarizacijskim preizkusom. Korozijske lastnosti raztopno `arjenih
vzorcev so bile primerjane s termomehansko izdelanimi vzorci. V kisli raztopini so bili vzorci manj korozijsko odporni kot v
kloridni raztopini, neodvisno od uporabljene toplotne obdelave. SEM in svetlobni posnetki so potrdili, da je bil napad korozije v
primerjavi s kloridno raztopino izrazitej{i v kisli raztopini. Na jeklu so bili dokazi za jami~asto in splo{no korozijo v obeh
raztopinah, kisli in bazi~ni. Korozijska odpornost prenasi~enih vzorcev v 0,1 M H2SO4 in v 3,5 % NaCl je bila manj{a v
primerjavi s termomehansko obdelanimi vzorci. Ugotovljeno je, da je korozijsko vedenje preiskovanega visokomanganskega
jekla odvisno od nagnjenosti k pasivaciji legirnih elementov (Mn, Al) in od velikosti zrn.
Klju~ne besede: visokomangansko jeklo, avstenitno jeklo, korozijska odpornost, pasivnost, termomehanska obdelava, preizkus
potenciodinami~ne polarizacije
1 INTRODUCTION
High-manganese austenitic steels are being deve-
loped as advanced automotive structural materials due to
their superior combination of strength, ductility, and
crashworthiness. However, their application for auto-
motive parts is limited because of their low corrosion
resistance. The corrosion behavior of high-manganese
austenitic steels depends on their chemical composition
and the corrosion medium applied. It was found that
additions of Al and Cr improve the corrosion resistance
of high-Mn steels.1–3 This is due to the passive-film-
forming tendency of the steel surface. It is reported that a
silicon addition to steel decreases its corrosion
resistance.4 The tendency to create a passive protective
layer on the steel surface depends on the type of corro-
sive environment. Kannan et al.1 reported that
29Mn-3.1Al-1.4Si austenitic steel showed a lower corro-
sion resistance in an acidic solution than in a chloride
solution.
The corrosion behavior of high-Mn austenitic steels
also depends on the heat treatment applied and plastic
deformation. Generally, cold working increases the
corrosion rate because of the deformation twins, which
represent the regions of different potentials from the
matrix.5,6 The corrosion resistance of steel is also related
to the dislocation density. In cold-rolled steels the dis-
location amount is much higher when compared to
hot-rolled steels because of the recrystallization effects
reducing the dislocation density.7–9 The corrosion resis-
tance of grain boundaries is poor because of a high dislo-
cation density present at these regions. Di Schino et al.10
reported that the corrosion resistance of steel is also
related to its grain size.
During the past decade, there were a number of
reports on the corrosion properties of Fe-(high-Mn)-
Materiali in tehnologije / Materials and technology 49 (2015) 6, 889–894 889
UDK 669.14:620.193 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)889(2015)
Al-Si alloys. Most investigations focused on the effect of
the alloying elements on the corrosion behaviour. There
are few reports about the effects of heat treatment or
plastic deformation on the corrosion resistance of high-
Mn austenitic steels.5,11 Therefore, the corrosion proper-
ties of solution-treated and thermomechanically pro-
cessed specimens were compared in this work.
2 EXPERIMENTAL PROCEDURE
The work addresses the corrosion behaviour of the
vacuum-melted high-manganese steel containing 0.04 %
C, 27.5 % Mn, 4.18 % Si, 1.96 % Al, 0.017 % S,
0.004 % P and 0.0028 % N. Carbon and manganese are
the main austenite stabilizers. Silicon and aluminium
should provide solid-solution strengthening. Addition-
ally, micro-additions of Nb (0.033 %) and Ti (0.01 %)
are added for precipitation strengthening and grain
refinement. The ingots were hot forged and roughly
rolled to a thickness of 4.5 mm. Two types of specimens
were prepared using the thermomechanical rolling and
the subsequent solution treatment. The thermomecha-
nical processing consisted of hot rolling of flat samples
in 3 passes to a final thickness of about 2 mm obtained at
the finishing rolling temperature of 850 °C. Then, the
samples were rapidly cooled in water to room tempe-
rature. The second group of samples was subsequently
annealed at 900 °C for 20 min.
The corrosion properties of the 27Mn-4Si-2Al steel
were evaluated using potentiodynamic polarization tests
in two environments: acidic 0.1 M H2SO4 and 3.5 %
NaCl solutions. The corrosion behavior of the solution-
treated specimens was compared to the thermomecha-
nically processed specimens. The corrosion tests were
examined using the average of the measurement results
for four supersaturated and two thermomechanically
treated specimens. The exposed specimen surfaces
(0.38 cm2) were ground with SiC paper of up to 800 grit.
The samples were washed in distilled water and rinsed in
acetone before the experiments. All the corrosion tests
were conducted using freshly prepared electrolytes.
Polarization studies were carried out using an Atlas 0531
electrochemical unit potentiostat/galvanostat driven by
the AtlasCorr05 software and an electrochemical corro-
sion cell consisting of a specimen as the working elec-
trode, stainless steel as the counter electrode and a
silver/silver chloride (Ag/AgCl) reference electrode. A
scan rate of 0.5 mV/s was employed during polarization.
Potentiodynamic-scan data were collected to determine
the electrochemical parameters – corrosion potential Ecorr
and corrosion current density icorr – using Tafel slope
extrapolation.
After the polarization tests, the samples were ob-
served using a Zeiss SUPRA 25 scanning electron
microscope (SEM) to assess the type of the corrosion
attack. Subsequently, the depth of the corrosion pits on
the cross-sectioned specimens using a Zeiss Axio Ob-
server Z1m light microscope was evaluated. For light
microscopy, the samples were mechanically ground with
SiC paper of up to 1500 grit, polished with Al2O3 with a
granularity of 0.1 μm and then etched using 5 % nital to
reveal the microstructure.
3 RESULTS AND DISCUSSION
A typical micrograph of a thermomechanically
treated specimen is shown in Figure 1a. The light
micrograph presents relatively coarse austenite grains
elongated in the direction of hot rolling. The average
grain size was estimated to be about 70 μm. Annealing
twins and elongated sulfide inclusions were also ob-
served. Figure 1b shows the light micrograph of a
solution-treated specimen. The microstructure of the
steel solution-treated from a temperature of 900 °C is
characterized by a bimodal distribution of the grains. The
average grain size was estimated to be about 40 μm.
Relatively large elongated austenite grains and small
recrystallized grains are apparent, indicating a strain
accumulation that remains after the thermomechanical
rolling. As a result, the driving force decreasing the
accumulated energy leads to a partial recrystallization of
A. GRAJCAR et al.: EFFECT OF THERMOMECHANICAL TREATMENT ON THE CORROSION BEHAVIOUR ...
890 Materiali in tehnologije / Materials and technology 49 (2015) 6, 889–894
Figure 1: Austenitic microstructures of: a) thermomechanically
treated and b) supersaturated steels
Slika 1: Avstenitna mikrostruktura: a) termomehansko obdelanega in
b) prenasi~enega jekla
the austenite grains during the subsequent annealing of
the samples at 900 °C.
Based on the potentiodynamic curves (Figures 2 to
5), the corrosion potential Ecorr and corrosion current
density icorr were determined. The average calculated
values are shown in Table 1. The results of the corrosion
tests are characterized by a small scatter. The lowest
corrosion resistance was obtained in the acidic solution,
independently of the heat-treatment type (thermomecha-
nically treated or supersaturated specimens). The highest
values of the corrosion current density were registered in
the acidic solution. Kannan et al.1 also reported that the
29Mn-3.1Al-1.4Si austenitic steel showed a lower corro-
sion resistance in an acidic medium than in a chloride
solution.
Table 1: Electrochemical-polarization data of thermomechanically
treated (t) and supersaturated (s) steels in two different environments
Tabela 1: Podatki o elektrokemijski polarizaciji termomehansko
obdelanega (t) in prenasi~enega (s) jekla v dveh razli~nih okoljih
Type of
heat
treatment
3.5 % NaCl 0.1 M H2SO4
Ecorr/
mV
icorr/
mA/cm2
Ecorr/
mV
icorr/
mA/cm2
average
value
s
–774.4 0.30 –584.7 12.3
standard
deviation 29.1 0.05 8.8 1.5
average
value
t
–787.6 0.09 –583.9 1.4
standard
deviation 6.5 0.007 6.2 0.2
The solution-treated specimens showed a higher
corrosion-current-density values (12.29 mA/cm2) than
the thermomechanically treated specimens (1.44 mA/cm2).
In the chloride solution, the supersaturated specimens
also showed a higher corrosion current density (0.24
mA/cm2) when compared to the thermomechanically
treated specimens (0.09 mA/cm2). Generally, it is re-
ported that cold working decreases the corrosion
resistance of steel.5,6 Interestingly, the thermomecha-
nically treated specimens showed a higher corrosion
resistance than the supersaturated specimens independ-
ently of the corrosion medium. It should be noted that
the examined steel was hot rolled. Hence, it has a much
lower dislocation density than the cold-rolled specimens.
Therefore, the higher corrosion resistance of the thermo-
mechanically treated specimens is related to other fac-
tors.
The corrosion-potential values of the thermomecha-
nically treated and supersaturated specimens are similar
in both the acidic (Figure 2) and chloride solutions
(Figure 3). The lowest corrosion-potential values were
A. GRAJCAR et al.: EFFECT OF THERMOMECHANICAL TREATMENT ON THE CORROSION BEHAVIOUR ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 889–894 891
Figure 4: Potentiodynamic-polarization curves of the thermomecha-
nically treated (t) and supersaturated (s) steels obtained in 3.5 % NaCl
solution
Slika 4: Potenciodinami~ni polarizacijski krivulji termomehansko
obdelanega (t) in prenasi~enega (s) jekla, posneti v raztopini 3,5 %
NaCl
Figure 2: Potentiodynamic-polarization curves of the supersaturated
steel obtained in 0.1 M H2SO4 solution (4 samples tested)
Slika 2: Potenciodinami~ne polarizacijske krivulje prenasi~enega
jekla, posnete v raztopini 0,1 M H2SO4 (preizku{eni so bili 4 vzorci)
Figure 3: Potentiodynamic-polarization curves of the thermomecha-
nically treated steel obtained in 3.5 % NaCl solution (2 samples
tested)
Slika 3: Potenciodinami~ni polarizacijski krivulji termomehansko
obdelanega jekla, posneti v raztopini 3,5 % NaCl (preizku{ena sta bila
2 vzorca)
registered in the chloride solution (Figure 4). They are
related to the earlier appearance of the corrosion pits in
the chloride environment. In the 0.1 M H2SO4 medium,
higher (right-shifted) corrosion-potential values were
registered (Figure 5). The lowest values of the corrosion
potential in the chloride solution are also reported for the
other high-Mn alloys.1 However, it should be noted that
the values of the corrosion current density were much
higher in the acidic solution (Table 1). They are related
to the stronger corrosion attack in this medium.
The SEM and light micrographs of the thermome-
chanically treated and supersaturated specimens after the
polarization tests in two environments are shown in
Figures 6 to 8. In the acidic solution, both steels showed
extensive uniform corrosion. In addition to uniform
corrosion, pitting corrosion can also be observed (Figure
6). In the thermomechanically treated specimens, after
the corrosion tests in 0.1 M H2SO4 the maximum depth
of corrosion pits was evaluated as 55 μm (Figure 6b). In
the supersaturated specimens, the maximum depth of
corrosion pits was slightly higher (Figure 7b). On both
types of specimens, wide and shallow corrosion pits
were observed. It was noted that corrosion pits were
usually formed in the coarse-grained regions of the
microstructure. Fine-grained regions were dominated by
uniform corrosion and a smaller density of the corrosion
pits. A similar relationship between the corrosion resis-
A. GRAJCAR et al.: EFFECT OF THERMOMECHANICAL TREATMENT ON THE CORROSION BEHAVIOUR ...
892 Materiali in tehnologije / Materials and technology 49 (2015) 6, 889–894
Figure 6: a) SEM micrograph of the surface and b) light micrograph
of the cross-section of thermomechanically treated steel specimens,
potentiodynamically polarized in acidic solution
Slika 6: a) SEM-posnetek povr{ine in b) svetlobni posnetek pre~nega
prereza termomehansko obdelanega vzorca jekla, potenciodinami~no
polariziranega v kisli raztopini
Figure 7: a) SEM micrograph of the surface and b) light micrograph
of the cross-section of supersaturated steel specimens, potentiodyna-
mically polarized in acidic solution
Slika 7: a) SEM-posnetek povr{ine in b) svetlobni posnetek pre~nega
prereza prenasi~enega vzorca jekla, potenciodinami~no polariziranega
v kisli raztopini
Figure 5: Potentiodynamic-polarization curves of the thermomecha-
nically treated (t) and supersaturated (s) steels obtained in 0.1 M
H2SO4 solution
Slika 5: Potenciodinami~ni polarizacijski krivulji termomehansko
obdelanega (t) in prenasi~enega (s) jekla, posneti v raztopini 0,1 M
H2SO4
tance and grain size was observed earlier by Di Schino et
al.10 They observed that the pitting-corrosion rate
decreased with the decreasing grain size, while the uni-
form-corrosion resistance was impaired by grain
refining.
In the chloride solution, the thermomechanically
treated specimens showed uniform corrosion. In addition
to uniform corrosion, pitting-corrosion evidence was also
observed (Figure 8a). Other authors1,12,13 also observed
corrosion pits in different high-manganese steels after
the polarization tests in a chloride solution. After the
corrosion tests in 3.5 % NaCl, the maximum depth of the
corrosion pits of the thermomechanically treated speci-
mens was evaluated as 44 μm (Figure 8b). Similar
corrosion pits can also be identified for the supersatu-
rated specimens. Small pits are usually formed at non-
metallic inclusions (Figure 9). The same nature of corro-
sion damages of high-Mn steels was also observed by
Grajcar et al.6 on the samples subjected to immersion
tests.
It was observed that pitting corrosion was usually
initiated at sulfide inclusions. The negative impact of
MnS inclusions on the corrosion resistance of austenitic
steels was also observed by Donik et al.14 The density of
corrosion pits and corrosion products are greater in the
acidic solution than in the chloride medium. Microscopic
observations are confirmed by the potentiodynamic test
results, which showed that the investigated steel has a
better corrosion resistance in the chloride medium when
compared to the acidic environment.
Corrosion resistance depends on the pH of corrosion
solutions and the values of the corrosion potential (Pour-
baix diagrams).15 In the 0.1 M H2SO4 (pH = 1) and 3.5 %
NaCl solutions, manganese dissolves as Mn2+. It can be
concluded that Mn showed a non-passivating tendency in
these environments. In the acidic solution, aluminium
dissolves as Al3+ ions, preventing the formation of the
passive layer. In the 3.5 % NaCl solution, aluminium
forms the oxide passive layer, slightly increasing the
corrosion resistance of the investigated steel.
4 CONCLUSIONS
Potentiodynamic polarization tests showed that the
27Mn-4Si-2Al type austenitic steel is characterized by
the lowest corrosion resistance in the 0.1 M H2SO4 solu-
tion, independently of the heat treatment applied (ther-
momechanically treated or solution-treated specimens).
A. GRAJCAR et al.: EFFECT OF THERMOMECHANICAL TREATMENT ON THE CORROSION BEHAVIOUR ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 889–894 893
Figure 9: a) SEM micrograph of the surface and b) light micrograph
of the cross-section of supersaturated steel specimens, potentiodyna-
mically polarized in chloride solution
Slika 9: a) SEM-posnetek povr{ine in b) svetlobni posnetek pre~nega
prereza prenasi~enega vzorca jekla, potenciodinami~no polariziranega
v raztopini klorida
Figure 8: a) SEM micrograph of the surface and b) light micrograph
of the cross-section of thermomechanically treated steel specimens,
potentiodynamically polarized in chloride solution
Slika 8: a) SEM-posnetek povr{ine in b) svetlobni posnetek pre~nega
prereza termomehansko obdelanega vzorca jekla, potenciodinami~no
polariziranega v raztopini klorida
Microscopic observations confirmed that corrosion
damages are more numerous in the acidic solution. In
both the acidic and chloride solutions wide and shallow
corrosion pits of various depths were identified. In
addition to pitting corrosion, extensive uniform corrosion
was observed. The high density of the corrosion damages
in the acidic solution is related to the non-passivating
tendency of Mn and Al in this environment. Independ-
ently of the corrosion media, the thermomechanically
treated specimens showed a better corrosion resistance
than the supersaturated specimens. This means that the
bimodal distribution of the grain size has a greater effect
on decreasing the corrosion resistance of the investigated
steel than the effect of hot working.
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SURFACE FREE ENERGY OF HYDROPHOBIC COATINGS OF
HYBRID-FIBER-REINFORCED HIGH-PERFORMANCE
CONCRETE
PROSTA ENERGIJA POVR[INE HIDROFOBNIH PREMAZOV NA
VISOKOZMOGLJIVEM BETONU, OJA^ANEM S HIBRIDNIMI
VLAKNI
Danuta Barnat-Hunek, Piotr Smarzewski
Lublin University of Technology, Faculty of Civil Engineering and Architecture, Nadbystrzycka Str. 40, 20-618 Lublin, Poland
d.barnat-hunek@pollub.pl
Prejem rokopisa – received: 2014-08-03; sprejem za objavo – accepted for publication: 2014-12-12
doi:10.17222/mit.2014.174
The aim of the research presented in the paper was to evaluate the feasibility of using hydrophobic preparations based on
organosilicon compounds for the protection treatment of hybrid-fiber-reinforced high-performance concrete (FRHPC) surfaces.
The wettability of concrete has a direct effect on the durability and corrosion resistance. The wetting properties of FRHPC were
evaluated through the measurement of the contact angle between the surfaces of these materials with either water or glycerine
used as probe liquids. On this basis, the surface free energy (SFE) was determined. The polar and disperse components of SFE
were obtained by means of the Owens-Wendt method. Three different siloxane preparations were deposited onto seven types of
concrete with the fiber content ranging from 0 % to 1 %. In order to investigate the effect on the strength, the granodiorite
aggregate in concrete mixes 5, 6, 7 was replaced with granite. The basic characteristics of the concrete strength were examined:
the tensile splitting strength, the compressive strength, the modulus of elasticity. A SEM examination of the coated concrete
surfaces confirmed that preparations A–C can effectively cover the voids and pores present in the concrete surfaces. The
presented results indicate that the surfaces of the concrete with a silane film had a wide range of SFE, depending on the kind of
agent. SFE depended on the chemical reactivity of the silanes used, the type of solvent, the viscosity and surface tension of the
solution. The evaluation of the contact angle and SFE helped to efficiently select the most appropriate preparation.
Keywords: surface free energy (SFE), contact angle, hydrophobization, high-performance concrete, hybrid fiber, Owens-Wendt
method
Namen raziskave, predstavljene v ~lanku, je bil oceniti izvedljivost uporabe hidrofobne obdelave na osnovi organosilikonskih
spojin za za{~ito povr{ine zmogljivega betona (FRHPC), oja~anega s hibridnimi vlakni. Omo~ljivost betona ima neposreden
vpliv na zdr`ljivost in odpornost proti koroziji. Omo~ljivost FRHPC je bila ocenjena z merjenjem sti~nega kota med povr{ino
teh materialov pri uporabi vode in glicerina kot preizkusne teko~ine. Na tej osnovi je bila dolo~ena prosta energija povr{ine
(SFE). Polarne in razpr{ilne komponente SFE so bile dolo~ene z Owen-Wendtovo metodo. Trije razli~ni pripravki siloksana so
bili naneseni na sedem vrst betona z vsebnostjo vlaken od 0 % do 1 %. Za preiskavo vpliva na trdnost je bil zrnati granodiorit
nadome{~en v betonskih me{anicah 5, 6 in 7 z granitom. Preiskovane so bile osnovne zna~ilnosti trdnosti betona: natezna
cepilna trdnost, tla~na trdnost in modul elasti~nosti. SEM-preiskave pokrite povr{ine betona so potrdile, da priprave A–C lahko
u~inkovito prekrijejo praznine in pore, ki so na povr{ini betona. Prikazani rezultati ka`ejo, da ima povr{ina betona s plastjo
silanov velik razpon SFE, odvisno od vrste predstavnika. SFE je odvisna od kemijske reaktivnosti uporabljenega silana, vrste
topila, viskoznosti in povr{inske napetosti raztopine. Ocena kota stika in SFE je pomagala pri izbiri najbolj primernega
preparata.
Klju~ne besede: prosta energija povr{ine (SFE), kot stika, hidrofobizacija, visokozmogljiv beton, hibridna vlakna, Owens-
Wendt-ova metoda
1 INTRODUCTION
High-performance fiber-reinforced concrete is a ce-
mentitious material with a low water/cement ratio and a
high cement content.1,2 Short, straight steel fibers are
usually used to enhance the tensile strength and increase
the ductility. As a result, the mechanical properties of
hybrid-fiber-reinforced high-performance concrete are
considerably enhanced compared to the normal con-
crete.2,3 High-performance concretes are often exposed to
aggressive impacts of the environment and, therefore,
they must have a high resistance to chemical corrosion,
frost corrosion, weathering, impact of aggressive water
and many other corrosive agents.
Moisture damage is a major factor in the deteriora-
tion of building materials. One of the methods used to
protect the concrete surface is hydrophobization.4,5 It
causes a decrease in the capillary water absorption, thus
allowing free vapor permeability. Organosilicon com-
pounds – siloxanes or methyl silicone resins6–8 – are
mostly used as concrete hydrophobing agents. Research9
confirmed that a polyethylhydrosiloxane admixture has a
beneficial effect on the durability of the concrete with
large volumes of supplementary cementitious materials.
Concrete is normally a hydrophilic material, which sig-
nificantly reduces the durability of concrete structures. In
the research to synthesize water-repellent concrete, the
emulsion was enriched with the polymethyl-hydrogen
Materiali in tehnologije / Materials and technology 49 (2015) 6, 895–902 895
UDK 691:620.1:691.3:620.193 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)895(2015)
siloxane-oil hydrophobic agent.10 There are also tests
aimed at increasing the strength and toughness using
special techniques like polymer impregnation of the
matrices of fibre-reinforced concretes.11
Sustainability is the necessity for concrete and nano-
technology is the chance for the future of concrete-poly-
mer composites.12 In the case of impregnation, polymer-
cement concrete (PCC) gained popularity. The most
often used polymer-cement concretes are modified using
styrene-butadiene co-polymer, acrylic polymers and
epoxy resin.13
These coatings make the concrete surface non-wett-
able by water and corrosive compounds such as water-
soluble salts. The fact that building materials can be
wetted by liquids is of special importance, for example,
during their hydrophobization, impregnation and in the
production of anti-graffiti agents. The wettability of
concrete by means of liquids which contain corrosive
components is of great importance in practice; it may
indicate adhesive properties of concrete as well as pro-
tective coatings applied to its surface. The wettability of
concretes has a direct effect on the durability and
corrosion resistance. In the research of concretes, their
wettability and surface free energy (SFE) are considered
to be important elements in assessing the adhesion pro-
perties. They are particularly useful in the analysis of the
effects of modifying the surfaces of high-strength
concretes by means of various protective coatings.
According to the literature data,14–16 the contact angle
of the materials is an indicator of their wettability
properties. High wettability – hydrophilicity – occurs at
a low contact angle < 90 °, and insufficient wettability –
hydrophobicity – occurs at a high contact angle > 90 °.
The contact angle can be used to determine the surface
tension12 and define the surface free energy17–19 and
adhesion operation.14,20,21
The contact angle is influenced by many factors
which include, but are not limited to, the following:
surface physical and chemical homogeneity, surface
roughness and impurities, type of the measured liquid,
drop size of the measured liquid, humidity or ambient
temperature.18–21 Among the most commonly used me-
thods for determining the contact angle, one can mention
the following: the air-bubble method, the geometric me-
thod, the liquid-capillary-rise methods tested on sample
materials (among others, the Wilhelmy method) and the
direct-measurement method.22–25 A very popular method
to measure the contact angle is direct measurement using
a contact-angle analyzer or a goniometer.26,27
The surface free energy (the surface tension) is the
key parameter while evaluating the physicochemical cha-
racteristics of solid surfaces. The surface may be of a
dispersive nature (a dispersion component) or polar (a
polar component). Knowing the properties of impregna-
ting agents one can decrease or increase the SFE and,
therefore, the surface tension of materials, causing their
non-wettability, which is related, among others, to the
chemical corrosion and frost resistance. The highest
decrease in SFE can be due to the coatings that hydro-
phobize a surface to the largest extent.28
The surface free energy (SFE) is one of the thermo-
dynamic quantities describing the state of equilibrium of
the atoms in the surface layers of materials.14,29 SFE
represents the state of imbalance of the intermolecular
interactions present at the phase boundary of two diffe-
rent mediums. There are numerous methods for direct
determination of the surface free energy of liquids.
Owing to the fact that there are no direct methods for
determining the SFE of solids, some indirect methods
are used, which include, among others, the contact-angle
measurement method, calculating the surface free energy
on the basis thereon.21,22,30
The main methods for determining SFE were for-
mulated by Neumann, Wu, Owens and Wendt, Zisman
and Fox, Fowkes, Van-Oss-Chaudhury-Good.21
The Owens-Wendt method is commonly used for
determining the surface free energy of materials.19 This
method consists of determining the dispersion and polar
components of SFE. The polar component (sp), which is
the measure of the surface polarity, is associated with,
among others, the bond strength between the materials.
The analysis of the nature of a hydrophobized-
concrete surface layer in terms of wettability presented
in the article allowed us to assess, inter alia, the material
behavior in the presence of water and corrosive com-
pounds. In the cases where a significant resistance of a
concrete surface layer to the impact of a corrosive envi-
ronment is required, it is desirable to use preparations of
the lowest SFE value.
2 MATERIALS AND METHODS
2.1 Concrete mixtures
In the laboratory, seven concrete mixtures were pre-
pared using Portland cement 670.5 kg/m3 (CEM I 52.5
N-HSR/NA), aggregate 990 kg/m3, sand 500 kg/m3,
water 178 L/m3, microsilica 74.5 kg/m3, superplasticizer
20 L/m3 and steel and polypropylene fibers in varied
amounts.
Table 1: Fractions of fibers in various concretes
Tabela 1: Dele` vlaken v razli~nih betonih
Concrete type
Fraction, %
Steel fibers Polypropylenefibers
granodiorite aggregate
HPC1 – –
SFHPC 1 % –
HFHPC1 0.75 % 0.25 %
granite aggregate
HFHPC2 0.5 % 0.5 %
HFHPC3 0.25 % 0.75 %
PFHPC – 1 %
HPC2 – –
D. BARNAT-HUNEK, P. SMARZEWSKI: SURFACE FREE ENERGY OF HYDROPHOBIC COATINGS ...
896 Materiali in tehnologije / Materials and technology 49 (2015) 6, 895–902
Table 1 lists abbreviated names of the concretes and
amounts of steel and polypropylene fibers for various
batches. Granodiorite aggregate was used for the first
three concretes and granite aggregate was used for the
remaining four concretes.
The concrete samples were made on the basis of a
recipe determined experimentally using the known mor-
tar according to EN 206:2014-04. The mixing procedure
was as follows: quartz sand and coarse aggregate were
homogenized together and mixed with half quantity of
water. Then, cement, silica fume and the remaining water
were added and, finally, superplasticizer was added.
After the components were thoroughly mixed, fibers
were gradually added by hand to obtain homogeneous
and workable mixtures. The fibers were dosed gradually
so as not to be tried and not to sink to the bottom of the
mixture. The samples were formed directly after the
concrete compounds were mixed according to EN
12390-2:2011.
Molds coated with an anti-adhesive substance were
filled with concrete batches and compacted on a vibra-
ting table. All the samples were stored at a temperature
of about 23 °C until removing them from the moulds
after 24 h and they were then placed in a water tank for 7
d to cure. After 7 d the samples were removed from the
tank to cure in laboratory conditions for up to 28 d.
2.2 Properties of the concrete
2.2.1 Compressive strength and splitting tensile strength
Cubic concrete samples with dimensions of 100 mm
× 100 mm × 100 mm were applied. Research was con-
ducted according to EN 12390-3:2002 regarding the
compressive strength and EN 12390-6:2001 regarding
the splitting tensile strength. The evaluation of the grades
of the concretes was carried out using a Walter-Baj AG
compression tester within 3 MN after 28 d of maturation,
when the average compressive strength of the samples
was obtained.
2.2.2 Modulus of elasticity
The determination of the modulus of elasticity was
carried out on cylinders with a diameter of 150 mm and a
height of 300 mm by measuring the deformation of the
samples in a stress range from 0.5 MPa to 30 % of the
concrete compressive strength. The examination was
conducted by means of a Walter-Baj AG press and a mo-
dulus-measuring device with an extensometer.
The strength properties of the concretes adopted for
the examination are shown in Table 2.
2.3 Hydrophobic materials used
Three hydrophobic preparations commonly used as
construction chemicals were selected for the laboratory
tests; they differed in the type of solvent, viscosity and
concentration:
A – the water-based solution of methylosilicone resin in
the potassium hydroxide (1 : 6),
B – the organic-solvent-based alkyl alkoxysilane oligo-
mer,
C – the organic-solvent-based methylosilicone resin.
Each preparation was applied with a brush in two
layers. The first preparation was diluted in the proportion
of 1 : 6 according to the manufacturer’s requirements.
The other hydrophobic preparations with organic sol-
vents were not diluted. Thereafter, all the hydrophobized
bricks were seasoned for a period of 7 d in the laboratory
conditions to control the process of hydrolytic polycon-
densation of the hydrophobic coatings.
Producers do not provide full product characteristics
of the examined hydrophobic preparations or how their
solutions are applied in practice. For that purpose the
following parameters were determined experimentally:
the viscosity factor and the surface tension for all the
examined preparations. The major characteristics of the
applied preparations used in the research are listed in the
Table 3. Viscosity factor  was determined by measuring
the time of the solution flow in the Ostwald viscosimeter.
The surface tension was measured by raising the fluid in
the capillary. The research was executed at a room tem-
perature of 22.5 °C.
Table 3: Basic characteristics of hydrophobic preparations
Tabela 3: Osnovne zna~ilnosti hidrofobnih preparatov
Type of
formu-
lation
Viscosity
/
(Pa s × 10–3)
Density at
20 °C
(g/cm3)
Surface
tension
 (N/
(m × 10–3)
Quotient of
surface
tension and
viscosity /
A 1.099 1.26 67.92 61.73
B 1.479 0.80 23.11 15.65
C 2.846 0.82 24.30 8.54
2.4 Determination of the contact angle and surface
free energy
In order to calculate the surface free energy, the
contact-angle measurements of the analyzed concretes
were conducted. The method of directly measuring the
angle formed by a drop of the measuring liquid with the
surface measured was used, using a computer program
for the image analysis. The measurement of the contact
D. BARNAT-HUNEK, P. SMARZEWSKI: SURFACE FREE ENERGY OF HYDROPHOBIC COATINGS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 895–902 897
Table 2: Mechanical properties of concretes
Tabela 2: Mehanske lastnosti betonov
HPC1 SFHPC HFHPC1 HFHPC2 HFHPC3 PFHPC HPC2
Compressive strength (MPa) 151.0 154.9 144.7 133.9 122.3 94.6 129.5
Splitting tensile strength (MPa) 8.9 13.8 13.5 10.0 9.3 7.6 6.8
Modulus of elasticity (GPa) 38.37 39.74 34.27 32.45 29.60 29.42 32.55
angles of the measuring-liquid drops was carried out on a
research stand consisting of a goniometer integrated with
a camera for taking photos of the drops put onto the sur-
faces of the samples. The stand was described in31.
In order to examine the contact angle two measuring
liquids were used – distilled water and glycerine required
by the Owens-Wendt model used in the analyses.
Measuring-liquid drops of 2 mm3 were deposited by
means of a micropipette.14,32 Due to the heterogeneity of
the material, six drops were put on each sample. The
measurements were carried out twice: at the time of the
application of the drops, i.e., at 0 min and at 40 min.
Standard and hydrophobized surfaces HFHPC1 during
the examination of the contact angle of a glycerine drop
are shown in Figure 1.
For the calculation of the wettability of the concrete
surface, the SFE values of the measuring liquids (L),
and their dispersion (Ld) and polar components (Lp)
were adopted as shown in Table 4.33
Table 4: SFE values of measuring liquids, their dispersion and polar
components33
Tabela 4: SFE-vrednosti izmerjenih teko~in in njihovih disperzijskih
in polarnih komponent33
Measuring liquid
SFE and its components (mJ/m2)
L Ld Lp
Distilled water 72.8 21.8 51.0
Glycerine 62.7 21.2 41.5
In the Owens-Wendt model, the following equations
were used – for the dispersion component:14
( )
 
  
  
   
S
d
g g w w g
p
w
p
g
d
g
p
w
d
w
p
=
+ − +
−
(cos ) (cos ) /
/
1 1
2
(1)
and for the polar component:

 
  

S
p w w S
d
w
d
w
p
=
+ −(cos )1 2
2
(2)
where w – the surface free energy of water, wd – the
dispersion component of water SFE, wp – the polar
component of water SFE, g – the SFE of glycerine, gd
– the dispersion component of glycerine SFE, gp – the
polar component of glycerine SFE, Sp – the polar com-
ponent of SFE of the examined material, Sd – the
dispersion component of SFE of the examined material,

g – the contact angle of glycerine, 
w – the contact
angle of water.
The total value of SFE (S) was determined as a sum
of the polar and dispersion components:
S = S
p + S
d (3)
2.5 Scanning electron microscopy of the hydropho-
bized concrete
A qualitative analysis of the chemical compositions
within the main mineral components of the standard and
898 Materiali in tehnologije / Materials and technology 49 (2015) 6, 895–902
D. BARNAT-HUNEK, P. SMARZEWSKI: SURFACE FREE ENERGY OF HYDROPHOBIC COATINGS ...
Figure 1: Standard and hydrophobized surfaces HFRHPC 1 during the examination of the contact angle of a glycerine drop: a) standard sample,
b) water-soluble preparation – A, c) alkyl alkoxysilane oligomers – B
Slika 1: Navadna in hidrofobizirana povr{ina HFRHPC 1 med preiskavo kontaktnega kota glicerinske kapljice: a) standardni vzorec, b)
vodotopen preparat – A, c) alkil alkoksilan oligomeri – B
hydrophobized concretes was carried out, and the
morphology and microtopography were determined
using a scanning electron microscope FEI Quanta 250
FEG equipped with a chemical-composition analysis
system based on energy dispersion spectroscopy (EDS).
The samples were prepared in the form of thin-layer
plates, on which X-ray microanalyses were performed in
the field mode and the compositions of the elements
were determined for the seven batches of the concretes.
The sample-preparation methodology excludes the for-
mation of the microdefects associated with the cracking
of the concrete surface and hydrophobic coatings. In
order to avoid the formation of other surface defects, low
vacuum and beam energy were used during the SEM
analysis.
3 RESULTS AND ANALYSIS
3.1 Contact angle and surface free energy
The measured contact angles of water and glycerine
and the calculated values of SFE and its components are
included in Tables 5 and 6, respectively.
Graphic illustrations of the results obtained are
shown in Figures 2 to 4.
3.2 Scanning electron microscopy of the hydropho-
bized concrete
SEM microscopic analyses were performed to verify
the distribution and effectiveness of hydrophobic
coatings A, B, C in the pores of the concrete. For the
analyses, the HFHPC1 concrete was adopted, for which
D. BARNAT-HUNEK, P. SMARZEWSKI: SURFACE FREE ENERGY OF HYDROPHOBIC COATINGS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 895–902 899
Figure 2: Total values of SFE of standard concretes at the beginning
of the examination and after 40 min
Slika 2: Skupna vrednost SFE standardnih betonov na za~etku preiz-
kusa in po 40 min
Table 5: Concrete contact angles with water and glycerine
Tabela 5: Kontaktni koti vode in glicerina z betonom
Type of
samples
Type of
prepara-
tion
Contact angle
water 
w/° glicerine 
g/°
t1 = 0 t2 = 40 t1 = 0 t2 = 40
HPC1
Standard 63.7 23.0 86.7 62.0
A 83.5 74.1 86.6 77.3
B 113.5 73.0 105.6 76.1
C 122.7 98.3 127.3 102.1
SFHPC
Standard 67.8 32.4 86.2 66.7
A 86.1 75.8 89.7 69.6
B 107.5 103.3 108.7 106.7
C 96.0 67.5 98.3 66.7
HFHPC1
Standard 39.4 20.1 67.0 43.5
A 83.7 74.3 87.0 77.7
B 109.3 104.0 109.7 105.7
C 82.0 73.2 80.7 73.1
HFHPC2
Standard 42.3 12.6 71.0 45.4
A 87.0 74.5 81.6 74.7
B 103.4 111.6 106.3 104.1
C 78.0 62.1 77.0 64.3
HFHPC3
Standard 55.5 26.0 73.4 47.8
A 92.8 75.8 94.7 78.4
B 113.4 103.3 106.8 103.6
C 78.9 67.7 80.0 70.7
PFHPC
Standard 47.1 14.2 73.3 54.0
A 100.7 78.3 103.3 82.1
B 105.6 95.5 110.0 97.0
C 101.8 83.7 94.5 85.7
HPC2
Standard 53.4 10.1 78.1 56.8
A 102.5 83.4 100.4 79.3
B 116.3 100.2 122.3 94.3
C 116.1 100.5 119.8 105.6
Table 6: SFE and its components for hydrophobized and standard
concretes
Tabela 6: SFE in njene komponente pri hidrofobiziranih in
standardnih betonih
Type
of
sample
Type of
prepara-
tion
SFE component Total SFE
S/(mJ/m2)
Dispersive
Sd/(mJ/m2)
Polar
Sp/(mJ/m2)
t1 = 0 t2 = 40 t1 = 0 t2 = 40 t1 = 0 t2 = 40
HPC1
Standard 1313.9 1873.8 266.9 342.5 1 580.8 2216.3
A 72.5 83.6 0.01 0.27 72.51 83.87
B 65.5 82.0 4.95 0.45 70.45 82.45
C 48.9 71.4 4.96 1.35 53.86 72.75
SFHPC
Standard 905.8 1833.5 160.04 345.60 1065.84 2179.1
A 82.0 12.5 0.23 16.30 82.23 28.80
B 17.9 56.1 0.63 0.94 18.53 57.04
C 43.2 17.7 0.07 18.48 43.27 36.18
HFHPC
1
Standard 1385.0 595.6 233.8 36.85 1618.8 632.45
A 77.1 88.5 0.01 0.11 77.11 88.61
B 9.4 26.6 2.00 0.24 11.40 26.86
C 6.3 22.5 17.30 12.03 23.60 34.53
HFHPC
2
Standard 1572.1 860.8 290.79 82.96 1862.89 943.79
A 12.3 25.6 9.44 9.91 21.74 35.51
B 46.4 57.0 0.29 2.95 46.69 59.95
C 10.3 68.7 16.47 4.25 26.77 72.95
HFHPC
3
Standard 803.5 641.1 111.2 47.24 914.70 688.34
A 39.0 68.0 0.59 0.92 39.59 68.92
B 40.4 11.0 1.17 3.08 41.57 14.08
C 36.2 83.3 4.59 1.13 40.79 84.46
PFHPC
Standard 1426.9 1438.8 260.06 217.8 1686.96 1656.6
A 43.9 95.4 0.03 0.07 43.93 95.47
B 74.0 30.2 3.60 1.03 77.60 31.23
C 49.8 48.6 0.31 1.20 50.11 49.80
HPC2
Standard 841.1 1798.1 127.99 309.91 1631.50 2108.0
A 0.1 2.0 15.45 22.55 15.46 24.55
B 88.3 25.4 10.92 0.81 99.22 26.21
C 44.5 100.3 2.27 5.66 46.77 105.96
the best hydrophobic properties of the coatings were
obtained. Table 7 shows the analyses of the chemical
compositions in the field mode for standard HFHPC1
and hydrophobized concretes using preparations A, B
and C, performed on the basis of energy dispersion
spectroscopy (the results from the entire area of the
study).
Table 7: Chemical compositions of standard and hydrophobized
HFHPC1
Tabela 7: Kemijska sestava standardnih in hidrofobiziranih HFHPC1
HFHPC1
Component, w/%
Na2O MgO Al2O5 SiO2 K2O CaO
Standard 0.98 0.87 0.63 22.66 – 72.39
A 0.75 1.31 1.60 68.20 25.38 2.29
B 0.49 2.07 1.81 13.46 – 82.18
C 0.69 1.57 0.72 29.34 – 67.69
The microstructures of HFHPC1 and HPC2 are
shown in Figure 5.
Distribution of polysiloxane gel in the structure of
hydrophobized HFHPC1 is shown in Figure 6.
4 DISCUSSION
4.1 Contact angle and surface free energy
When analyzing the examination results presented in
Tables 5 and 6, it can be noticed that the values of the
contact angles and the surface free energy depend on the
D. BARNAT-HUNEK, P. SMARZEWSKI: SURFACE FREE ENERGY OF HYDROPHOBIC COATINGS ...
900 Materiali in tehnologije / Materials and technology 49 (2015) 6, 895–902
Figure 6: Organosilicon compounds in the microstructure of HFHPC1: a) water-soluble preparation – A, b) alkyl alkoxysilane oligomers – B,
c) methylosilicone resin – C, SEM
Slika 6: Organosilikonske spojine v mikrostrukturi HFHPC1: a) vodotopen preparat – A, b) alkil alkilosilan oligomer – B, c) metilsilikonska
smola – C, SEM
Figure 4: Total values of SFE of hydrophobized concretes after 40
min
Slika 4: Skupna vrednost SFE pri hidrofobiziranih betonih po 40 min
Figure 3: Total values of SFE of hydrophobized concretes at the
beginning of the examination (t = 0)
Slika 3: Skupna vrednost SFE pri hidrofobiziranih betonih na za~etku
preizkusa (t = 0)
Figure 5: Microstructures of HFHPC1 and HPC2 concretes prior to
hydrophobization, SEM
Slika 5: Mikrostruktura HFHPC1 in HPC2 betona pred hidrofobi-
zacijo, SEM
type of hydrophobic preparations and also on the type of
concrete.
The results of the contact-angle measurements
proved that in most cases the contact angle of glycerine
(
g) is higher than the contact angle of water (
w), and it
decreases in the course of time. The contact angles of
water obtained for the standard samples at t1 = 0 range
from 
w = 39.4 ° for HFHPC1 to 
w = 67.8 ° for SFHPC;
at t2 = 40 min, they range from 
w = 10.1 ° for HPC2 to

w = 32.4 ° for SFHPC. In the case of the standard
samples, the contact angles of glycerine are higher than
the contact angles of water (67.0 for HFHPC1 and 86.7
for HPC1 at t1 = 0; 29.9 for HFHPC2 and 62.0 for HPC1
at t2 = 40 min).
The highest contact angles of water, 
w = 122.7 °, and
glycerine, 
g = 127.3 °, at t1 = 0 were obtained for
methylosilicone resin (C) used for HPC1. In all the other
cases, the largest contact angle was obtained with alkyl
alkoxysilane (B) and it ranged from 103.4 ° to 116.3 °
(t1 = 0) and from 73.0 ° to 111.6 ° (t1 = 40 min), which
proved that a very good surface hydrophobicity was
obtained with this preparation.
For the non-hydrophobized concretes, the value of
the surface free energy is the highest and it amounts to S
= 914.69 mJ/m2 for HFHPC3 and S = 1862.93 mJ/m2 for
HFHPC2. With respect to the non-hydrophobized
concretes, the SFE value is up to 142 times higher for the
concrete with granite and up to 105 times higher for the
concrete with granodiorite aggregate than in the case of
the impregnated surfaces. The lowest value of SFE, S =
11.40 mJ/m2 (the weakest adhesion properties), was
obtained for the HFHPC1 concrete hydrophobized with
alkyl alkoxysilane in the organic solvent (B). In the cases
of the PFHPC and HPC2 concretes with granite aggre-
gate, preparation B showed much lower hydrophobic
properties, since the value of SFE was higher by
27.55–83.77 mJ/m2 than those of the other preparations.
The concretes with granite aggregate obtained the
highest hydrophobicity using the water-based solution of
methylosilicone resin in potassium hydroxide (S =
15.46–43.93 mJ/m2).
Furthermore, it can be noticed that in all the cases the
SFE dispersive component (Sd) constitutes a far larger
share in the total SFE value (S) than the polar compo-
nent (Sp).
Considering the variations in time it was observed
that in the course of time (after 40 minutes) the total SFE
value decreased in the case of the concretes containing
steel and polypropylene fibers with both granodiorite and
granite aggregates. In the other concretes without fibers,
the SFE value decreased after 40 min, proving a decrease
in the hydrophobicity.
4.2 Scanning electron microscopy of the hydropho-
bized concrete
A uniformly distributed silicon coating is formed in
the microstructure of the concrete due to the water-
soluble preparation (A); however, it is too thick and
shows numerous cracks. Analyzing the microstructure of
the coating obtained from preparation A, it can be con-
cluded that it creates the sealing for the fine subsurface
pores of the concrete, resulting in a decrease in the
water-vapor permeability of the concrete. The coating
that is too thick does not provide an adequate adhesion to
the concrete minerals, making the hydrophobization less
effective. As a result of frost or chemical corrosion the
coating will be damaged within a very short period of
time.
The methylosilicone-resin coating (C) is of a homo-
geneous nature, as it is made from fine particles and does
not create the sealing for the HFHPC1 structure. This
ensures a good efficiency of the impregnation and does
not disturb the diffusion of gases and vapors.
The alkyl-alkoxysilane-oligomer preparation (B),
based on an organic solvent, is characterized by a fine-
pore structure composed of even smaller particles than
the macromolecular resin (C). Microscopic observations
proved that the coating was uniformly distributed and it
did not show any defects.
5 CONCLUSION
The measurement of the contact angle is one of the
methods for monitoring the changes in the wettability of
hydrophobized building materials.
The use of various preparations results in obtaining
different wetting and adhesion properties of HFHPC,
determined by the surface free energy. The SFE value
decreases significantly on a hydrophobized surface, in
particular when using a small molecule oligomer for the
concretes with granodiorite aggregate or a water-based
solution of methylosilicone resin in the potassium hydro-
xide for the concretes with granite aggregate. As proved,
this may result from a more effective hydrophobization,
due to macromolecular resins, of the materials characte-
rized by the pores of bigger diameters like granite-aggre-
gate-based concretes. Adding polypropylene and steel
fibers contributes to an increased hydrophobicity in the
course of time.
The application of organosilicon compounds in the
near-concrete surface area results in a reduction in the
SFE and surface tension of the concrete, depending on
the chemical composition of the preparation. This causes
a reduced penetration of the corrosive substances into the
concrete structure, thus affecting its durability.
Evenly distributed silica gel is formed in the fine
pores of the high-strength concrete by organic-solvent-
based preparations B and C. This coating is characte-
rized by a fine-pore microstructure; it shows an adequate
adhesion to the concrete surface. Such properties of
silicone resin ensure a good vapor permeability and
effective hydrophobization.
D. BARNAT-HUNEK, P. SMARZEWSKI: SURFACE FREE ENERGY OF HYDROPHOBIC COATINGS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 895–902 901
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D. BARNAT-HUNEK, P. SMARZEWSKI: SURFACE FREE ENERGY OF HYDROPHOBIC COATINGS ...
902 Materiali in tehnologije / Materials and technology 49 (2015) 6, 895–902
J. FALKUS, K. MI£KOWSKA-PISZCZEK: DEVELOPING CONTINUOUS-CASTING-PROCESS CONTROL BASED ...
903–912
DEVELOPING CONTINUOUS-CASTING-PROCESS CONTROL
BASED ON ADVANCED MATHEMATICAL MODELLING
UPORABA NAPREDNEGA MATEMATI^NEGA MODELIRANJA ZA
RAZVOJ KONTROLE POSTOPKA KONTINUIRNEGA ULIVANJA
Jan Falkus, Katarzyna Mi³kowska-Piszczek
Faculty of Metals Engineering and Industrial Computer Science, AGH University of Science and Technology, Al. Mickiewicza 30, 30-059
Kraków, Poland
jfalkus@agh.edu.pl
Prejem rokopisa – received: 2014-08-03; sprejem za objavo – accepted for publication: 2014-12-16
doi:10.17222/mit.2014.176
The method of continuous casting of steel – due to its ability to maximize the yield of liquid steel, along with substantially
reducing the energy consumption of the production process – has become the fundamental method for obtaining steel
semi-products. Nowadays, over 90 % of the global steel is cast with the continuous method. In recent years the ability to
numerically model metallurgical processes – including the continuous process of steel casting – has been very important for
creating new technologies, along with modifying those that already exist. The mathematical modelling of solidification
processes with numerical methods allowed a full comprehensive reconstruction of the complex physical and chemical nature of
the solidification processes. However, having to formulate a numerical model of the continuous-casting process is an extremely
complex task because the requirements stipulate that a correct set of material parameters, along with the process data, have to be
implemented. As regards the formulation of a mathematical model of the steel continuous-casting process, a comprehensive
description of the heat transfer during the continuous casting is an important item. The complexity of this issue requires that
conscious simplifications are made when formulating mathematical models for the calculation of the cast-strand solidification
process. The number and type of the simplifications – which are necessary in this case – are the keys to the correctness of the
results obtained, and they also influence the scope and accuracy when it comes to verifying the model. It is crucial to define the
problem on the basis of the finite-element theory, and the elimination of numerical errors is obviously a necessary condition to
ensure the correctness of the analysis performed. This paper contains a description of a number of solutions that are based on
the finite-element method (FEM) for the models using the Euler, Lagrange and MiLE meshes. All the model concepts are
illustrated with the examples of the calculations that were completed using the actual industrial data, along with the properties
of the materials as determined by laboratory tests. The pivot of the considerations conducted is related to the verification of the
correctness of the calculations, together with the sensitivity analysis of individual model types. The conclusions present an
assessment of the progress of the current numerical models of the continuous-casting process, along with the directions for their
further development.
Keywords: continuous casting, mathematical modelling, solidification, process control
Metoda kontinuirnega ulivanja jekla je postala osnovna metoda za pridobivanje jeklenih polproizvodov zaradi mo`nosti
maksimalnega izkoristka staljenega jekla ter zmanj{anja porabe energije za proizvodni proces. Danes je okrog 90 % svetovne
proizvodnje jekla ulitega z metodo kontinuirnega ulivanja. V preteklih letih je bila mo`nost modeliranja postopka ulivanja jekla
zelo pomembna pri nastajanju novih tehnologij vklju~no z modificiranjem `e obstoje~ih. Matemati~no modeliranje procesa
strjevanja z numeri~nimi metodami je omogo~ilo celovito rekonstrukcijo kompleksne fizikalne in kemijske narave procesa
strjevanja. Vendar pa je postavitev numeri~nega modela postopka kontinuirnega ulivanja zelo zahtevna naloga, ker zahteve
dolo~ajo, da je treba uporabiti pravilne parametre materiala vklju~no s podatki procesa. Glede na postavitev matemati~nega
modela postopka kontinuirnega ulivanja je pomembna postavka celovit opis prenosa toplote med kontinuirnim ulivanjem.
Kompleksnost tega vpra{anja zahteva zavestne poenostavitve pri postavitvi matemati~nih modelov za izra~un procesa strjevanja
`ile. [tevilo in vrsta poenostavitve, ki so v tem primeru nujne, sta klju~ za pravilnost dobljenih rezultatov, vplivajo pa tudi na
obseg in natan~nost, ko pride do preverjanja modela. Klju~nega pomena je definiranje problema, ki temelji na teoriji metode
kon~nih elementov, odprava numeri~nih napak pa je o~itno potrebni pogoj za zagotovitev pravilnosti izvr{enih analiz. ^lanek
vsebuje opis {tevilnih re{itev, ki temeljijo na metodi kon~nih elementov (FEM), za modele, ki uporabljajo Euler-jevo,
Lagrange-jevo in MiLE-mre`o. Prikazani so vsi koncepti modelov z izra~uni, ki so bili dopolnjeni z uporabo pravih industrij-
skih podatkov, skupaj z lastnostmi materialov, kot je bilo dolo~eno v laboratorijskih preizkusih. Te`i{~e izvr{enih obravnav je
povezano s preverjanjem pravilnosti izra~unov, skupaj z ob~utljivo analizo posameznih vrst modelov. V sklepih je predstavljena
ocena glede napredovanja sedanjih numeri~nih modelov postopka kontinuirnega ulivanja vklju~no z napotki za njihov nadaljnji
razvoj.
Klju~ne besede: kontinuirno ulivanje, matemati~no modeliranje, strjevanje, kontrola procesa
1 INTRODUCTION
The history of the development of the steel conti-
nuous-casting process is an interesting example of the
implementation of an idea that was well ahead of its
time. After H. Bessemer had applied for the first patent
in 1846, 34 years had to pass before the inventor made
his first trials of steel casting at the semi-technical scale.1
The reasons why the idea of the metal continuous-cast-
ing process was so inspiring for many metallurgical
engineers are well known. The implementation of this
process to steel casting significantly shortened the
process line, definitely reducing both the process and the
investment costs. The other indisputable reason for the
advantage of steel continuous casting over ingot casting
is the yield. Thanks to a yield increase of over 10 %, it
Materiali in tehnologije / Materials and technology 49 (2015) 6, 903–912 903
UDK 519.61/.64:621.74.047:536.421.4 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)903(2015)
became possible to increase the steelworks output
without increasing the steelmaking furnace capacity.
The reasons why the implementation of the steel
continuous-casting process for the production took so
long are as well understood as the reasons for its
implementation. From the outset, the mechanical
characteristics of a strand cast have been the main
difficulty, as a strand is prone to cracking, especially at
the initial formation stage; in an extreme case it may
even break out. A small shell thickness of a strand
leaving the mould is of key importance. This problem is
exceptionally serious as regards steel and it arises from a
relatively low value of the heat-conductivity coefficient
of steel as compared to the other metals. For most
non-ferrous metals, the liquid core of a strand cast
continuously is much shorter and the solidification
process ends right under the mould. For steel, the length
of the liquid core most often ranges from a few metres to
something between ten and twenty. This fact brings
about the necessity of using advanced techniques of the
process control and ensuring stable process conditions.
Note that despite the indisputable successes, the
current development of the steel continuous-casting
process cannot be considered finished. The problems
related to the strand breakout or the occurrence of
various defects of cast strands are still valid. The
financial means earmarked for scientific research related
to the continuous casting of steel indirectly prove that the
problems still exist. An example of a research area that
constitutes a huge challenge, both in theory and in
practice, is the continuous casting of peritectic steels.
Engineering new steel grades also involves the need for
developing their casting process from scratch. Therefore,
any projects related to the improvement of the
mathematical description of the steel continuous-casting
process are valuable for facilitating the process control
problem.
2 STEEL CONTINUOUS-CASTING PROCESS AS
AN OBJECT OF MODELLING
The effect of a continuous-casting-machine design on
the method of mathematical modelling of the steel soli-
dification process is significant. Without the knowledge
of the machine technical documentation, it is difficult to
make any attempt at modelling, aiming at obtaining the
results intended for use for the process control.
Despite the existing differences in the design of
individual machines, one should note that it is possible to
generalise the problem and form the construction of a
mathematical model assuming the occurrence of two
cooling zones of the strand, i.e.:
• the primary cooling zone (including the mould)
• the secondary cooling zone (including the so-called
cooling chamber and the division into the cooling
zones independent of the cooling chamber).
Each of the two mentioned zones is characterised by
two groups of parameters – the first one may be defined
as a geometrical characteristic, and the other one as a set
of parameters related to the heat transfer and the
properties of the steel cast. The description of the
heat-transfer model for the steel continuous-casting
process is a complex task as all three heat-transfer
mechanisms occur – conduction, radiation and
convection2–4. The following processes impact the heat
transfer in the primary zone:
• conduction and convection in the liquid-steel area,
• conduction in the solidified shell,
• heat transfer between the outer layer of the solidified
shell and the mould wall surface through an air gap
forming in the mould,
• heat conduction in the mould,
• heat transfer in the mould between the channel walls
and the cooling water.
The main processes in the secondary cooling zone
are:
• heat transfer by convection,
• heat transfer by radiation,
• heat transfer by conduction between the solidifying
strand and the rolls.
Additionally, thermal effects related to the phase
transitions that accompany the solidification have a
significant influence on the heat-transfer model.
The heat-transfer model was used for further
considerations, wherein the temperature field could be
determined by solving the Fourier Equation:
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
( )
  
c T
t x
T
x y
T
y z
T
z
p
= ⎛⎝
⎜ ⎞
⎠
⎟ +
⎛
⎝
⎜
⎞
⎠
⎟ + ⎛⎝
⎜ ⎞
⎠
⎟ +Q (1)
where  is the density, kg m–1; cp is the specific heat,
kJ kg–1 K–1;  is the thermal conductivity, W m–1 K–1;
t is the time, s; T is the temperature, K; Q is the heat-
source term, W m–3 ; x, y, z are 3D coordinate axes. The
solution of the Fourier Equation should meet those
boundary conditions declared on the strand surface. In
the numerical model of the steel continuous-casting pro-
cess discussed, these boundary conditions may be
declared in three various ways. The Equation below
describes the second and third types of boundary
conditions:
Q Flux T T T T= + − + − ( ) ( )a a
4 4 (2)
where Flux is the heat flux, W m–2;  is the heat-
transfer coefficient, W m–2 K–1; Ta is the ambient
temperature, K;  is the Stefan-Bolzmann constant,
W m–2 K–4;  is the emissivity. The heat flux in the
model may be defined directly as the Flux value (the
Neumann condition), as well as with the convection
( – substitute heat-transfer coefficient) and with the
radiation models ( – emissivity).
It is a complicated task to formulate a model of the
heat transfer in the mould. The main difficulty is the
J. FALKUS, K. MI£KOWSKA-PISZCZEK: DEVELOPING CONTINUOUS-CASTING-PROCESS CONTROL BASED ...
904 Materiali in tehnologije / Materials and technology 49 (2015) 6, 903–912
effect of an air gap formed immediately after the steel
has cooled to the solidus temperature, under the liquid-
steel meniscus surface. The oscillatory movement of the
mould and, in addition, the strand movement in the
mould with a variable casting speed, strongly influence
the actual dimensions of the gap. The presence of the
mould powder and gases within the gap makes this des-
cription even more problematic. The physical properties
and the chemical composition of the mould powders
applied significantly influence the heat-resistance value.
The development of the air gap causes a very high tem-
perature gradient between the solidifying strand shell
and the mould wall. Also, the change in the air-gap
dimensions during the continuous-casting process may
influence the stability of the process, having an impact
on the thickness of the shell leaving the mould. The
dimensions of the air gap increase with the distance from
the liquid-steel meniscus, consequently causing an in-
crease in the heat resistance (Figure 1).
The heat-transport mechanism between the soli-
difying strand shell and the mould wall may be divided
into two components – conduction and radiation. The
convection in this area may be neglected due to the small
size of the air gap. The total air-transfer coefficient on
the strand mould path may be presented with the
following Equation:
qwk = qc + qr (3)
where:
qwk – total heat-flux density,
qc – conducted heat-flux density,
qr – density of the heat flux transferred by radiation.
TM – temperature of the mould, TS – temperature of
the strand surface, TSS – temperature of the solid layer of
the slag, TMS – the melting point of the slag, TLS – tem-
perature of the liquid layer of the slag, TSO – solidus
temperature.5
Finally, the flux density of the heat flowing from the
strand to the mould is calculated from the Newton law:
( )q T T
r
T Tc
wl kr
sz
sz wl kr=
−
= − (4)
where:
sz – coefficient of the heat transfer through the air gap,
rsz – air gap thermal resistance.
The outer side of the mould is intensely cooled with
water, flowing through the channels. Here, the heat is
transferred by forced convection. The calculation of the
heat-transfer coefficient, with the water cooling in the
mould channels, on the basis of the available formulas, is
complex because of the method of heat transfer to the
water flowing through the channel. Assuming the distri-
bution of the channel density at the perimeter of the
mould varies, it is possible to obtain a few cooling pro-
grams. To determine the average heat-transfer coeffi-
cient, the following formula may be applied to the outer
surface of the mould:6


w
w
k
k=
Nu
d
x (5)
where:
xk – water-cooled mould-surface share,
dk – mould-channel diameter,
w – heat-transfer coefficient for water,
Nu – Nusselt number.
For the forced convection (the water flowing in the
mould channels) the Nusselt number is represented by
the relationship between the Reynolds (Re) number and
the Prandtl (Pr) number:
Nu = f (Re, Pr) (6)
The Mikheyev formula may be selected to determine
the Nusselt number:7
Nu Re Pr
Pr
Pr
=
⎛
⎝
⎜
⎞
⎠
⎟0 021 0 8 0 43
0 25
. . .
.
w
w
s
(7)
The Reynolds Re and Prandlt Pr numbers in Equation
(7) are determined at the water properties for the mould-
wall temperature (the s index) and the mean water tem-
perature in the channel (the w index).
After leaving the mould, the slab surface is cooled
with a water spray and in the air. The heat flux that is
carried away from the surface of the solidifying strand is
proportional to the temperature difference between the
strand surface and the cooling medium temperature. In
this zone it is recommended to maintain the cooling
intensity which leads to gradual temperature changes. It
ensures that the cracks generated by thermal stresses are
avoided.8
In the secondary cooling zone, the heat exchange
with the environment is accompanied by several mecha-
nisms:
• direct impact of the water stream on the strand,
• cooling with the water flowing on the surface and
with defected drops,
J. FALKUS, K. MI£KOWSKA-PISZCZEK: DEVELOPING CONTINUOUS-CASTING-PROCESS CONTROL BASED ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 903–912 905
Figure 1: Structure of the layer between the mould and the shell
Slika 1: Zgradba podro~ja med kokilo in strjeno skorjo
• cooling with the water lingering on the roll in the
formed hollow,
• cooling the places located directly under the roll with
the ambient air,
• cooling through the contact faces of the roll.
The heat transfer from the surface of the solidifying
strand is a complex process due to the nature of the heat
exchange accompanying the boiling effect. When the
surface of the strand leaving the mould is cooled, film
boiling prevails. At this stage, the main mechanism of
the heat transfer is the conduction through the vapour
film resulting from supplying water to the hot surface of
the strand via the nozzles. The heat-transfer mechanism
is also accompanied by radiation, therefore, the lower the
strand surface temperature, the lower is the value of the
heat-transfer coefficient.9 Drops of cooling water that fall
onto the hot strand surface evaporate, forming a film that
restricts water access at this spot. The water momentum
is the highest in the centre of the cooling area, enabling
the vapour film to be broken and allowing a direct
contact between the water and the cooled strand surface.
It intensifies the cooling process in this area, therefore,
the amount of the heat received strongly depends on the
liquid velocity.5
During further cooling, as the temperature falls below
the Leidenfrost point, the transition-boiling effect occurs.
In the vapour film, vapour bubbles form, consequently
leading to a break and decay of the vapour film. It causes
an increase in the heat-transfer coefficient.10 During fur-
ther cooling we observe the effect of bubble boiling,
along with a further decline in the heat-transfer coeffi-
cient. After the completion of the bubble-boiling stage
and the cessation of the conditions for vapour forming,
the strand surface is cooled by forced convection.11
Figure 2 presents the change in the heat-transfer coeffi-
cient for the water-boiling process.
A very important mechanism of the heat exchange in
the secondary cooling zone is radiation due to the surface
temperature of the strand that leaves the mould. To ob-
tain a complete description of the effects, it is necessary
to consider the heat-flux density resulting from the heat
transfer by radiation. The radiation term may be ex-
pressed with the Stefan-Boltzman law:
qrad = wl  (Twl
4 – Ta
4) (8)
where:
qrad – heat-flux density lost by the strand surface to the
environment,
 – Stefan-Boltzman radiation constant,
wl – total emissivity of the strand surface,
Ta – ambient temperature.
The strand emissivity-coefficient value may be
adopted in the range of 0–1. Due to a simplification, the
constant may be assumed to be 0.85.
The flux density of the heat transferred to the cooling
water may be calculated from the following relationship:
qspray = spray (Twl – Tspray) (9)
where:
spray – coefficient of heat transfer by water,
Twl – strand surface temperature,
Tspray – temperature of the water flowing through the
nozzles.
In order to calculate the flux density of the heat trans-
ferred, it is necessary to know the heat-transfer coeffi-
cient spray. The relationship describing the heat-transfer
coefficient of the heat transferred as a result of the
water-spray impact may be expressed in a general form:4
( ) spray sprayc spray= −AV bt 1 (10)
where:
spray – heat-transfer coefficient,
A, b, c – empirical constants,
Vspray
c – cooling-water flux,
tspray – temperature of the water flowing through the
nozzles.
The heat-transfer coefficient depends on the condi-
tions that occur during the contact of the water with the
surface of the solidifying strand. The water-flux density,
the velocity of the water flowing out of the spray nozzle,
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906 Materiali in tehnologije / Materials and technology 49 (2015) 6, 903–912
Figure 3: Heat-transfer coefficient during the cooling process with a
water spray for Vspray
c = 3 dm3 m2 s–1
Slika 3: Koeficient prenosa toplote med ohlajanjem z brizganjem vode
pri Vspray
c = 3 dm3 m 2 s–1
Figure 2: Flow diagram for the heat-transfer coefficient changing dur-
ing the boiling process3
Slika 2: Potek spreminjanja koeficienta prenosa toplote med vrenjem3
the nozzle type and the water pressure all influence the
value of the coefficient. A typical course of the changes
in the heat-transfer coefficient as a function of the tem-
perature is presented in Figure 3.
If the solidifying cast-strand surface temperature is
not known, a simplified formula may be applied:1
 spray spray= + +10 107 0688v v V( . )  (11)
where: v – water-drop speed.
Equation (11) may be applied for water fluxes from
0.3 dm3 m –2 s–1 to 9.0 dm3 m –2 s–1 and speeds from 11 m s–1
to 32 m s–1. The values obtained for a flux of = 3 dm3 m–2 s–1
and the average water speeds are about 600 W m–2 K–1.
In the zone beyond the water sprays, in most of the
continuous-casting machines, after a strand leaves the
secondary-cooling chamber, a heat transfer occurs bet-
ween the hot surface of the strand cooled and the
ambient atmospheric air. Radiation is the prevailing
mechanism of heat transfer; however, the heat is partly
transported from the strand surface with free or forced
convection.3
Even an abbreviated description of the heat-transfer
mechanisms in the steel continuous-casting process indi-
cates two basic difficulties accompanying the mathema-
tical modelling of the process. First, it is impossible to
identify every component of the model. If the description
is very detailed, taking into account all the partial pro-
cesses, then, at the verification stage, simplifications are
usually made and substitute heat-transfer coefficients are
determined. However, these simplifications do not cover
all the problems that are encountered. The second stage
of verification always requires us to find the relationship
between the identified values of the heat-transfer coeffi-
cients and the intensity of the cooling-media supply. The
fulfilment of this requirement is the key to the correct-
ness of the forecast of each continuous-casting model
and it is a problem most difficult to solve.
3 TYPES OF MATHEMATICAL MODELS
3.1 Lagrangian mesh model
The motion of a continuum may be described from
two different standpoints – either with the Lagrangian or
Eulerian method. The motion analysis, together with
these methods, allows one to determine the position of
any moving element of a body; in addition, these ele-
ments are treated as material points. The Lagrangian
description is preferred in solid-body mechanics. In this
model the "observer" follows the body’s material points,
observing their temperature and velocity being affected
by specific boundary conditions. The initial positions of
the material points influence the course of the process.
As a result, the history of a numerically modelled pro-
cess can be created on the basis of the temperature field
and velocity (strain) field, allowing a calculation of the
stress field in the body and the changes in the micro-
structure and porosity. Another advantage of the Lagran-
gian description is the fact that the mesh deforms
together with the body, which is particularly relevant in
the numerical modelling of body strains. In the Lagran-
gian method an object’s motion (physical continuum) is
described by the dependence of the coordinates of all the
particles on the time and the initial position for  = 0 or
the initial time 1. A necessary condition is also the
determination of all the properties of the object, taking
into account the positions of its points at a specific time.
By eliminating the time, we can obtain an equation of
the trajectory of the object particles. In this method, the
observer moving together with the object maintains
contact with the same object particles at the set time.5
This is schematically presented in Figure 4.
3.2 Eulerian mesh model
The Eulerian method is usually used in fluid mecha-
nics. Contrary to the Lagrangian description, the refe-
rence point is located outside the area, so the "observer"
follows the behaviour of the area from the outside. In the
Eulerian method, bodies move on the background of a
mesh that represents a specific area. If, during the motion
of the fluid in the system, the temperature, the velocity
vector and the concentrations do not change with the
time at the fixed points, the external "observer" interprets
the system as a steady system. Only by analysing the
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Materiali in tehnologije / Materials and technology 49 (2015) 6, 903–912 907
Figure 5: Euler Method – description of the motion of the object5
Slika 5: Euler-jeva metoda – opis gibanja predmeta5
Figure 4: Lagrangian Method – description of the motion of the
object5
Slika 4: Lagrange-jeva metoda – opis gibanja predmeta5
data concerning the temperature field and the velocity,
does it turn out that, in fact, the process of the system is
transient.
The time and coordinates of the current configuration
are independent variables. Such a description is called
the Eulerian description or spatial description because
the changes in the medium parameters at a specific place
in the space are tracked. In Figure 5 this place is deter-
mined by the coordinates x, y, z. At the initial moment,
these coordinates describe the position of element P1 in
the object, whereas after time , at the moment of the
current observation, they describe the position of a diffe-
rent element, P2. So, at various times various elements
are located at the same place. This method of description
is very favourable for the objects whose initial configu-
ration has no influence, or their influence on the state of
the configuration at a later time is of little significance. It
is particularly significant for the description of the move-
ment of fluids, i.e., liquids and gases.9
3.3 MiLE method
In the process of steel continuous casting, the solid
and liquid phases occur simultaneously and, in the
metallurgical part of the strand, these phases are
adjacent. So, the problem of numerical modelling of this
part of the process must be solved with both methods –
the Lagrangian and Eulerian methods – or their com-
bination. In the numerical model using the finite-element
method, the Eulerian approximation is used to describe
the temperature field and the liquid-metal flow. A fixed
mesh with invariable geometry may be used here. The
speed and temperature vectors are determined in fixed
node points. To model the strain and stress of the soli-
dified metal simultaneously, it is necessary to use the
description with the Lagrangian coordinates.
At present, a new method combining the Lagrangian
and Eulerian methods is used for a numerical simulation
of the steel continuous-casting process – the MiLE
(mixed Lagrangian-Eulerian method) method. In this
method the continuous-casting-process area is divided
into two parts. The upper, stationary, part is described by
the Eulerian scheme. The lower part is described by the
Lagrangian scheme as it moves at the casting speed. To
maintain the strand integrity, at consecutive time steps,
layers are placed under the upper part; these layers bring,
into the lower part, a mass of steel equivalent to the mass
of the steel brought to the mould from the tundish.
The MiLE method allows linked modelling of the
non-stationary temperature fields and of the liquid-steel
flow in the process of steel continuous casting and
modelling of the stresses caused by thermal effects.
The principle of the MiLE algorithm is presented in
Figure 6. The first action is the division of the strand
into two areas. After starting the casting process, the
lower area moves downwards, whereas the upper area
stays in its initial position. The calculations within this
area are performed with the Eulerian model. To maintain
the integrity between the areas, the mesh must be
successively supplemented with additional layers of
elements, forming a dynamic area. Therefore, a certain
number of elements with the initial thickness of zero are
stored within the area between the first two areas (the
third area). In order to ensure the integrity of the tem-
perature field and the velocity field between areas 1 and
3, in all the nodes of the “stored” (not yet used) layers,
momentary boundary conditions are imposed. The same
values of the temperature and velocity are assigned to the
nodes with identical coordinates.5,12 In areas 2 and 3, the
Lagrangian model is used for the calculations (Figure 6).
4 SENSITIVITY OF STEEL-CONTINUOUS-
CASTING-PROCESS MODELS
4.1 Numerical parameters
The procedure of a mesh application arises from
calculation tests and not from formal guidelines. For
each new case, applying a FEM mesh is an individual
issue for a specific model. For a new mesh (use of
different elements or their sizes), a number of calculation
tests should be performed in order to eliminate potential
numerical errors. At the stage of the FEM mesh design,
the places of the implementation of the boundary con-
ditions should be taken into account due to the refine-
ment of the mesh at those places, regardless of the
applied mesh type. An observation of the obtained
results, in particular the temperature distribution, allows
an easy identification of the places where a numerical
error related to the mesh-element size is generated.
Usually a temperature-distribution asymmetry is notice-
able. In addition, a high jump in the temperature values
can be noticed. This is neither correlated with the boun-
dary conditions nor justified by the change in the main
process parameters in this area.
4.2 Material parameters
The thermophysical properties of steel – as deter-
mined by experimental research or calculated using
thermodynamic databases – are the key input parameters
for building a numerical model of the continuous-casting
process. Based on the chemical composition of a steel
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908 Materiali in tehnologije / Materials and technology 49 (2015) 6, 903–912
Figure 6: MiLE algorithm5
Slika 6: MiLE-algoritem5
grade, and the algorithms implemented in the thermo-
dynamic databases, a number of material properties, i.e.,
enthalpy, thermal conductivity, density, viscosity, solidus
and liquidus temperatures, may be determined. The cal-
culation of the material parameters of the steel grade
tested based on its chemical composition is a common
approach. However, it should be emphasised that the
temperature distribution obtained with the numerical
modelling is extremely sensitive to any changes in the
basic thermophysical properties.
The starting point for the sensitivity analysis was the
base variant, which took account of the specific heat
values from the experimental tests, along with the latent
heat value that was implemented in the numerical model
as a numerical value. The calculations obtained for the
base variant allowed the actual continuous-casting pro-
cess for the S235 steel grade to be fully mapped for the
declared strand casting speed of 1 m/min. The other
variants were defined by the cp declaration, as follows:
Variant 1, where the magnetic-transformation heat and
the austenite-ferrite-transformation heat were consi-
dered. The heat of the fusion was declared as 113
kJ/kg.
Variant 2, where the magnetic-transformation heat, the
austenite-ferrite-transformation heat and, addition-
ally, the peritectic-transformation heat were consi-
dered. The heat of the fusion was declared as 113
kJ/kg.
Variant 3, where the magnetic-transformation heat, the
austenite-ferrite-transformation heat, the peritectic-
transformation heat and the heat of fusion were con-
sidered.
The obtained calculation results are presented in
Table 1.
Table 1: Metallurgical length and shell thickness calculated for the
selected variants
Tabela 1: Metalur{ka dol`ina in debelina skorje, izra~unana za
izbrane variante
Specific heat,
kJ/(kg K)
Metallurgical
length, m
Thickness of the
shell, cm
Basic 12.6 2.2
Variant 1 12.6 2.2
Variant 2 13.7 2.18
Variant 3 11.4 2.21
The method of determining the specific heat value,
along with the solidifying heat, has a significant influ-
ence on the obtained results of the numerical model. By
taking into account the values coming from the experi-
mental research, the amount of the heat accompanying
the individual phase transformations occurring during
the solidification process may be accurately determined.
In addition, the method determining the latent heat is
very important. In extreme cases, an error in the en-
thalpy-value determination in relation to the verified
specific-heat values may be as high as 178 °C.
4.3 Process parameters (boundary conditions)
The influence of the boundary conditions assumed
for the calculations, apart from the thermophysical para-
meters, is the most important element of any mathema-
tical model of the continuous-casting process. However,
in professional references there is an extremely huge
range of the values assumed. To illustrate the scale of the
problem, the values of the heat-transfer coefficients
(HTC) for 12 various models describing the heat transfer
in a mould were collected in Table 2.3,13–25 Next, cal-
culations with the specified HTC values were performed,
maintaining the constant value of the other process para-
meters. The obtained calculation results are presented in
Table 3.
Table 2: Heat-transfer coefficient in the primary cooling zone
Tabela 2: Koeficient prenosa toplote v primarni hladilni coni
No. Average or maximum value Reference
Group I – the average HTC value along the whole length of
the mould
1 h = 1200 W/(m2 K) 13,14
2 h = 1300 W/(m2 K) 15
3 h = 1500 W/(m2 K) 16,17
Group II – two values of heat-transfer coefficients
4
h1 = 1163 W/(m2 K) for z = 0.6 m
h2 = 1395.6 W/(m2 K) for z > 0.6 m
18,19
Group III – linear variable
5 h = 2000–800 W/(m2 K) 20
6 h = 1500–600 W/(m2 K) 20
Group IV – variable (various heat-transfer mechanisms)
7 hmax = 1300 W/(m2 K) 21
8 hmax = 2500 W/(m2 K) 22
9 hmax = 2000 W/(m2 K) 23
10 hmax = 1300 W/(m2 K) 3
11 hmax = 3097 W/(m2 K) 24
12 hmax = 1600 W/(m2 K) 25
Table 3: Comparison of the shell thicknesses and temperatures for the
selected models of the heat-transfer coefficient
Tabela 3: Primerjava debeline skorje in temperature pri izbranih
modelih koeficienta prenosa toplote
Model Thickness of the shell afterleaving the mould, cm Temperature, °C
1 2 836
2 2.3 812
3 2.75 772
4 2.27 833
5 2.38 900
6 1.94 956
7 1.98 976
8 1.92 1050
9 1.86 1090
10 1.82 1099
11 2.51 874
12 2.52 938
It should be stressed that model 12 is the author’s
own model, verified under industrial conditions. Its accu-
racy was checked with test-strand temperature measure-
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Materiali in tehnologije / Materials and technology 49 (2015) 6, 903–912 909
ments. The results of this verification are presented in
Table 4.
Table 4: Values of the strand surface temperature calculated and
measured at the reference points
Tabela 4: Vrednosti izmerjene in izra~unane temperature na povr{ini
`ile v referen~nih to~kah
The average
measured
temperature, °C
The calculated
temperature, °C
Measurement point I 855 861
Measurement point II 905 912
5 STRATEGY OF SELECTING THE STRAND
COOLING PARAMETER
From the perspective of the applicable technology
and the design of the slab-continuous-casting machine, it
is essential to maintain the metallurgical length at a con-
stant level. This is mostly related to the location of the
soft reduction zone. The use of soft reduction of a strand
allows an elimination of the axial porosity, consequently,
improving the quality of the steel semi-products ob-
tained. Maintaining a comparable metallurgical length
for various strand casting speeds guarantees the safety
during the continuous-casting process as the strand is
fully solidified before shearing. For most of the conti-
nuous-casting machines, the best area where the total
solidification of the strand should occur is the area
before the exit from the secondary-cooling chamber. It is
related to the location of the soft-reduction zone and the
location of the last point of strand straightening.
The first step of the proposed method for determining
the new cooling values is to calculate the influence of a
change in the strand withdrawal speed on the metallur-
gical length, while verifying the shell thickness and the
temperature at the reference points. An analysis of the
sensitivity of the numerical model of the steel conti-
nuous-casting process to the casting speed allows a
determination of the percentage impact of the change in
the casting speed on the metallurgical length. Knowing
this dependence allows us to calculate the influence of a
change in the cooling parameters, which must always be
correlated with the casting speed.
The simplest method of determining the new cooling
values for the spray zones is using the percentage change
in individual heat-transfer coefficients, taking into
account the percentage change in the metallurgical
length as a function of the casting speed. From the nu-
merical perspective, the described method is an effective
approach. However, determining the new sets of cooling
values must be correlated with the existing technology
and the relationships of the changes in the water-flow
rates for individual spray zones. When the new values of
heat-transfer coefficients were determined, minimum
changes were applied to the heat-transfer coefficients
within the first two cooling zones; more significant
changes were introduced to the other spray zones, while
the temperature increase within the spray zones was con-
trolled. Such an approach is extremely important for
maintaining the safety of the steel continuous-casting
process.
In the example discussed, for a speed of 1 m/min, the
set of cooling factors remained unchanged, because the
desired values of the metallurgical length, the tempera-
ture at the reference points and the shell thickness after
leaving the mould were obtained. The simulation for the
speed of 1 m/min was considered the reference for fur-
ther considerations. Figure 7 presents a diagram of the
algorithm for determining the new values of heat-transfer
coefficients.
The starting point of the algorithm includes the heat-
transfer coefficients for seven spray zones; these were
calculated on the basis of the actual flows of the water
for the three speeds examined. Next, one heat-transfer
coefficient for the whole secondary cooling zone average
was calculated, using the proposed dependence presented
with the formula below:
 average =
∑
∑
a s
s
n n
n
n
n
1
1
(12)
where:
n – heat-transfer coefficient for the selected spray
zone,
sn – surface area of the specific zone,
n – zone number.
The average heat-transfer coefficients for the secon-
dary cooling zone were calculated from dependence 12.
They were (305, 395 and 410) W m–2 K–1, for the speeds
of (0.6, 0.8 and 1) m min–1, respectively. Knowing the
percentage impact of the casting speed on the metallur-
gical length, the new heat-transfer coefficients for the
secondary cooling zone for the speeds of 0.6 m min–1 and
0.8 m min–1 were calculated; similar metallurgical
lengths were obtained as for the speed of 1 m min–1. The
following dependence was used:


' average
average
=
k
(13)
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910 Materiali in tehnologije / Materials and technology 49 (2015) 6, 903–912
Figure 7: Diagram for a calculation of new heat-transfer coefficients25
Slika 7: Prikaz izra~una novih koeficientov prenosa toplote25
where: k – coefficient based on the percentage relation-
ship for the metallurgical length as a function of the
casting speed.
The new heat-transfer coefficients for the secondary
cooling zone were calculated from dependence 13. For
the speed of 0.8 m min–1 the average heat-transfer
coefficient was 315 W m –1 K–1 and for the speed of 0.6
m min–1 the average heat-transfer coefficient was about
205 W m–2 K–1. The numerical calculations conducted for
the new average values of the heat-transfer coefficient
for the secondary cooling zone 'average confirmed the
above method was correct. Similar metallurgical lengths
of 15.7 m and 16.5 m were obtained for the speeds of 0.6
m min–1 and 0.8 m min –1, respectively.
At the next stage, the heat-transfer coefficient was
calculated using dependence 12 for each spray zone. The
calculations resulted in a few sets of heat-transfer coeffi-
cients. The values of the heat-transfer coefficients that
met the boundary conditions within the allowed percen-
tage change in the heat-transfer coefficients for the indi-
vidual spray zones were selected. The boundary condi-
tions were determined on the basis of the actual change
in the flows of the cooling water in the spray zones for a
selected speed.
6 CONCLUSION
The development of a versatile method for the opti-
mum determination of strand cooling parameters in the
steel continuous-casting process is an extremely complex
issue, requiring that many mutually contradictory criteria
be met. The quality requirements related mainly to the
structure of the solidified cast strand impose a clear
restriction to increasing the casting speed. However, the
continuous-casting-machine efficiency must be synchro-
nised with the other nodes of the production line. The
flexibility required in this case is related to the necessary
response to unexpected events that may change the pro-
duction rhythm of a steelmaking shop.
The cooling-program determination method in the
steel continuous-casting process is based on the verifi-
cation of the numerical models describing the strand
solidification process. The determination of, possibly, all
the properties of the cast steel as the function of the tem-
perature is a very important step at the stage of deter-
mining the model parameters for the conditions defined
as the standard. It mainly concerns the specific heat and
the solidification heat, the dynamic viscosity, the density
and the heat-transfer coefficient. Carrying out the
measurements of the listed properties is, unfortunately, a
very complex task, requiring the use of very specialised
equipment.
For the mathematical description of the steel con-
tinuous-casting process at any particular level, we need
to find the appropriate method of selecting the boundary
conditions. The determination of the boundary condi-
tions based on the process information is one of the fac-
tors that are crucial for the correctness of model calcu-
lations. Practically, this means that the amount of the
used cooling water needs to be converted into the suit-
able heat-transfer coefficient.
The task of determining the cooling programme in
the steel continuous-casting process is not a determi-
nistic task. The selection of the variant for practical
implementation depends largely on the experience of the
user of the specific machine. However, the criteria that
allow an assessment of the correctness of the assumed
solution are strictly defined. The most important are:
• the shell thickness under the mould,
• the strand metallurgical length,
• the strand surface temperature at the selected
measurement points,
• the critical stress at the temperature of the shell under
the mould,
• the strand structure – in particular, the thickness of
the chilled grain zone.
Acknowledgments
This research project was supported by the European
Regional Development Fund – the Operational Program-
me "Innovative Economy", as a project New Concept for
Selection of the Cooling Parameters in Continuous
Casting of Steel, POIG.01.03.01-12-009/09.
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912 Materiali in tehnologije / Materials and technology 49 (2015) 6, 903–912
I. JASTRZÊBSKA et al.: ELASTIC BEHAVIOUR OF MAGNESIA-CHROME REFRACTORIES ...
913–918
ELASTIC BEHAVIOUR OF MAGNESIA-CHROME REFRACTORIES
AT ELEVATED TEMPERATURES
ELASTI^NO VEDENJE OGNJEVZDR@NIH GRADIV
MAGNEZIJA-KROM PRI POVI[ANIH TEMPERATURAH
Ilona Jastrzêbska, Jacek Szczerba, Jakub Szlêzak, Edyta Œnie¿ek,
Zbigniew Pêdzich
AGH University of Science and Technology, al. A. Mickiewicza 30, 30-059 Kraków, Poland
ijastrz@agh.edu.pl
Prejem rokopisa – received: 2014-08-08; sprejem za objavo – accepted for publication: 2015-01-14
doi:10.17222/mit.2014.186
An investigation of the high-temperature elastic behaviour of four varieties of magnesia-chrome products was conducted in the
present work. The starting materials were characterized with XRF, XRD and SEM/EDS analyses. The measurements of the
dynamic Young’s modulus were performed in a temperature range of 20–1300 °C, using the acoustic method. Moreover, the
dynamic Young’s moduli, shear moduli and Poisson’s ratios were established at ambient temperature for the original and
after-heating samples. The obtained results showed similar characteristics, but the elastic behaviours of all the refractory test
materials were distinguishable. The curve of the Young’s modulus changed as a function of temperature, exhibiting a
close-to-hysteresis character, with a decreasing area between the lower and upper curve as the Cr2O3 content increased in the
material. The Young’s modulus and the shear modulus at ambient temperature were lower for the after-heating samples when
compared to the original ones.
Keywords: Young’s modulus, elastic properties, refractories, shear modulus, Poisson’s ratio
V tem delu je opisana preiskava visokotemperaturnih elasti~nih lastnosti magnezija-kromovih proizvodov. Za~etni materiali so
bili ocenjeni z analizo XRF, XRD in SEM/EDS. V obmo~ju temperatur 20–1300 °C so bile izvr{ene meritve dinami~nega
Young-ovega modula z akusti~no metodo. Nadalje so bili dolo~eni dinami~ni Young-ov modul, stri`ni modul in Poisson-ovo
{tevilo pri sobni temperaturi originalnih in toplotno obdelanih vzorcih. Dobljeni rezultati so pokazali podoben karakter, vendar
pa se je elasti~no vedenje vseh preizku{anih ognjevzdr`nih materialov razlikovalo. Krivulja Young-ovega modula se spreminja v
odvisnosti od temperature, ima ozek histerezni karakter z zmanj{anjem podro~ja med spodnjo in zgornjo krivuljo, ko vsebnost
Cr2O3 v materialu nara{~a. Young-ov stri`ni modul je bil pri sobni temperaturi ni`ji pri `arjenih vzorcih v primerjavi z original-
nimi.
Klju~ne besede: Young-ov modul, elasti~ne lastnosti, ognjevzdr`na gradiva, stri`ni modul, Poisson-ovo {tevilo
1 INTRODUCTION
The rapidly growing interest in the high-temperature
elastic properties arises from their great usefulness in the
area of designing of refractory materials and predicting
their lifetime. Notwithstanding, the literature about this
issue is still scarce. The Young’s modulus (E), as one of
the most relevant elastic-material constants, is directly
associated with the thermal-shock resistance, which is a
very important property during the first heating stage of
the refractory furnace lining, or when it is subjected to
permanent thermal stresses as in steel ladles, tundishes
or oxygen convertors used for pig-iron production in the
industrial metallurgy.
The commonly applied method for determining the
Young’s modulus includes the measurements of the
speed of the ultrasonic-wave propagation in a material.
This method allows us to measure the property at
ambient temperatures only. However, in real working
conditions the knowledge of the properties at elevated
temperatures is more valuable.
The magnesia-chrome refractories (MgO-Cr2O3) are
specific materials with multi-component complex micro-
structures composed of magnesia, introduced as a
clinker, and a solid solution of spinels like chromite,
magnesiochromite, magnesioferrite, a regular spinel or
hercynite, introduced with a chromite-ore concentrate.
Both components may also be introduced with a mag-
nesia-chromite co-clinker, which is a burning product of
these two components. This group of refractory materials
is no longer developed to be used for rotary-kiln linings
because of the toxicity of hexavalent chromium
compounds, but it is still very important in the
metallurgy of iron, lead or copper.1,2
Therefore, the purpose of this work is to assess the
elastic behaviour of a few magnesia-chrome refractory
materials with slightly different chemical compositions,
both at elevated and ambient temperatures.
2 EXPERIMENTAL WORK
The four varieties of magnesia-chrome (MCr) refrac-
tory products designated in order of increasing content of
Cr2O3 as MCr1, MCr2, MCr3 and MCr4, industrially
shaped and fired, were investigated. The test materials
Materiali in tehnologije / Materials and technology 49 (2015) 6, 913–918 913
UDK 666.76:620.17:620.187 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)913(2015)
were produced with the use of magnesia-chrome
co-clinker, fused-magnesia clinker and a chrome-ore
concentrate. As a result of different proportions between
the used raw materials, the products contained slightly
different amounts of Cr2O3 in the range of 18.0–21.8 %
and MgO in the range of 61.3–68.4 %. The CaO/SiO2
ratio for all of the materials was below 1.
Detailed chemical compositions determined with the
XRF method, together with the basic properties of the
examined products, are presented in Table 1. It can be
observed from this table that the test samples exhibit
slightly different open porosities which increase with the
decreasing ferric-oxide content. The average thermal-
shock resistance (TSR) of MCr1 is the same as for MCr4
(10 cycles) while for MCr2 it takes on the same value as
for MCr3 (6 cycles).
Table 1: Basic properties of the examined materials
Tabela 1: Osnovne lastnosti preiskovanih materialov
Property Material designation
MCr1 MCr2 MCr3 MCr4
Open porosity, % 14.0 13.8 12.1 15.6
Apparent density, g/cm3 3.11 3.17 3.23 3.07
Compressive strength, MPa 62.2 94.8 100.7 48.3
Bending strength at 20 °C,
MPa 6.8 9.7 8.6 7.2
Bending strength at
1450 °C, MPa 2.5 4.9 4.7 3.3
Average thermal-shock
resistance (TSR), cycles at
950 °C, water
10 6 6 10
Oxide Oxide content, w/%
MgO 68.4 66.4 61.3 64.2
Cr2O3 18.0 18.45 19.9 21.8
Fe2O3 7.0 8.1 10.2 6.7
Al2O3 4.0 4.4 5.4 4.2
CaO 0.88 0.82 1.07 0.85
SiO2 1.28 1.34 1.23 1.79
The XRD analysis of the investigated products was
performed at room temperature, using a PANanalytical
X’Pert Pro MPD X-ray diffractometer with Bragg-Bren-
tano geometry, and Cu-K radiation ( = 0.154056 nm),
in a range of 10 °  2
 < 90 °. The obtained XRD
patterns, illustrated in Figure 1, show two main compo-
nents which correspond to periclase (marked as P on the
diffractogram) and a complex solid solution of spinels
with an approximate chemical formula that can be
expressed as (Mg, Fe)(Cr, Al, Fe)2O4 (marked as S on the
diffractogram). The main difference between the
registered patterns is associated with different intensities
of the reflexes of the spinel solid solution, which gene-
rally increase with an increased content of Cr2O3 in the
material.
The microstructures of the starting materials were
observed under an ultra-high-definition scanning elec-
tron microscope, NovaNanosem200 equipped with an
energy dispersive spectrometer, EDS. The samples for
the BSE-SEM/EDS analysis were prepared as resin-em-
bedded cross-sections using the traditional ceramogra-
phic method. The exemplary microstructures of the mag-
nesia-chrome products are presented in Figures 2 to 5.
As it can be observed from the figures, the test products
do not differ significantly in their microstructures.
Periclase grains (visible in the SEM images as
dark-grey areas) contain inclusions of the (Mg, Mn,
Fe)(Al, Cr, Fe)2O4 solid solution (Figures 2 and 3 – point
1, Figure 5 – point 4). They represent magnesia
co-clinker used as the raw material. Pure grains of mag-
nesia clinker are also present in the material (Figure 2 –
914 Materiali in tehnologije / Materials and technology 49 (2015) 6, 913–918
I. JASTRZÊBSKA et al.: ELASTIC BEHAVIOUR OF MAGNESIA-CHROME REFRACTORIES ...
Figure 1: XRD patterns of test materials MCr1, MCr2, MCr3 and
MCr4
Slika 1: XRD-posnetki preizkusnih materialov MCr1, MCr2, MCr3 in
MCr4
Figure 2: SEM images of MCr1 product: 1 – MgO with inclusions of
solid solution (Mg, Fe)(Al, Cr, Fe)2O4, 2 – solid solution (Mg, Mn,
Fe)(Al, Cr, Fe)2O4, 3 – MgO
Slika 2: SEM-posnetka proizvoda MCr1: 1 – MgO z vklju~ki trdne
raztopine (Mg, Fe)(Al, Cr, Fe)2O4, 2 – trdna raztopina (Mg, Mn, Fe)
(Al, Cr, Fe)2O4, 3 – MgO
Materiali in tehnologije / Materials and technology 49 (2015) 6, 913–918 915
I. JASTRZÊBSKA et al.: ELASTIC BEHAVIOUR OF MAGNESIA-CHROME REFRACTORIES ...
Figure 4: SEM images of the MCr3 product along with the EDS analysis for points 1, 2 and 3: 1 – CaMgSiO4, 2 – (Mg, Fe)(Al, Cr, Fe)2O4, 3 –
MgO enriched with iron and chromium
Slika 4: SEM-posnetka produkta MCr3 z EDS-analizo v to~kah 1, 2 in 3: 1 – CaMgSiO4, 2 – (Mg, Fe)(Al, Cr, Fe)2O4, 3 – MgO, obogaten z
`elezom in kromom
Figure 3: SEM images of MCr2 product: 1 – MgO with inclusions of
solid solution (Mg, Fe)(Al, Cr, Fe)2O4, 2 – solid solution (Mg, Mn)
(Al, Cr)2O4
Slika 3: SEM-posnetka produkta MCr2: 1 – MgO z vklju~ki trdne
raztopine (Mg, Fe)(Al, Cr, Fe)2O4, 2 – trdna raztopina (Mg, Mn)(Al,
Cr)2O4
Figure 5: SEM images of MCr4 product: 1 – CaMgSiO4, 2 – MgO,
3 – solid solution (Mg, Mn, Fe)(Al, Cr, Fe)2O4, 4 – MgO with large
inclusions of solid solution (Mg, Mn, Fe)(Al, Cr, Fe)2O4
Slika 5: SEM-posnetka produkta MCr4: 1 – CaMgSiO4, 2 – MgO, 3 –
trdna raztopina (Mg, Mn, Fe)(Al, Cr, Fe)2O4, 4 – MgO z velikim
vklju~kom trdne raztopine (Mg, Mn, Fe)(Al, Cr, Fe)2O4
point 3, Figure 5 – point 2). The lightest areas comprise
a complex solid solution containing iron, (Mg, Mn,
Fe2+)(Al, Cr, Fe3+)2O4 (Figures 2 and 4 – point 2, Figure
5 – point 3), or without iron, (Mg, Mn) (Al, Cr)2O4
(Figure 3 – point 2). Calcium magnesium silicate
(monticellite) was also detected in the microstructures of
the MCr3 and MCr4 products, visible as the light-grey
areas located between the magnesia-chrome co-clinker
grains (Figures 4 and 5 – point 1).
The changes in the Young’s modulus, depending on
the temperature, were measured in the heating conditions
at the maximum temperature of 1300 °C using a reso-
nant-frequency and damping analyser RFDA HT-1600.
The test samples, of rectangular cross-sections, had
dimensions of 100 mm × 40 mm × 20 mm. The tempe-
rature increment during the heating was set to 2 °C/min
while for the cooling it was 5 °C/min. The holding time
at the maximum temperature was 20 min.
The ambient-temperature measurements of the dyna-
mic Young’s modulus, shear modulus (G) and Poisson’s
ratio (v) were carried out on both the original samples
and on the samples after the E-modulus measurements at
elevated temperatures. Two samples, with rectangular
cross-sections and the same dimensions, 100 mm × 40
mm × 20 mm, of each type of the material were sub-
jected to an investigation. Therefore, the average
modulus value for each sample was accounted for by the
average value of the two measurements. The test was
conducted with a RFTA Professional apparatus, in the
flexural and torsion modes of vibrations, in accordance
with the ASTM E 1876-09 standard.3
3 RESULTS
Figures 6 to 9 depict the dependence between the
Young’s modulus (E) and the temperature during the
process of heating and cooling the samples. At the first
sight, the sequences of the measured points for all of the
examined materials look similar, exhibiting a "dolphin-
like" shape. These measured-point sequences arrange-
ment also resembles a hysteresis function, with a de-
creasing area between the lower and upper points as the
Cr2O3 content increases in the material.
The common feature of all the investigated products
is an evident increase of the heating slope which begins
at about 600 °C. The MCr1 product, characterized by the
lowest Cr2O3 content and a high open porosity, exhibited
the widest hysteresis and the lowest value of the initial
Young’s modulus read off from the E(T) dependence as
12.7 GPa (Figure 6). During the heating of this sample,
916 Materiali in tehnologije / Materials and technology 49 (2015) 6, 913–918
I. JASTRZÊBSKA et al.: ELASTIC BEHAVIOUR OF MAGNESIA-CHROME REFRACTORIES ...
Figure 8: Young’s modulus E versus the temperature during heating
(triangles) and cooling (dots) for the material MCr3
Slika 8: Young-ov modul E v odvisnosti od temperature med ogreva-
njem (trikotniki) in ohlajanjem (to~ke) za material MCr3
Figure 6: Young’s modulus E versus the temperature during heating
(triangles) and cooling (dots) for the material MCr1
Slika 6: Young-ov modul E v odvisnosti od temperature med ogreva-
njem (trikotniki) in ohlajanjem (to~ke) za material MCr1
Figure 7: Young’s modulus E versus the temperature during heating
(triangles) and cooling (dots) for the material MCr2
Slika 7: Young-ov modul E v odvisnosti od temperature med ogreva-
njem (trikotniki) in ohlajanjem (to~ke) za material MCr2
Figure 9: Young’s modulus E versus the temperature during heating
(triangles) and cooling (dots) for the material MCr4
Slika 9: Young-ov modul E v odvisnosti od temperature med ogreva-
njem (trikotniki) in ohlajanjem (to~ke) za material MCr4
the maximum E value of 60.7 GPa was achieved at 1172
°C. Then, the elastic modulus decreased and reached the
minimum value of 50 GPa at 1252 °C. Such an increase
followed by a decrease in the Young’s modulus during
the heating cycles was observed for all the investigated
samples. The maximum E value of 80.7 GPa, during the
heating cycle, was registered for the MCr2 material at
1181 °C (Figure 7).
The cooling-cycle point sequence, for all the exa-
mined materials, showed an evident increase in the
elastic modulus and a following gradual decrease, finally
reaching an E value that was lower than the starting one,
registered at the beginning of the test. In spite of the fact
that the MCr1 and MCr4 products differ mostly in the
area between the heating and cooling slopes, their values
of the initial and final E moduli are the closest and so are
their elastic moduli at the maximum temperature. On the
other hand, the MCr2 and MCr3 materials, exhibiting
similar sequences of the measured points and the areas
between them, achieve similar E-modulus values both at
the ambient and maximum temperatures of the test. It is
worth to note that these two "pairs" of the materials
exhibit the same thermal-shock resistance that can be
observed from Table 1 (TSR: MCr2 = MCr3 = 6 and
MCr1 = MCr4 = 10).
Table 2: Results of Young’s modulus (E), shear modulus (G) and
Poisson’s (v) ratio measured at ambient temperature
Tabela 2: Young-ov modul (E), stri`ni modul (G) in Poisson-ovo
{tevilo (v), izmerjeno pri sobni temperaturi
Sample
designation E/GPa
E decrease
after heat-
ing, %
G/GPa
G decrease
after heat-
ing, %
v
MCr1o 13.80 ±0.14
16.1
7.28 ±
0.07
17.4
–0.052
MCr1h 11.58 ±0.12
6.01 ±
0.06 –0.038
MCr2o 21.51 ±0.22
20.9
10.43 ±
0.10
18.5
0.031
MCr2h 17.02 ±0.17
8.50 ±
0.09 0.001
MCr3o 24.41 ±0.24
22.4
12.29 ±
0.12
22.8
–0.007
MCr3h 18.93 ±0.19
9.49 ±
0.09 –0.003
MCr4o 17.62 ±0.18
20.8
8.08 ±
0.08
20.4
0.091
MCr4h 13.96 ±0.14
6.43 ±
0.06 0.086
Designation indexes: o – original sample, h – after-heating sample
Table 2 presents the results of the Young’s moduli
and shear moduli determined with the standard deviation
as well as Poisson’s ratios calculated with the equations
given in ASTM E 1876-09 for the samples with a rec-
tangular geometry. It can be observed that in each case
the Young’s and shear moduli of the after-heating
samples are lower in comparison to the original ones.
Here, the "pairs" of the test materials with similar E and
G moduli can also be distinguished. The moduli for
MCr1 are the most similar to the ones for MCr4. A
similar situation occurs when comparing the MCr2 and
MCr3 materials.
It is worth emphasizing that the drops in the E values,
for all the materials, entail a decrease in the shear
moduli. The highest values of the E and G moduli for
both the original and after-heating samples were ob-
tained for the MCr3 sample, while the lowest values
were registered for MCr1. The results obtained for
Poisson’s ratio v showed that in each case its value was
lower for the original sample when compared to the
after-heating one. Moreover, for the MCr1 and MCr3
samples the v ratios even reached negative values, where-
as for the rest of the samples they were positive.
4 DISCUSSION
Firstly, it should be emphasized that magnesia-
chrome materials are highly complex systems because
they are composed of magnesia and a few spinel com-
pounds that often form solid solutions with different
proportions, making them even more complex. Among
these spinels the following can be found: chromite
FeCr2O4, magnesiochromite MgCr2O4, regular spinel
MgAl2O4, magnesioferrite MgFe2O4, galaxite MnAl2O4
and hercynite FeAl2O4.
The present mixture of the mentioned spinels was
found with the SEM observations and was also con-
firmed by the XRD analysis. It is worth noting that there
is a good repeatability of the achieved results with regard
to the curve shape created (Figures 6 to 9) by the
obtained point sequences, despite the fact that these were
industrial samples, in which an unavoidable spread of
commercial properties may occur.
All the measured point sequences resemble a hyste-
resis function, which arises from the thermal history of a
sample. The initial increasing character of an E(T) curve
is associated with a densification of the microstructure
due to a partial oxidation of Fe2+ into Fe3+, which starts
above 300 °C,1 leading to a modification of the existing
spinel solid solution due to an incorporation of oxygen
into the structure, or due to a grain reorganization as a
result of the stress during the test and the increased
temperature.
Above the temperature of 500 °C a more intense
growth of the heating slope probably originates from an
"order-disorder" phase transformation, occurring in the
solid mixtures with a spinel structure. This phenomenon
derives from the propensity of the ions to change their
local positions in the structure. The ions in the 2+ oxi-
dation states, which are regularly located in the tetra-
hedral (T) sites, pass to the octahedral (M) ones that are
normally occupied by the 3+ ions, and vice versa.4 It is
called a transition into an "inversion structure", which is
followed by the changes in the structural-parameter
values, like the oxygen parameter or the cation-anion
distance in the T and M sites.
Materiali in tehnologije / Materials and technology 49 (2015) 6, 913–918 917
I. JASTRZÊBSKA et al.: ELASTIC BEHAVIOUR OF MAGNESIA-CHROME REFRACTORIES ...
The literature quotes that cation disordering consi-
derably affects the elasticity,5 the compressibility and the
thermal-expansion coefficient.6–8 The beginning of an
inversion transformation corresponds to different tempe-
ratures, depending on the type of spinel. After reaching
the maximum value of the Young’s modulus, the curve
(the author has in mind a continuous sequence of
measured points) starts to drop until reaching the maxi-
mum temperature of the test, which needs a further
investigation. The cooling cycle after the maximum tem-
perature starts with an increase in the E modulus reach-
ing the maximum value for all the test materials. This
behaviour exhibits the "strengthening" of the material in
the average temperature range of 800–1000 °C, which
may be the effect of the recrystallization process and
filling the voids.
In the case of the MCr1 material the E-modulus
change exhibits an almost continuous character during
the cooling process, whereas in the cases of the other
samples this dependence is discontinuous, requiring a
further investigation. The following gradual drop in the
E(T) dependence is probably related with the stress rela-
xation due to a microcrack formation or, as previously
reported by Podwórny et al.9, a channel-like pore for-
mation, which triggers the loosening of the microstruc-
ture. Even though the test materials are not considerably
different in their chemical and phase compositions, they
exhibit distinguishable behaviours of the elastic-modulus
change. Therefore, the E(T) dependence may represent a
"finger print" of the material. It proves that this method
can be highly helpful for assessing the high-temperature
behaviour of refractory materials.
The obtained lower values of the E moduli at ambient
temperature for all the after-heating MCr samples (Table
2) are the results of microcracks or other defects formed
during the heating of the materials. The decreased
Young’s moduli (E) are accompanied by simultaneous
decreases in the shear moduli (G). These two moduli,
together with Poisson’s ratio (v), are expressed with a
well-known equation, v = (E/2G) – 1. It can be observed
that the MCr1 and MCr3 products exhibit negative
values of Poisson’s ratio for both the original and
after-heating samples. Such a behaviour of a material
was also previously reported.9,10 The negative values of
Poisson’s ratio show that the material unfolds when it is
subjected to stretching11 and the negative v obtained in
this research can be ascribed to the rearrangement and
rotation of the microstructural components such as large
grains, used for the production of the material, or the
pores that are always present in the microstructures of
refractory materials.
Notwithstanding, it is difficult to relate these ratios
with the elastic behaviour, chemical composition and
type of the material investigated in this study. The cal-
culated percentage changes in the E and G values after
the heating (during the test at elevated temperatures)
(Table 2) show that the tensile, compressive and shear
stresses have almost equally destructive effects on
materials MCr3 and MCr4. On the other hand, according
to the obtained results, the properties of the MCr1 mate-
rial are more influenced by the shear stress, while for the
MCr2 material the tensile and compressive stresses play
prevailing roles.
5 CONCLUSIONS
1. The high-temperature investigation of the elastic pro-
perties of magnesia-chrome refractories showed
hysteresis-like behaviours. The hysteresis range was
the widest for the MCr1 product with the lowest
amount of Cr2O3 and the narrowest for the MCr4
product with the largest amount of Cr2O3.
2. Different E moduli obtained at elevated temperatures
for the magnesia-chrome products prove that the used
method may be helpful for predicting the high-tem-
perature behaviour and lifetime of this kind of widely
applied refractory materials.
3. The dynamic Young’s moduli and shear moduli,
measured at ambient temperature, were found to be
lower for the after-heating samples when compared
to the original ones, which was the result of the
defects formed during the heating process.
Acknowledgements
This work was supported by grant no. UDA-POIG.
01.04.00-18-028/11-00.
6 REFERENCES
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ture and property relations, Journal of American Ceramic Society, 82
(1999) 12, 3279–3292, doi:10.1111/j.1151-2916.1999.tb02241.x
5 R. C. Liebermann, I. Jackson, A. Ringwood, Elasticity and phase
equilibria of spinel disproportionation reactions, Geophys. J. Int., 50
(1977) 3, 553–586, doi:10.1111/j.1365-246X.1977.tb01335.x
6 R. M. Hazen, H. Yang, Effects of cation substitution and order-dis-
order on P-V-T equations of state of cubic spinels, American Mine-
ralogist, 84 (1999), 1956–1960
7 J. Podwórny, Order-disorder phase transformation in 2:3 spinels,
Ceramika – Ceramics, vol. 117, Polish Ceramic Society, Kraków
2014 (in Polish)
8 F. Martignago, A. Dal Negro, S. Carbonin, How Cr3+ and Fe3+ affect
Mg-Al order-disorder transformations at high temperature in natural
spinels, Physics and Chemistry of Minerals, 3 (2003), 401–408,
doi:10.1007/s00269-003-0336-0
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tory materials with thermal shocks, Ceramics International, 37
(2011), 2221–2227, doi:10.1016/j.ceramint.2011.03.070
10 R. Lakes, Deformation mechanism in negative Poisson’s ratio mate-
rials: structural aspects, Journal of Materials Science, 26 (1991) 9,
2287–2292, doi:10.1007/BF01130170
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235 (1987) 4792, 1038–1040, doi:10.1126/science.235.4792.1038
918 Materiali in tehnologije / Materials and technology 49 (2015) 6, 913–918
I. JASTRZÊBSKA et al.: ELASTIC BEHAVIOUR OF MAGNESIA-CHROME REFRACTORIES ...
M. DOŒPIAL et al.: STUDY ON THE MAGNETIZATION-REVERSAL BEHAVIOR ...
919–923
STUDY ON THE MAGNETIZATION-REVERSAL BEHAVIOR OF
ANNEALED Sm-Fe-Co-Si-Cu RIBBONS
[TUDIJ VEDENJA PRI OBRATU MAGNETIZACIJE @ARJENIH
TRAKOV Sm-Fe-Co-Si-Cu
Marcin Doœpial, Sebastian Garus, Marcin Nabialek
Institute of Physics, Czestochowa University of Technology, Armii Krajowej Av. 19, 42-200 Czestochowa, Poland
mdospial@wp.pl
Prejem rokopisa – received: 2014-09-04; sprejem za objavo – accepted for publication: 2014-12-09
doi:10.17222/mit.2014.219
The paper presents the studies of Sm-Fe-Co-Si-Cu ribbons obtained with the melt-spinning technique, annealed at 1123 K for
3 h. The phase-composition studies were made using a D8 Advance Bruker X-ray diffractometer. It was found that the studied
alloy has a multi-phase composition. These studies were of crucial importance in the interpretation of the magnetization
reversal. Magnetic measurements, i.e., the major hysteresis loop and recoil curves were performed using a LakeShore
vibrating-sample magnetometer with the maximum magnetic field of up to 2 T. On the basis of the recoil curves, the hysteresis
loop was decomposed into the reversible and irreversible magnetization components. The decomposed curve was used to
describe the processes that influence the reversal magnetization in the studied permanent magnets. Further, these components
were used to model the recoil curves, using a modified hyperbolic T(x) model based on the method described by Doœpial. The
modeled hysteresis loop and recoil curves revealed a high compliance with the experimental data, proving the validity of the
assumptions made in the modeling procedure.
Keywords: permanent magnets, TbCu7 structure, magnetization reversal, hysteresis model
^lanek predstavlja {tudij trakov Sm-Fe-Co-Si-Cu, dobljenih s tehniko "melt-spining", 3 h `arjenih pri 1123 K. [tudij sestave faz
je bil izvr{en z rentgenskim difraktometrom D8 Advance Bruker. Ugotovljeno je bilo, da je preu~evana zlitina sestavljena iz ve~
faz. Te {tudije so bile klju~nega pomena pri razlagi obrata magnetizacije. Magnetne meritve, to so glavna histerezna zanka in
povratne krivulje, so bile izvr{ene z magnetometrom LakeShore z vibrirajo~im vzorcem z uporabo najve~jih magnetnih polj do
2 T. Na podlagi povratnih krivulj je bila histerezna zanka razdeljena v komponente reverzibilne in ireverzibilne magnetizacije.
Razstavljene krivulje so bile uporabljene za opis procesov, ki vplivajo na obrat magnetizacije pri preu~evanih permanentnih
magnetih. Nadalje so bile te komponente uporabljene za modeliranje povratnih krivulj z modificiranim hiperboli~nim modelom
T(x) na podlagi metode, ki jo je opisal Doœpial. Modelirane histerezne zanke in povratne krivulje so odkrile veliko ujemanje z
eksperimentalnimi podatki, kar potrjuje veljavnost pribli`kov, uporabljenih pri razvoju modela.
Klju~ne besede: permanentni magneti, strukture TbCu7, obrat magnetizacije, histerezni model
1 INTRODUCTION
Alloys with the chemical composition close to
SmCo7–8.5 are used for fabricating the materials with a
TbCu7 meta-stable structure that cannot exist steadily.1–3
Doping elements, such as Si, Zr, Cu, etc., promote the
crystallization of this type of disordered structure, whose
main feature is a positive  intrinsic-coercivity tempe-
rature coefficient.4,5 The annealing process applied to
these materials leads to the decomposition of the SmCo7
phase into more stable structures, composed of the
SmCo5 and Sm2Co17 phases.6–8
One of the most popular methods for determining the
reversal-magnetization process is an analysis of recoil
loops. Reversal magnetization in multiphase, nanocom-
posite permanent magnets, such as annealed
Sm12.5Fe8Co66.5Si2Cu11 ribbons, is very complex and often
described with more than one type of process.2 On the
basis of recoil curves, it is not only possible to determine
the number of reversible and irreversible magnetization
processes occurring in these materials, but also des-
cribing their parameters.6
Such information applied to hysteresis models can be
used for simulating the major and minor hysteresis
loops, the initial magnetization or recoil curves.
The aim of this paper is to present the theoretical
description of the major hysteresis loop and recoil curves
obtained using a hyperbolic T(x) model modified by
Doœpial and its comparison with the experimental data.
2 MATERIALS AND EXPERIMENTAL
PROCEDURE
Samples of the Sm12.5Fe8Co66.5Si2Cu11 alloy were
obtained from high-purity elements, by arc melting, in a
protective argon atmosphere. The studied ribbons were
prepared by rapidly quenching the liquid alloy on a
rotating copper wheel with a high linear velocity of
20 m/s. Both ingots and the samples were prepared in a
protective gas atmosphere under a pressure of 0.4 ×
105 Pa. The obtained ribbons were encapsulated in the
argon atmosphere, annealed at 1123 K for 3 h and slowly
cooled to room temperature.
Materiali in tehnologije / Materials and technology 49 (2015) 6, 919–923 919
UDK 621.318.1:537.624:669.018.582 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)919(2015)
XRD patterns were measured using a Bruker X-ray
diffractometer equipped with a Lynx Eye semiconductor
counter. Diffraction patterns were made using a Cu-K
radiation source with a characteristic wave length of
0.1541 nm in the Bragg-Brentano geometry. The samples
for the X-ray measurements were scanned in a 2
 range
from 30 ° to 120 ° with an angle step of 0.02 ° and an ex-
position time of 3 s. A quantitative and qualitative
analysis of the phase composition was carried out using
the Brass evaluation program applying the Rietveld
profile-matching method.9
The major hysteresis loop and recoil curve were
measured using a LakeShore VSM with the maximum
external magnetic field of 2 T. The method of the recoil-
curve decomposition into the constituent magnetization
components was described elsewhere.10,11 The samples
used for the magnetic measurements were in the form of
ribbons of known dimensions. The demagnetization field
resulting from their shape was taken into account and
evaluated with the method described in12.
2.1 Hysteresis model
In the original T(x) model,13,14 the hysteresis loop can
be described with the sum of sigmoid and linear func-
tions. The sigmoid hysteretic function characterizes the
irreversible magnetization changes. The linear one is
used for describing the reversible magnetization changes.
In this research the authors used the T(x) model modified
by Doœpial,15,16 describing the reversible magnetization
component with an anhysteretic sigmoid function. Based
on this assumption, the whole reversal magnetization
process was described with following Equations:
[ ]f M B C x a b br
i
n
i i i, , , , , ,(tanh ( )+
=
+= − + −∑hys R
irr
 0 0 1
1
0 0 1 i ) (1a)
[ ]
f M
B
C x a
r
j
j
n
j j
,
max
,
, ,tanh ( )
+
=
= ⋅
⋅
+ +
∑
anhys
rev
rev
 0
0
1
0 0 [ ]tanh ( ), ,C x aj j0 0
2
−⎛
⎝
⎜
⎜
⎞
⎠
⎟
⎟
(1b)
[ ] [ ]b C x a C x ai i k i i k i+ = + − −, , , , ,tanh ( ) tanh ( )0 0 0 0 (1c)
[ ]f M B C x a b br
i
n
i i i, , , , , ,(tanh ( )−
=
−= + + +∑hys R
irr
 0 0 1
1
0 0 1 i ) (2a)
[ ]
f M
B
C x a
r
j
j
n
j j
,
max
,
, ,tanh ( )
−
=
= ⋅
⋅
+ +
∑
anhys
rev
rev
 0
0
1
0 0 [ ]tanh ( ), ,C x aj j0 0
2
−⎛
⎝
⎜
⎜
⎞
⎠
⎟
⎟
(2b)
[ ] [ ]b C x a C x ai i k i i k i− = − − +, , , , ,tanh ( ) tanh ( )0 0 0 0 (2c)
[ ] [ ]
b
C x a C x a
i
i m i i m i
1
0 0 0 0
2,
, , , ,tanh ( ) tanh ( )=
+ − −
(3)
where (+) and (–) in the f±
hys, anhys
functions represent the
ascending and descending changes of the reversible
(anhys) and irreversible (hys) components, respectively; x
is the external-magnetic-field excitation, a0i is the center
of the ith pinning/nucleation site, a0i is the center of the
jth reversible process; B0,i, B0,j are the amplitudes of the
ith and jth magnetization components; C0,i, C0,j are the
sheering factors, while xm represents the maximum-
external-magnetic-field excitation. The i and j indexes
refer to individual reversible and irreversible magne-
tization components, respectively, and nirr,rev is their
total number.16
3 RESULTS AND DISCUSSION
Figure 1 presents the experimental X-ray diffraction
pattern compared with the results of the Rietveld refine-
ment simulation, obtained for the annealed
Sm12.5Fe8Co66.5Si2Cu11 thin ribbons.
M. DOŒPIAL et al.: STUDY ON THE MAGNETIZATION-REVERSAL BEHAVIOR ...
920 Materiali in tehnologije / Materials and technology 49 (2015) 6, 919–923
Figure 1: X-ray diffraction patterns: measured and calculated from
the Rietveld refinement and the difference curve for a
Sm12.5Fe8Co66.5Si2Cu11 ribbon annealed at 850 °C for 3 h
Slika 1: Rentgenogram, izmerjen in izra~unan iz Rietveld-ovega
pribli`ka, ter diferen~na krivulja za trak Sm12,5Fe8Co66,5Si2Cu11,
`arjen 3 h na 850 °C
Figure 2: Measured demagnetization curve and the calculated
hysteresis loop, obtained with the modified T(x) model, for the
Sm12.5Fe8Co66.5Si2Cu11 ribbon annealed at 850 °C for 3 h
Slika 2: Izmerjena krivulja razmagnetenja, izra~unana z modificira-
nim modelom T(x) histerezne zanke za trak Sm12,5Fe8Co66,5Si2Cu11,
`arjen 3 h na 850 °C
According to the Rietveld refinement, it was found
that the studied alloy was composed from Sm2Co17
(22.92 %), SmCo7 (37.45 %) and SmCo5 (39.63 %)
phases. A lack of one of the less intense peaks on the
experimental diffraction pattern, as compared with the
simulation, can be associated with the preferred position,
occupied by the Cu atoms in the TbCu7 structure and the
use of a copper X-ray source. Due to the overlapping
peaks, originating from the presence of different crystal-
line phases, it was not possible to estimate the average
grain size using the Bragg equation. However, it was
possible to state that it was less than about 120 nm.
With the measured demagnetization curve and the
simulated one, obtained with the modified hyperbolic
T(x) model, the major hysteresis loop is presented in
Figure 2.
The experimentally determined hysteresis loop was
used to estimate the basic magnetization parameters:
saturation of the magnetization μ0MS (0.90 T), rema-
nence μ0MR (0.70 T) and coercivity JHC (0.41 T).
The saturation of the magnetization and the rema-
nence were also compared with those calculated from the
initial magnetization curve and the irreversible magne-
tization dependence after the extrapolation to the infinite
field, using the method described elsewhere.17 The
calculated parameters were as follows: μ0MS() = 0.91 T
and μ0MR() = 0.70 T; they were used to obtain the
Mr()/Ms() ratio which was 0.77.
On the basis of the Mr()/Ms() ratio combined with
the shape of the demagnetization curve and the estimated
grain size, it was possible to conclude which type of in-
teractions between the particles is dominant. According
to the literature,18,19 a multiphase material with a smooth
demagnetization curve (as observed on Figure 2) and the
value of the MR/MS ratio higher than 0.5 is characteristic
for exchange-coupled nanocomposites and/or anisotropic
permanent magnets.
The measured demagnetization curve (the lower arm
of the hysteresis is symmetric) was compared with the
theoretically simulated hysteresis loop (Figure 2). The
simulation was done using the Mathematica software and
Equations (1) to (3). The obtained results showed a high
compliance with the experiment.
The startup data for simulating the major hysteresis
loop and recoil curve was determined on the basis of the
data obtained from the analysis of the reversible and
irreversible magnetic susceptibilities (Figure 3).
As it can be seen on Figure 3, both reversible and
irreversible susceptibilities are composed of at least three
distribution sites. The fitting parameters determined from
susceptibility curves are gathered in Table 1.
In order to simulate the irreversible magnetization
changes, it was necessary to use a combination of three
hysteretic functions, properly representing three pinning/
nucleation sites. On the other hand, the reversible
magnetization was represented by a combination of three
anhysteretic functions, sourcing from the rotation of
magnetization vectors, the free domain-wall movement
of unpinned domain walls or the bowing of strongly
pinned domain walls.
Table 1: Fitting parameters used in the simulation of the major
hysteresis loop and recoil curve applying the modified T(x) model,
where: a0i – the center of the ith pinning/nucleation site (irreversible)
or the peak resulting from the reversible process, B0i – the amplitude
of the ith magnetization component, C0i – the shearing factor of the ith
magnetization component
Tabela 1: Parametri ujemanja, uporabljeni pri simulaciji glavne histe-
rezne zanke in povratne krivulje z uporabo modificiranega modela
T(x), kjer je: a0i – sredina ith nukleacije (ireverzibilno) ali vrh pri re-
verzibilnem procesu, B0i – amplituda ith magnetizacijske komponente,
C0i – stri`ni faktor ith magnetizacijske komponente
Component C0i a0i B0i
rev
1 2.705 0.001 0.079
2 0.107 0.320 0.354
3 0.114 1.200 0.204
irr
1 0.039 0.022 0.355
2 5.618 0.413 0.589
3 2.672 0.749 0.081
M. DOŒPIAL et al.: STUDY ON THE MAGNETIZATION-REVERSAL BEHAVIOR ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 919–923 921
Figure 3: Reversible, irreversible and total susceptibilities determined
from the magnetization components for a Sm12.5Fe8Co66.5Si2Cu11
ribbon annealed at 850 °C for 3 h
Slika 3: Reverzibilna, ireverzibilna in skupna ob~utljivost, dolo~ena iz
komponent magnetizacije za trak Sm12,5Fe8Co66,5Si2Cu11, `arjen 3 h
na 850 °C
Figure 4: Decomposition of the hysteresis loop simulated with the
modified T(x) model into the hysteretic and anhysteretic functions
Slika 4: Razstavljanje z modificiranim modelom T(x) simulirane
histerezne zanke v histerezno in antihisterezno funkcijo
The decomposition of the hysteresis loop into the
hysteretic and anhysteretic curves, representing irrever-
sible and reversible magnetization processes, respec-
tively, is presented in Figure 4.
The obtained results were also used for simulating
the recoil curve in the demagnetization direction by
applying Equations (1) to (3). The comparison of the
simulated and measured recoil loops is presented in
Figure 5.
As can be seen in Figure 5, the simulated and
measured curves are similar. The observed small diffe-
rences between the experiment and the simulation can be
related to the recoil-loop openness effect that is, in turn,
associated with the magnetic viscosity.
4 CONCLUSION
From the X-ray diffraction it was found that the
studied sample was composed of three different phases,
i.e., Sm2Co17 (22.92 %), SmCo7 (37.45 %) and SmCo5.
The analysis of the parameters determined from the
hysteresis loop and its shape revealed that the Mr/Ms ratio
was higher than 0.5 (0.77). Such an increase in the value
of the aforementioned ratio is typically met in the sam-
ples characterized by a high anisotropy20 or strong
exchange coupling between nanosized grains.18,19 The
multiphase composition combined with a smooth, sin-
gle-step demagnetization curve can be treated as a proof
of strong exchange coupling between the grains of the
constituent phases.21
The decomposition of the demagnetization curve into
the reversible and irreversible magnetization components
provided information on the quantity and type of the
reversal-magnetization processes occurring in the studied
material. Furthermore, using the obtained results and the
T(x) model modified by Dospial, it was possible to
determine the anhysteretic and hysteretic curves forming
the hysteresis. The same data was used to simulate the
recoil curve that showed a high compliance with the
experimentally obtained one.
Acknowledgement
This work was supported by the Ministry of Science
and Higher Education of Poland through Grant No. N
N507 234940.
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Materiali in tehnologije / Materials and technology 49 (2015) 6, 919–923 923

U. T. KALYONCUOGLU et al.: EVALUATION OF THE CHITOSAN-COATING EFFECTIVENESS ...
925–931
EVALUATION OF THE CHITOSAN-COATING EFFECTIVENESS
ON A DENTAL TITANIUM ALLOY IN TERMS OF MICROBIAL
AND FIBROBLASTIC ATTACHMENT AND THE EFFECT OF
AGING
OCENA U^INKOVITOSTI NANOSA HITOZANA NA
OPRIJEMANJE MIKROBOV IN FIBROBLASTOV NA DENTALNI
TITANOVI ZLITINI TER NA POJAV STARANJA
Ulku Tugba Kalyoncuoglu1, Bengi Yilmaz2, Serap Gungor3, Zafer Evis4,
Pembegul Uyar5,6, Gulcin Akca7, Gulay Kansu8
1Balgat Oral and Dental Health Center, Baris Manco Street 22, 06520 Ankara, Turkey
2Middle East Technical University, Department of Biomedical Engineering, 06800 Ankara, Turkey
3Yuzuncu Yil University, Department of Mechanical Engineering, 65080 Van, Turkey
4Middle East Technical University, Department of Engineering Sciences, 06800 Ankara, Turkey
5Selcuk University, Department of Biology, 4207 Konya, Turkey
6Selcuk University, Advanced Technology Research and Application Center, 4207 Konya, Turkey
7Gazi University, Faculty of Dentistry, Department of Microbiology, 06500 Ankara, Turkey
8Ankara University, Faculty of Dentistry, Department of Prosthodontics, 06500 Ankara, Turkey
ulkutugbaterzi@gmail.com
Prejem rokopisa – received: 2014-09-23; sprejem za objavo – accepted for publication: 2015-01-08
doi:10.17222/mit.2014.239
The aim of this study was to obtain a biocompatible and antimicrobial implant surface by coating Ti6Al4V with chitosan which
can be used to create a smooth transmucosal region for a faster and better wound healing and an increased bioactivity. Ti6Al4V
plates were first abraded and ultrasonically cleaned and then coated with chitosan. In order to simulate the conditions of an oral
environment, a group of coated plates were treated in a thermocycle apparatus. The coatings were evaluated with SEM, EDS,
XRD and FTIR spectroscopy. The fibroblastic cell behavior was determined using HGF-1 cells. P. gingivalis was used to assess
the effectiveness of chitosan as an antimicrobial coating.
It can be said that the Ti6Al4V plates were successfully coated with chitosan, indicated by the presence of the C, H and O
elements in the EDS results. There were no significant differences between the XRD patterns of the coated and uncoated plates;
however, the characteristic bands of chitosan were observed in the FTIR patterns of both the coated and aged samples. The
fibroblast-cell attachment and proliferation were enhanced while the bacterial proliferation was inhibited by the chitosan
coating. Chitosan was shown to be a biologically useful material that can be used as the coating material for transmucosal
regions of dental implants.
Keywords: chitosan coating, dental implants, Ti6Al4V, HGF-1, P. gingivalis
Namen te {tudije je bil dobiti biokompatibilno in antimikrobno povr{ino implantata z nanosom hitozana na Ti6Al4V, ki je
primeren za gladko transmukozno podro~je, za hitrej{e in bolj{e celjenje ran ter pove~ano bioaktivnost. Pri plo{~ah Ti6Al4V je
bila najprej pove~ana hrapavost, nato so bile o~i{~ene z ultrazvokom, potem pa je bil nanesen hitozan. Da bi simulirali razmere
v ustih, je bila skupina plo{~ obdelana v napravi za termocikliranje. Nanosi so bili ocenjeni s SEM-, EDS-, XRD- in
FTIR-spektroskopijo. Vedenje celic fibroblastov je bilo dolo~eno z uporabo celic HGF-1. P. gingivalis je bil uporabljen za oceno
u~inkovitosti hitozana kot protimikrobnega nanosa.
Lahko trdimo, da so bile plo{~e Ti6Al4V uspe{no prekrite s hitozanom, kar potrjuje prisotnost elementov C, H in O v
EDS-rezultatih. Ni bilo opa`ene ve~je razlike pri rentgenski analizi vzorcev z nanosom in brez nanosa, vendar so bili opa`eni
karakteristi~ni signali hitozana pri FTIR-analizi vzorcev tako pri vzorcih z nanosom kot tudi pri staranih vzorcih. Oprijemanje
in {irjenje fibroblasti~nih celic je bilo pospe{eno, medtem ko je nanos hitozana zaviral {irjenje bakterij. Pokazalo se je, da je
hitozan biolo{ko koristen material, ki ga je mogo~e uporabiti za nanos na transmukozna podro~ja dentalnih implantatov.
Klju~ne besede: nanos hitozana, dentalni implantati, Ti6Al4V, HGF-1, P. gingivalis
1 INTRODUCTION
Peri-implantitis, defined as an inflammatory reaction
of the tissues surrounding a dental implant to a loss of
the supporting bone,1 is still the major challenge for the
implant dentistry. It is known that the interface between
an implant and the healthy soft tissue is similar to the
one involving natural teeth.2 A peri-implant soft-tissue
cuff has to provide the same functions with the peri-
dental gingiva, such as inflammatory and immunological
defenses, growth factors and cytokine productions, filter-
ing the seal around a tooth, an implant or prosthetic.3
Titanium (Ti), which is currently used for the
production of dental implants, has remarkable properties,
such as a good corrosion resistance, a very good bio-
vcompatibility and a high strength-to-weight ratio. A
dioxide (or trioxide) layer that is a few nanometers thick
(3–5 nm) spontaneously forms on the Ti surface. This is
the layer that the biological fluids and tissues are in
Materiali in tehnologije / Materials and technology 49 (2015) 6, 925–931 925
UDK 577:669.295:621.793/.795:532.6 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)925(2015)
contact with when Ti implants are inserted into the
body.4 However, Ti is not able to function as an anti-mi-
crobial agent and, therefore, it cannot prevent the peri-
implant infections.5
Surface modifications of implants with bioactive
coatings stimulating the bone-cell attachment and growth
is one way to increase the osseointegration and help sta-
bilize the implant.6 Furthermore, in recent years, in order
to decrease the bacterial activity around the Ti implants,
studies focused on coating the transmucosal components
with antimicrobial and biocompatible coating materials
which are also efficient at wound healing and cellular
attachment. On the other hand, the adhesion of gingival
and epithelial cells is a desirable situation, providing a
seal around the transmucosal component; however, a
bacterial adhesion, which can provoke a breakdown of
the attachment, is not wanted.4
For these kinds of applications, chitosan has many
advantages, such as biocompatibility, antimicrobial effi-
ciency and cheapness. It is a chitin-derived natural poly-
mer produced by a deacetylation reaction. Chitin is
mainly found in the exoskeleton of crustaceans and also
in some fungi. Chitosan has been used in many biome-
dical applications including wound healing, skin graft-
ing, homeostasis, hemodialysis, drug delivery, preventing
dental plaque, calcium adsorption, etc.7
The objective of this study was to fabricate and cha-
racterize chitosan-coated Ti6Al4V plates. It is hypothe-
sized that a chitosan coating would help with the
human-gingival-fibroblast (HGF) attachment and antimi-
crobial effects. The influence of aging which simulates
an oral environment was investigated in vitro as well.
2 MATERIALS AND METHODS
Titanium-alloy (Ti6Al4V) plates with dimensions of
20 mm × 20 mm × 2 mm were used in this study. The
samples were divided into three groups: the control or
un-treated Ti6Al4V (n = 5) (group 1), the chitosan-
coated Ti6Al4V (n = 5) (group 2) and the chitosan-
coated and aged Ti6Al4V (n = 5) (group 3). Before the
coating process, the surfaces of the plates were air-
borne-particle abraded with 50 μm aluminum-oxide
particles for 10 s under 413.7 kPa (60 psi), with a dist-
ance from the tip of 0.5 mm, in a sandblaster (Heraeus
Kulzer Combilabor, CL FSG 3, Germany). The aim of
sandblasting was to increase the surface contact area and
to obtain a higher surface roughness. All the samples
were ultrasonically cleaned for 10 min in distilled water,
70 % ethanol, acetone, 70 % ethanol and distilled water
sequence, respectively.
2.1 Chitosan coating
The coating procedure was performed as described
by Yuan et al.8 Firstly, the plates were submerged in a
volume fraction water : ethanol 5 % : 95 % solution.
3-iso-cyanatopropyltriethoxysilane was added at a con-
centration of  = 2 % for 10 min at room temperature
while maintaining the pH at 4.5–5.5 with NaOH or
acetic acid. The samples were rinsed with ethanol and
treated at 110 °C for 10 min in a vacuum oven (Nüve EV
018, Turkey). The samples were submerged into a vol-
ume fraction  = 2 % gluteraldehyde solution at room
temperature overnight and rinsed with distilled water. A
mass fraction 1 % acetic-acid solution (aqueous) and
1 % chitosan (a medium molecular weight, Sigma Ald-
rich, Germany) in the acetic-acid solution were prepared.
These solutions were mixed and stirred for 10 min and
the resulting solution turned into a gel. The plates were
submerged in this chitosan/acetic-acid solution at 4 °C
overnight. The excess water was allowed to evaporate
over 7 d (at ambient conditions).
The plates were briefly rinsed in 0.005 M NaOH
followed by distilled water. After the coating process, a
group of the coated plates (group 3) were treated with a
thermocycle apparatus (Nüve, Turkey) to simulate a
one-year oral environment. The aging process was
performed for 1 week and the temperature of the water
bath was alternated between 5–55 °C.9
The surface morphologies of the plates were ob-
served using SEM (QUANTA 400F Field Emission
SEM) at a voltage of 20 kV and the chemical compo-
sition was determined with EDS using a 30 keV ion
beam. For the SEM observation the samples were coated
with 5 μm AuPd. The XRD structural analysis was per-
formed using an Ultima-IV X-ray diffractometer (Riga-
ku, Tokyo, Japan). The XRD was operated with a Cu-K
radiation (40 kV/40 mA) and spectra were collected in
the 2
 range of 10 ° to 80 ° with a scan speed of 2 ° min–1.
The surface functional groups of the coated samples
were determined with a FTIR spectroscope (Bruker IFS
66/s, Bruker Optics, Germany) in wavenumber regions
of 3800 cm–1 to 300 cm–1.
2.2 Cell culture and the MTT test
Before the biological experiments, for sterilization,
the samples were washed in 95 % ethanol and placed
under UV for 15 min, rinsed twice with sterile demine-
ralized water and PBS. The HGF-1 cells (ATCC,
CRL-2014) were donated by the Ankara University
Faculty of Veterinary Medicine. The cells were cultured
in Dulbecco’s Modified Eagle Medium (DMEM) supple-
mented with 10 % of fetal-calf serum (FCS), 100 U/mL
penicillin and 100 μg/mL streptomycin in a  = 5 % CO2
incubator at 37 °C (all from Biochrom Ltd, Cambridge,
UK). The proliferation of the cells was determined with
a MTT (3-(4.5-dimethylthiazol-2-yl)-2.5-diphenyl-tetra-
zolium salt) test assay (Sigma, St Louis, MO, USA) and
the morphology of the cells was examined with SEM.
HGF-1 cells were seeded on the Ti6Al4V plates,
placed onto 6-well plates, at a density of cells 4 × 105
mL–1. The cells were incubated in 5 % CO2 at 37 °C for
96 h. The Ti6Al4V plates were moved onto new 6-well
plates after 96 h of incubation, and fresh media were
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926 Materiali in tehnologije / Materials and technology 49 (2015) 6, 925–931
added. The media were then removed, a diluted MTT
(5 mg/mL) solution was added into the wells and the
incubation was continued in 5 % CO2 at 37 °C for 4 h.
After that, the incubation medium was removed and 400
μL of isopropanol with 0.04 N HCl was added to each
well in order to dissolve the resulting formazan crystals.
The absorbance of the formazan product at 570 nm and
690 nm was measured with a microplate reader (Biotek
Epoch, Germany). The experiment was repeated inde-
pendently in triplicate.
2.3 Cellular attachment and the morphology
The surfaces were analyzed using SEM in order to
determine the cellular attachment and morphology of the
cells. The HGF-1 cells cultured for 96 h on the Ti6Al4V
plates were washed twice with a 0.1 M sodium-cacody-
late buffer (pH 7.4), and the cells were fixed with 2.5 %
glutaraldehyde prepared in 0.1 M sodium cacodylate for
1 h at room temperature. The excess glutaraldehyde
solution was removed and the cells were rinsed twice in
sterile distilled water and kept at –80 °C overnight before
lyophilization. After the cells were dried to a critical
point, the samples were coated with AuPd. The fixed
cells on the discs were observed with SEM (Zeiss LS-10,
Germany). The SEM images were recorded at 500-times
and 5000-times magnifications.
2.4 Microbiological evaluation
The microbiological processes were done in the
Medical Microbiology Laboratory at the Gazi University,
the Faculty of Dentistry. In order to assess the microbial
inhibition, Porphyromonas gingivalis (ATCC 33277) was
used and cultured in Columbia Broth (Merck, Germany)
supplemented with vitamin K (1 μg/mL), hemin
(5 μg/mL) and 5 % sheep blood in an automated anaero-
bic chamber (Electrotek, United Kingdom) at 37 °C for
3–5 d with an atmosphere of 10 % H2, 10 % CO2 and
80 % N2. The optical density of bacterial inoculum was
adjusted to a 5 × 108 colony-forming unit per mL
(cfu/mL) using a spectrophotometer (BioTek ELx800,
USA) according to the turbidity of the McFarland stan-
dard.
The same bacterial inoculum was spread onto a
Schaedler agar media (Merck, Germany) supplemented
with vitamin K (1 μg/mL), hemin (5 μg/mL) and 5 %
sheep blood in the automated anaerobic chamber (Elec-
trotek, UK) and the Ti6Al4V samples, coated with chito-
san, were embedded into the infected agar plates with
chitosan-coated faces. Then, the plates were incubated at
37 °C in an atmosphere of 10 % H2, 10 % CO2 and 80 %
N2 in the anaerobic chamber for 3–5 d. During the
incubation period, the day after the incubation and every
following day, a loopful of a sample was tested for the
viability of the bacteria by culturing them on another
Schaedler agar media, separately in the anaerobic cham-
ber under the same conditions as mentioned before.
Then, the grown colonies of the bacteria were counted,
calculated as cfu/mL and evaluated for the viability of
the bacteria. All the samples were studied in triplicates in
the experiment.
3 RESULTS
The SEM images of the uncoated, chitosan-coated
and chitosan-coated and aged Ti6Al4V plates are given
in Figure 1. As expected, the uncoated Ti6Al4V plate
(Figures 1a and 1b) was observed to have a rougher
surface than the coated plates.
A chitosan aggregate was identified in the SEM
micrographs of the chitosan-coated and aged plate. The
chitosan was homogeneously distributed on the sand-
blasted Ti6Al4V surface and conserved to a degree after
the aging process. The chitosan aggregate can be seen
more clearly at a high magnification even on the aged
plate (Figure 1f). From Figures 1c and 1d, it can be
seen that the chitosan completely covered the Ti6Al4V
substrate before the aging process was applied. The
microcracks in the chitosan coating (Figure 1c) can be
explained with the drying procedure in the ambient
conditions. These cracks would have been prevented by
freeze-drying the samples after the coating process. The
coating was expected to be thicker than the chitosan
phase on the aged plate due to the possible erosion
U. T. KALYONCUOGLU et al.: EVALUATION OF THE CHITOSAN-COATING EFFECTIVENESS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 925–931 927
Figure 1: SEM images of Ti6Al4V plates: a) uncoated, 500-times, b)
uncoated, 5000-times, c) chitosan-coated, 500-times, d) chito-
san-coated, 5000-times, e) chitosan-coated and aged, 500-times, f)
chitosan coated and aged, 5000-times
Slika 1: SEM-posnetki Ti6Al4V plo{~: a) brez nanosa, pove~ava
500-kratna, b) brez nanosa, 5000-kratna, c) nanos hitozana, 500-krat-
na, d) nanos hitozana, 5000-kratna, e) nanos hitozana in starano, 500-
kratna, f) nanos hitozana in starano, 5000-kratna
caused by the aging process. When Figures 1c to 1f are
compared, it can be seen that the rough-surface morpho-
logy, completely covered with the chitosan coating,
reappears in the aged plates.
EDS was used to identify the elements and obtain
semi-quantitative compositional information from the
surfaces of the plates. The EDS results for the surfaces of
the uncoated, chitosan-coated and chitosan-coated and
aged Ti6Al4V plates are given in Figure 2.
The EDS analysis confirms the presence of the
chitosan coating with the detected C and O elements on
the Ti6Al4V plates even after the aging process. Without
chitosan, due to the chemical content of the substrate
material, the Ti and Al elements were detected together
with the O element. The EDS spectrum of the chitosan-
coated plate (Figure 2b) was observed to contain a low
amount of the Na element.
The XRD patterns of the surfaces of the uncoated,
chitosan-coated, chitosan-coated and aged Ti6Al4V
plates are given, with the XRD peak positions of stan-
dard Ti and chitosan (Figure 3).
On Figures 3a and 3c, the substrate plates have the
main standard Ti peaks at the 2
 values of 35.02 °, 38.44 °
and 40.23 °. The common XRD peaks at 25.42 ° and
42.22 ° were also attributed to the Ti6Al4V substrate.
There were no distinct peaks of chitosan detected in the
XRD patterns of both the coated or coated and aged
plates.
The FTIR spectra of the surfaces of the chitosan-
coated and chitosan-coated and aged Ti6Al4V plates are
given in Figure 4. The FTIR spectra of the chitosan-
coated Ti6Al4V plates show obvious differences after
the aging process. If the FTIR spectrum of the coated
and aged sample (Figure 4b) is compared with the one
that was only coated, it can be seen that there are two
additional bands at 2921 cm–1 and 2851 cm–1 which
belong to the -CH2- and -CH3 stretching vibrations, res-
pectively.10 In addition, the intensities of the FTIR bands
were decreased after the aging process. In the FTIR
spectra of the chitosan-coated and aged sample, the band
U. T. KALYONCUOGLU et al.: EVALUATION OF THE CHITOSAN-COATING EFFECTIVENESS ...
928 Materiali in tehnologije / Materials and technology 49 (2015) 6, 925–931
Figure 3: XRD patterns of: a) standard Ti (JCPDS # 01-1197), b) chi-
tosan (JCPDS # 039-1894), c) uncoated Ti6Al4V plate, d) chitosan-
coated Ti6Al4V plate, e) chitosan-coated and aged Ti6Al4V plate
Slika 3: Rentgenski posnetki vzorcev: a) Ti standard (JCPDS #
01-1197), b) hitozan (JCPDS # 039-1894), c) plo{~a Ti6Al4V brez
nanosa, d) plo{~a Ti6Al4V z nanosom hitozana, e) plo{~a Ti6Al4V z
nanosom hitozana in starana
Figure 4: FTIR patterns of Ti6Al4V plates: a) chitosan-coated, b) chi-
tosan-coated and aged
Slika 4: FTIR posnetka vzorcev Ti6Al4V-plo{~: a) z nanosom hito-
zana, b) z nanosom hitozana in starano
Figure 2: EDS spectra of Ti6Al4V plates: a) uncoated, b) chitosan-coated, c) chitosan-coated and aged
Slika 2: EDS-spektri Ti6Al4V plo{~: a) brez nanosa, b) z nanosom hitozana, c) z nanosom hitozana in starano
at 1372 cm–1 was assigned to -NHCO of amide and the
bands at 1069 cm–1 and 1022 cm–1 were ascribed to the
saccharide structure.11
The cellular viability of the HGF-1 cells of both the
coated and coated and aged Ti6Al4V plates was
evaluated using a MTT assay. The absorbances of the
formazan produced by metabolically active HGF-1 cells
on the experimental groups are given in Figure 5.
The cell morphology, cytoskeletal structure and
adhesion behavior of the HGF-1 cells were observed
after 96 h of incubation on the control (untreated),
chitosan-coated, and chitosan-coated and aged Ti6Al4V
plates. The SEM micrographs, at two different magnifi-
cations, for each group are given in Figure 6.
In order to observe the inhibition effect of chitosan
against the bacterial growth, the untreated Ti6Al4V plate
(control), the chitosan-coated Ti6Al4V plate and the chi-
tosan-coated Ti6Al4V plate after the ageing process
were used as substrates. The numbers of viable P.
gingivalis colonies proliferated on these substrates are
given in Figure 7.
As can be seen from Figure 7, no P. gingivalis
colonies were detected in the agar medium, on which the
chitosan-coated samples were placed. On the other hand,
U. T. KALYONCUOGLU et al.: EVALUATION OF THE CHITOSAN-COATING EFFECTIVENESS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 925–931 929
Figure 7: Number of viable P. gingivalis colonies (interpreted as lg10)
Slika 7: [tevilo `ivih kolonij P. gingivalis (prikazanih kot lg10)
Figure 6: SEM micrographs of cultured human gingival fibroblast cells after 96 h of incubation on the Ti6Al4V plates: a), b) control (untreated),
c), d) chitosan-coated, e), f) chitosan-coated and aged. Magnifications: 500-times (images on the left-hand side) and 5000-times (images on the
right-hand side)
Slika 6: SEM-posnetki kultiviranih ~love{kih gingivalnih fibroblasti~nih celic po 96 h inkubacije na Ti6Al4V-plo{~ah: a), b) kontrolni (brez
nanosa), c), d) z nanosom hitozana, e), f) z nanosom hitozana in starano. Pove~ave: 500-kratna (slike na levi strani) in 5000-kratna (slike na desni
strani)
Figure 5: Survival/population of HGF-1 cells after 96 h of incubation
on the control (untreated), chitosan-coated and chitosan-coated and
aged Ti6Al4V plates. Data were expressed as the mean values (MV)
of the ± standard deviation (SD) of three independent experiments.
Slika 5: Pre`ivela populacija celic HGF-1 po 96 h inkubacije na kon-
trolni plo{~i Ti6Al4V (brez nanosa), z nanosom hitozana in z nanosom
hitozana ter staranju. Podatki za tri neodvisne preizkuse so prikazani
kot srednja vrednost (MV) ± standardna deviacija (SD).
the bacterial growth was detected in the agar medium for
the chitosan-coated and aged Ti6Al4V plates; however,
the growth was much smaller than in the agar medium
for the untreated Ti6Al4V plates.
4 DISCUSSION
Following the definition of osseointegration as a
close contact between the bone and Ti and its alloys at
the bone level,12 the researchers focused on the surface-
modification methods and implant-body design.
The major purposes for modifying dental implant
surfaces are to positively modulate the host/implant
tissue responses, prevent a microorganism attachment, a
bone destruction, and also a failure of the implant, and
decrease the healing period for osseointegration.12 The
aim of this study was to obtain an antibacterial coating
for dental/craniofacial and orthopaedic implants and test
the ability of the coating to promote the fibroblastic-cell
growth before and after an aging process. For this pur-
pose, chitosan, which is a biocompatible, biodegradable,
antibacterial and inexpensive natural polymer with a
good film-forming ability13, was applied on the Ti6Al4V
substrate via silanization. The chitosan was successfully
deposited on the plates, forming a uniform and yellowish
transparent layer with some microcracks, possibly
formed due to the ambient drying process.
The SEM micrographs revealed the existence of a
chitosan layer on the plates even after the aging process.
In addition, sandblasting increased the surface roughness
and allowed a proper mechanical interlock with the coat-
ing. The elements belonging to the chitosan coating were
detected with the EDS method on both the coated, and
coated and aged plates. The presence of the oxygen (O)
element in the EDS measurements of the Ti6Al4V plates
(Figure 2a) can be attributed to the Al2O3 grains em-
bedded in the surface during sandblasting14 or the oxide
layer naturally forming on Ti and its alloys. The sodium
(Na) element, which was detected in the EDS spectra of
the chitosan-coated plate, was thought to originate from
the NaOH solution that was used to remove the
remaining acetic acid after the coating process.
The XRD pattern of the chitosan is characteristic of
an amorphous polymer15 and yields broad peaks. Conse-
quently, in the chitosan-coated plates there were no
distinctive differences when compared to the substrate
material. However, in the FTIR spectra of the chitosan-
coated Ti6Al4V plates, characteristic bands of chitosan
were observed and the aged sample showed obvious
differences. The FTIR spectra showed molecular
changes induced by the aging process. The main change
observed was a decrease in the amide group 1372 cm–1
and the formation of two -CH2- and -CH3 bands.
From the MTT absorbance values, it can be con-
cluded that the degree of cell proliferation for the
chitosan-coated Ti6Al4V group was higher than that of
the control and chitosan-coated and aged groups after
96 h. Although the chitosan coating induced a prolife-
ration of the HGF-1 cells, after the aging process their
vitality decreased significantly. The general shape and
growth pattern of the fibroblast cells can be seen directly
from the SEM micrographs (Figure 6). After 96 h of cul-
turing, the cells attached to all the different surfaces; the
cells on the control samples were more flattened than on
the other surfaces, with a broader contact area. In con-
trast, the cells on the chitosan-coated Ti6Al4V surface
had a polygonal morphology with extensions in multiple
directions.
In addition, the cells on the chitosan-coated and aged
surface showed a spindle-like and elongated morpho-
logy, but they did not completely adhere to the surface.
From the above we can conclude that the cells cultured
on the chitosan coating showed higher initial-adhesion
properties with an increased number of extremities of the
cell bodies when compared to the control group. Since
the thermal process applied during the aging induced
chemical modifications in the structure of chitosan, the
attachment and proliferation of the cells were then
different. However, one week of aging, which was a
simulation of one year of the oral environment, did not
erode the coating much and the cellular response to this
surface was not significantly affected.
In an in-vitro study, in which the initial attachment of
oral bacteria on Ti surfaces was investigated16, it was
shown that comparatively large amounts of P. gingivalis
and A. actinomycetemcomitans adhered to a Ti surface
even after polishing. These findings indicate that there is
a considerable risk of adhesion of periodontopathic bac-
teria on Ti implants. The adhesion of bacteria is gene-
rally influenced by the physicochemical properties of the
material surface, including the surface roughness,
hydrophobicity (surface wettability) and electrical charge
(zeta potential)17. Generally, rough surfaces allow a
greater bacterial adhesion than smooth surfaces.
For the Ti6Al4V implant material, the electrical
charge of the material’s surface influences the adhesion
capacity of bacteria. Bacteria are generally negatively
charged, as are the Ti6Al4V surfaces. In this study,
although having the same charge, the bacteria adhered to
the untreated Ti6Al4V control samples. This proves that
Ti and its alloys were unable to prevent a bacterial adhe-
sion in long-term oral applications.
As chitosan was used for coating the Ti6Al4V
material in this study, the electrically charged interaction
between chitosan and the bacteria is also another import-
ant factor for assessing the bacterial adhesion. Never-
theless, positively charged chitosan can react easily with
the negatively charged molecules and particles. There-
fore, an electrical attraction can be expected between the
positively charged chitosan-coated surface of the
Ti6Al4V and negatively charged P. gingivalis. However,
because of its different physicochemical properties,
chitosan has an antibacterial effect instead of allowing a
bacterial adhesion.
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930 Materiali in tehnologije / Materials and technology 49 (2015) 6, 925–931
An ionic interaction between the cations due to the
amino groups of chitosan and anionic parts of bacterial
cell walls such as phospholipids and carboxylic acids
was proposed as the mechanism for the antimicrobial
activity of chitosan.18
It was reported that the antimicrobial effect between
chitosan and bacteria was related to the following pro-
bable mechanisms: in the case of gram-positive bacteria,
chitosan on the surfaces of cell walls forms a polymeric
membrane which inhibits the food ingestion into the
cells. Therefore, the cells cannot get food. In the case of
gram-negative bacteria (such as E. coli), low-mole-
cular-weight chitosan can pass into the cells easier; there
it breaks the cell metabolism, forms flocculation and
kills the bacteria by changing their physiological activi-
ties.19 In this study, the antibacterial effect of chitosan
against P. gingivalis can be explained with the above
mechanisms. On the other hand, bacterial inhibition was
not provided on the samples that were exposed to the
aging process, which can be explained with the chemical
effects of the aging process.
In summary, chitosan was successfully applied to the
Ti alloy and the aging process did not significantly erode
the coating material. The chitosan coating allowed the
adhesion and proliferation of human gingival fibroblast
cells and it showed a high level of cytocompatibility
while preventing the growth of the P. gingivalis bacteria.
As determined with the FTIR studies, a one-week aging
process, simulating a one-year oral environment, altered
the chemical structure of the chitosan coating. The cell
attachment decreased slightly; however, the coating was
not able to perform its antibacterial activity after the
aging process even though it was still better than the
uncoated metal.
Acknowledgements
This research was supported by TUBITAK (the
Scientific and Technological Research Council of Tur-
key) through project 111S515. The authors would also
like to thank the Advanced Technology Research and
Application Center at the Selcuk University for pro-
viding the resources (the cell culture and SEM facilities)
necessary for the completion of this work.
5 REFERENCES
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U. T. KALYONCUOGLU et al.: EVALUATION OF THE CHITOSAN-COATING EFFECTIVENESS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 925–931 931

E. SKOHEK et al.: STRUCTURE AND PROPERTIES OF THE CARBURISED SURFACE LAYER ON 35CrSiMn5-5-4 ...
933–939
STRUCTURE AND PROPERTIES OF THE CARBURISED SURFACE
LAYER ON 35CrSiMn5-5-4 STEEL AFTER
NANOSTRUCTURIZATION TREATMENT
STRUKTURA IN LASTNOSTI NAOGLJI^ENE POVR[INE JEKLA
35CrSiMn5-5-4 PO NANOSTRUKTURNI OBDELAVI
Emilia Sko³ek, Krzysztof Wasiak, Wies³aw A. Œwi¹tnicki
Warsaw University of Technology, Faculty of Materials Science and Engineering, Wo³oska 141, 02-507 Warsaw, Poland
emilia.skolek@nanostal.eu
Prejem rokopisa – received: 2014-10-08; sprejem za objavo – accepted for publication: 2014-12-09
doi:10.17222/mit.2014.255
The aim of the paper was to investigate the structure and properties of the carburized surface layer of the 35CrSiMn5-5-4 steel
after the nanostructurisation with the austempering heat treatment. During vacuum carburizing the surface layer of the steel was
enriched with carbon above w = 0.6 %. Steel samples were subsequently austenitized, quenched at two different temperatures,
260 °C and 320 °C, and annealed at these temperatures for the time necessary for the completion of the bainitic transformation.
For comparison, one set of carbonized samples was subjected to the conventional heat treatment: martensitic quenching and low
tempering. The microstructural characterisation of the steel after different heat treatments was performed using scanning (SEM)
and transmission (TEM) electron microscopes. It was shown that both austempering treatments led to a carbide-free
nano-bainitic structure composed of nanometric ferrite plates separated by thin layers of retained austenite. The microhardness
and wear resistance of three kinds of steel samples were investigated: the two subjected to the austempering treatment at
different temperatures and the one subjected to the conventional treatment. It was shown that the nano-bainitic structure
containing an increased amount of retained austenite displays a higher wear resistance than the tempered martensite. The results
confirm that austempering can be a competitive method of thermal treatment, in comparison to the conventional heat treatment,
for the steels after the carburizing process.
Keywords: vacuum carburizing treatment, carburized surface layer, carbide-free bainite, austempering, wear resistance
Namen tega dela je preiskati strukturo in lastnosti naoglji~ene povr{ine jekla 35CrSiMn5-5-4 po nanostrukturiranju s toplotno
obdelavo austempranja. Med naoglji~enjem v vakuumu se je povr{ina jekla obogatila z ogljikom nad masnim dele`em w = 0,6
%. Vzorci jekla so bili segreti v avstenitno podro~je, ga{eni na dve razli~ni temperaturi: 260 °C in 320 °C, in zadr`ani na teh
temperaturah, potrebnih za popolno bainitno pretvorbo. Za primerjavo je bila serija naoglji~enih vzorcev toplotno obdelana po
navadni metodi: ga{enje v martenzit in popu{~ano pri nizki temperaturi. Karakterizacija mikrostrukture jekla po razli~nih
toplotnih obdelavah je bila izvr{ena z vrsti~nim (SEM) in presevnim (TEM) elektronskim mikroskopom. Izkazalo se je, da obe
toplotni obdelavi austempranja povzro~ita nastanek nanobainitne strukture brez karbidov in sestavljene iz nanometrskih feritnih
plo{~ic, lo~enih s tanko plastjo zaostalega avstenita. Pri treh razli~nih vzorcih je bila izmerjena mikrotrdota in odpornost proti
obrabi; pri dveh z austempranjem pri razli~nih temperaturah in enem z navadno toplotno obdelavo. Izkazalo se je, da ima
nanobainitna struktura s pove~anim dele`em zaostalega avstenita ve~jo odpornost proti obrabi kot pa popu{~eni martenzit.
Rezultati so potrdili, da je toplotna obdelava z austempranjem bolj{a metoda toplotne obdelave kot navadna toplotna obdelava
za jekla po postopku naoglji~enja.
Klju~ne besede: naoglji~enje v vakuumu, naoglji~ena plast, bainit brez karbidov, austempranje, odpornost proti obrabi
1 INTRODUCTION
Carbide-free bainite obtained in the bottom range of
bainitic transformation is characterised by exceptionally
high mechanical properties such as hardness and tensile
strength.1–5 Moreover, its ductility is kept due to a high
volume fraction of the retained austenite.5–7 Simultane-
ously, a lack of cementite precipitations as well as the
presence of a great amount of the retained austenite in
the form of thin layers placed between the bainitic plates,
result in a high fracture toughness.8,9 The presence of the
retained austenite may also increase the frictional wear
resistance.10,11 It was postulated10–12 that during the wear
tests austenite may transform into the strain-induced
martensite due to the stresses occurring during the
friction. This would increase the hardness in the contact
zone as well as the frictional wear resistance in
comparison to the samples containing tempered marten-
site after the conventional treatment.10–12
Initially, the carbide-free structure was produced in
the types of steel with specifically designed chemical
compositions. These contained the appropriate amounts
of carbon, silicon and manganese.13 However, recently an
attempt was made to obtain such a structure also in the
commercial types of steel.5,14,15 The main objective of the
present study was to produce a structure of carbide-free
bainite in the surface layer of the 35CrSiMn5-5-4 steel
after the carburising. The second objective was to
determine the effect of the isothermal-quenching
temperature on the structure and the surface properties of
the examined material.
Materiali in tehnologije / Materials and technology 49 (2015) 6, 933–939 933
UDK 621.78:621.785.37:539.538 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)933(2015)
2 EXPERIMENTAL PROCEDURE
2.1 Vacuum carburizing
The carburizing process was carried out in a
15.0VPT-4022/24N vacuum furnace at the Seco/War-
wick Company. The carburizing treatment was con-
ducted with the FineCarb® technology in accordance
with the patent.16 It was divided into ten subsequent
processes of saturation and diffusion. The times of parti-
cular saturation and diffusion processes were selected in
accordance with the simulations conducted with the
SimVaC Plus® programme. The carburising atmosphere
was a mixture of acetylene, ethylene and hydrogen. Due
to the possibility of a significant growth of austenite
grains, the pearlitisation process was carried out directly
after the carburisation process. Afterwards, the furnace
feed was heated up to 850 °C, held for 20 min and,
finally, slowly cooled to the room temperature at the
furnace-cooling rate.
2.2 Carbon-content measurements
The amount of carbon after the carburisation process
was determined on the cross-section of the layer, using
the Magellan Q8 Bruker optical emission spectrometer
with spark excitation. The study was conducted from the
front of the sample. 21 measurements were carried out at
0.1 mm – a surface layer of about 0.1 mm was removed
after each measurement to obtain the cross-section of the
carbon content. The measurements were performed in an
accredited laboratory, in accordance with the 3/CHEM
procedure, 3 ed., Sep 2010, with a precision of 0.02 %
(0.001 % for sulphur and phosphorus).
2.3 Heat treatment
In order to form a nano-bainitic microstructure, the
carburised samples of steel were austenitized for 30 min
at a temperature of 900 °C, then cooled to the tempe-
rature range of the bainitic transformation and, finally,
annealed at this temperature (Figure 1). The cooling
medium used for isothermal quenching was composed of
a liquid alloy of Sn-Ag heated up to the temperature of
the isothermal step (260 °C and 320 °C). The time of iso-
thermal quenching was selected in the way that ensured
the end of the bainitic transformation at the temperature
of isothermal annealing. The conventional heat treatment
consisting of the martensitic quenching from the tem-
perature of 900 °C and the subsequent tempering process
at a temperature of 200 °C for 1h was performed in order
to compare the properties of the layers obtained with two
treatments. In this case, the oil quenching (WODOL) in-
tended for structural steels was used as a cooling
medium.
2.4 Microstructural observations
The microstructures of the austempered samples
were examined with the use of light microscopy, SEM
Hitachi S-3500N scanning electron microscopy and the
TEM JEOL 1200 transmission electron microscope
working at 120 kV. The LM and SEM observations were
conducted just under the surface of the layer. At first, 50
μm of the examined material of the sample surface were
grinded. Subsequently, the surfaces were etched with the
Nital agent. The TEM observations were conducted at
the surface zone of the layers at a depth of about 50–100
μm from the surface and in the core of the samples. Thin
plates with a thickness of 200 μm were cut out of the
samples, grinded to a thickness of 100 μm and, finally,
electrolytically polished until perforation occurred.
Phase constituents were identified according to the ima-
ges of the electron diffraction analysis. The observations
were carried out both in the bright field (BF) and in the
dark field (DF), using the reflexions obtained from diffe-
rent phases.
The thickness of the ferrite plates and the austenite
layers, observed on the TEM images, was determined in
accordance with the following stereological Equation:17
d L=
2
π
(1)
where d stands for the real size of the element of the
microstructure (in the analysed case, the real thickness
of a plate) and L means the size of the microstructure
measured on the TEM image (in this case, the width
measured on the image). The plate widths (L) were
measured perpendicularly to the interphase boundaries.
The relative volume fraction of the phases was
calculated on the assumption that the volume fraction of
a given phase is equivalent to its area fraction observed
on the image. Therefore, the n numbers of the secants of
the l length were marked on the image of the micro-
structure. The volume fractions of phases VV were
calculated with the following formula:
E. SKOHEK et al.: STRUCTURE AND PROPERTIES OF THE CARBURISED SURFACE LAYER ON 35CrSiMn5-5-4 ...
934 Materiali in tehnologije / Materials and technology 49 (2015) 6, 933–939
Figure 1: a) Scheme of the austempering and b) quenching and low-
tempering treatment
Slika 1: Shematski prikaz toplotne obdelave: a) austempranje, b) ga-
{enje in popu{~anje pri ni`ji temperaturi
V
c
nlv
ik=
∑
(2)
where cik is the sum of the widths of all the intersec-
tions of the secant line l with a given phase, l – the
length of the secant line.
2.5 Microhardness measurements
The measurements of the microhardness on the
intersection of a layer were conducted with a LECO LM
248AT semi-automatic microhardness tester with the
AMH43 v1.81 software in accordance with the PN-EN
ISO 6507-1 norm.18 1.96 N and 9.81 N loads, indenting
the material for 10 s, were used in the investigation.
2.6 Wear-resistance measurements
The wear resistance measurements were carried out
with the block-on-ring method using a T-05 tester, in
conformity with the ASTM G77 norm.19 The examined
sample was of a block shape and a width of 6.35 mm.
The ring with an outer diameter of 34.99 mm was used
as a counter specimen. The ring used in the research was
made of the 100Cr6 bearing steel with a hardness of 62
HRC. The investigation lasted for 100 min and the
applied unit load was 200 N mm–2 and 400 N mm–2. The
rotation speed was 316 r/min (5.26 Hz) and the rubbing
speed was 0.25 m/s. The track of the wear radius was
17.5 mm. Lux 10 oil was used as the grease.
3 RESULTS AND DISCUSSION
The previous research conducted on the
35CrSiMn5-5-4 steel indicated that the formation of car-
bide-free bainite structure with nanometric or submicron
grain sizes can be achieved during isothermal quench-
ing.15 However, the amount of the retained austenite is
relatively low due to a relatively low carbon content is
this steel.15 In order to increase the carbon content in the
surface layer, the 35CrSiMn5-5-4 steel with the chemical
composition presented in Table 1 was submitted to the
vacuum carburizing treatment using the injection me-
thod. The increase in the carbon content facilitated the
formation of a nano-bainitic structure in the steel.
The analysis of the chemical composition of the
cross-section of the carburised layer of a steel sample
indicated that the carbon content at the surface is w =
0.64 %. This value remains constant up to a depth of 0.5
mm from the surface. At a depth of 2 mm the carbon
content is w = 0.39 % (Figure 2).
A fine-grained acicular structure was formed during
the isothermal quenching at the temperature of 260 °C,
in the carburised layer of steel. The bainitic ferrite plates
within this structure were either parallel to each other or
arranged at different angles to each other. They formed
groups or packages. An increase in the temperature of
the treatment resulted in the microstructure in which
small, unetched areas of the retained austenite occurred,
E. SKOHEK et al.: STRUCTURE AND PROPERTIES OF THE CARBURISED SURFACE LAYER ON 35CrSiMn5-5-4 ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 933–939 935
Figure 3: Microstructure of the carburized surface layer on
35CrSiMn5-5-4 steel after austempering at: a) 260 °C and b) 320 °C –
LM images
Slika 3: Mikrostruktura naoglji~ene povr{ine jekla 35CrSiMn5-5-4 po
austempranju na temperaturi: a) 260 °C in b) 320 °C (svetlobna mikro-
skopija)
Figure 2: Carbon-content profile for the carburized surface layer on
35CrSiMn5-5-4 steel
Slika 2: Profil vsebnosti ogljika v naoglji~eni plasti na jeklu
35CrSiMn5-5-4
Table 1: Chemical composition of the 35CrSiMn5-5-4 steel in mass fractions, w/%
Tabela 1: Kemijska sestava jekla 35CrSiMn5-5-4 v masnih dele`ih, w/%
C Cr Mn Si Ni Cu Al Mo W Fe
w/% 0.35 1.31 0.95 1.3 0.14 0.15 0.04 0.018 <0.03 balance
apart from the fine bainite plates arranged parallelly to
each other (Figures 3 and 4).
A detailed investigation, conducted with TEM, of the
microstructure of a steel sample treated at the tempera-
ture of 260 °C revealed the presence of martensite and
carbide-free nano-bainite. The martensite and bainite
were separated from each other with the layers of re-
tained austenite (Figure 5). The average measured width
of the ferritic and martensitic plates was (102 ± 7) nm,
whereas the width of the layers of the retained austenite
was (33 ± 3) nm. The increase in the temperature of the
E. SKOHEK et al.: STRUCTURE AND PROPERTIES OF THE CARBURISED SURFACE LAYER ON 35CrSiMn5-5-4 ...
936 Materiali in tehnologije / Materials and technology 49 (2015) 6, 933–939
Figure 6: Microstructure of the carburized surface layer on
35CrSiMn5-5-4 steel after austempering at 320 °C: a) TEM BF image,
b) TEM DF image of austenite reflection
Slika 6: Mikrostruktura naoglji~ene povr{ine jekla 35CrSiMn5-5-4 po
austempranju na 320 °C: a) TEM BF-posnetek, b) TEM DF-posnetek
odboja avstenita
Figure 5: a) Microstructure of the carburized surface layer on
35CrSiMn5-5-4 steel after austempering at 260 °C, b) TEM BF image,
c) TEM DF image of austenite reflection
Slika 5: a) Mikrostruktura naoglji~ene plasti na jeklu 35CrSiMn5-5-4
po austempranju na temperaturi 260 °C, b) TEM BF-posnetek, c)
TEM DF-posnetek odboja avstenita
Figure 4: Microstructure of the carburized surface layer on 35CrSiMn5-5-4 steel after austempering at: a) 260 °C and b) 320 °C – SEM images
Slika 4: Mikrostruktura naoglji~ene plasti jekla 35CrSiMn5-5-4 po austempranju na temperaturi: a) 260 °C in b) 320 °C (SEM-posnetka)
treatment to 320 °C allowed the formation of a structure
consisting of carbide-free bainite with the retained auste-
nite. The bainitic plates have the average width of 65 nm
± 4 nm, whereas the layers of the retained austenite are
26 nm ± 2 nm thick (Figure 6). Additionally, blocks of
austenite with an area not exceeding 1 μm2 were ob-
served (Figure 7).
The relative volume fraction of the retained austenite
in the surface layer, estimated on the basis of the TEM
images was (20.5 ± 3.5) % and (20.2 ± 3.5) % for the
samples isothermally quenched at 260 °C and 320 °C,
respectively. The microstructure obtained through iso-
thermal quenching at 320 °C was relatively homogenous
in terms of the grain size. However, areas with different
amounts of austenite were observed in the structure, as
well as some individual secondary carbides that did not
form clusters.
In the cores of the samples treated at 260 °C a mar-
tensitic-bainitic structure was formed. The presence of
martensite resulted from the Ms temperature of 307 °C
which is higher than the austempering temperature. The
presence of highly dense carbides in martensite laths
indicates that the tempering process occurred in the core
of a sample during the austempering (Figure 8). The
E. SKOHEK et al.: STRUCTURE AND PROPERTIES OF THE CARBURISED SURFACE LAYER ON 35CrSiMn5-5-4 ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 933–939 937
Figure 9: Microstructure of the core of the carburized steel after
austempering at 320 °C
Slika 9: Mikrostruktura jedra pri naoglji~enem jeklu po austempranju
na 320 °C
Figure 7: Block of retained austenite in the carburized surface layer
on 35CrSiMn5-5-4 steel after austempering at 320 °C
Slika 7: Kos zaostalega avstenita v naoglji~eni povr{ini jekla
35CrSiMn5-5-4 po austempranju na 320 °C
Figure 10: Austenite areas in the core of the carburized steel after
austempering at 320 °C: a) TEM BF image, b) TEM DF image of aus-
tenite reflection
Slika 10: Avstenitna podro~ja v jedru pri naoglji~enem jeklu po aus-
tempranju na 320 oC: a) TEM BF-posnetek, b) TEM DF-posnetek
odboja avstenita
Figure 8: Microstructure of the core of the carburized steel after
austempering at 260 °C
Slika 8: Mikrostruktura jedra pri naoglji~enem jeklu po austempranju
na 260 °C
average grain size was (152 ± 23) nm. During the aus-
tempering at 320 °C carbide-free bainite of a submicron
grain size 140 nm ± 10 nm was formed in the core
(Figure 9). However, a relatively large number of larger
ferrite grains of about 0.5 μm were also observed (Fig-
ure 10). In both cases, the relative austenite volume frac-
tion in the core, equal to (13 ± 3) % and (11.4 ± 2.3) %
for the austempering at 260 °C and 320 °C, respectively,
was significantly lower than in the carburised layer.
Austenite mainly occurred in three different forms: as
very thin layers (29 ± 4) nm in the case of the treatment
carried out at 260 °C; as thicker layers (46 ± 6) nm; and
small blocks (0.086 μm2–0.288 μm2) in the case of the
treatment conducted at 320 °C.
In the case of the austempering at 320 °C, which is
higher than the Ms temperature, various factors may
affect the structure of the core as well as the volume
fractions of certain phases. The previous studies con-
ducted on the 35CrSiMn5-5-4 steel revealed that the time
necessary for completing the process of the bainitic
transformation at 320 °C was slightly below 2 h.15
According to the literature,20,21 extending the time of iso-
thermal treatment over the critical value results in a
coalescence of ferritic grains and leads to the formation
of austenite in the form of blocks. The increase in the
grain sizes of particular phases may be also affected by
the temperature of the treatment.21 Simultaneously, the
relative volume fraction of the retained austenite has a
tendency to decrease with the increase in the time of the
treatment.20
The heat treatment described in the present study
lasted for 24 h. Therefore, the grain sizes of ferrite and
austenite in the core increased, whereas the relative
austenite volume fraction decreased during the austem-
pering in comparison with the previous treatment of this
type of steel conducted for 2 h.15
The hardness of the layers was measured at different
loads. In the case of a carburised, quenched and low-
tempered layer, the hardness was higher than the hard-
ness of the carburised steel submitted to isothermal
quenching at 260 °C of approximately 130 HV0.2 and 125
HV1 units. It was also higher than the hardness of the
steel after the treatment at 320 °C of about 225 HV0.2 and
205 HV1 units (Figure 11). This may be associated with
a high amount of the relatively soft retained austenite in
the nano-bainitic structure. The hardness does not
change on the cross-section of a layer in the case of the
steel subjected to the conventional treatment. After the
nanostructurisation process, the hardness of the steel
slightly increased when the distance from the surface
increased. With the increasing distance from the surface,
the grain size also increases which should result in a de-
crease in the hardness. However, both the carbon content
and the relative austenite volume fraction in the nano-
bainitic structure decreased; therefore, the hardness in
the core of the austempered samples is slightly higher.
The results of the investigations on the frictional
wear resistance with the applied loads of 200 MPa and
400 MPa are presented in Figure 12. The level of the
volume wear increases with an increase in the applied
load. The wear tests indicate that the layer that was car-
burised and then austempered is characterised by a signi-
ficantly higher frictional wear resistance in comparison
to a much harder carburised layer, quenched and tem-
pered in the conventional way. This difference is particu-
larly pronounced in the case of the layer austempered at
260 °C. Most likely, this is associated with the amount of
E. SKOHEK et al.: STRUCTURE AND PROPERTIES OF THE CARBURISED SURFACE LAYER ON 35CrSiMn5-5-4 ...
938 Materiali in tehnologije / Materials and technology 49 (2015) 6, 933–939
Figure 12: Volumetric wear of the carburized surface layer on
35CrSiMn5-5-4 steel after various heat treatments
Slika 12: Volumenska obraba naoglji~ene plasti jekla 35CrSiMn5-5-4
po razli~nih toplotnih obdelavah
Figure 11: Hardness: a) HV0.2 and b) HV1 profiles for the carburized
surface layer on 35CrSiMn5-5-4 steel after various heat treatments
Slika 11: Profili trdote: a) HV0,2 in b) HV1 v naoglji~eni plasti jekla
35CrSiMn5-5-4 po razli~nih toplotnih obdelavah
the retained austenite in the layer with a nano-bainitic
structure. Although the austenite is a relatively soft pha-
se, it may transform into martensite under the stresses
occurring during the friction due to the TRIP effect.9–12
As a result of this effect, the hardness in the contact zone
would increase significantly. This, in turn, may improve
the frictional-wear resistance in comparison with the
samples treated in the conventional way.10–12 However, in
order to prove this theory, a further study of the phase
composition performed with XRD is required.
4 CONCLUSIONS
The austempering process conducted on the car-
burised 35CrSiMn5-5-4 steel allowed the formation of a
nano-bainitic structure in the carburised layer as well as
the formation of a martensite or carbide-free bainite
structure with submicron grain sizes in the core of the
examined material.
The increase in the carbon content in the layer during
the carburizing favours the formation of the retained aus-
tenite, which may undergo a martensitic transformation
due to the stresses during the wear test. The formed
martensite may significantly increase the frictional-wear
resistance of the examined steel.
Low-temperature isothermal quenching may be a
new heat-treatment method applied on steel after carburi-
sation. It is an alternative to the conventional treatment
of quenching and low tempering. It seems beneficial in
terms of the properties of a layer such as the increase in
the frictional-wear resistance and lower quenching dis-
tortions.
Acknowledgements
The results presented in this paper were obtained
within the project "Production of nanocrystalline steels
using phase transformations" – NANOSTAL (contract
no. POIG 01.01.02-14-100/09 made with the Polish
Ministry of Science and Higher Education). The project
is co-financed by the European Union from the European
Regional Development Fund, within Operational
Programme Innovative Economy 2007-2013.
5 REFERENCES
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E. SKOHEK et al.: STRUCTURE AND PROPERTIES OF THE CARBURISED SURFACE LAYER ON 35CrSiMn5-5-4 ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 933–939 939

M. SARÝKAYA: OPTIMIZATION OF THE SURFACE ROUGHNESS BY APPLYING THE TAGUCHI TECHNIQUE ...
941–948
OPTIMIZATION OF THE SURFACE ROUGHNESS BY APPLYING
THE TAGUCHI TECHNIQUE FOR THE TURNING OF STAINLESS
STEEL UNDER COOLING CONDITIONS
UPORABA TAGUCHI-JEVE METODE ZA OPTIMIRANJE
HRAPAVOSTI POVR[INE PRI STRU@ENJU NERJAVNEGA JEKLA
Z OHLAJANJEM
Murat Sarýkaya
Department of Mechanical Engineering, Sinop University, 57030 Sinop, Turkey
msarikaya@sinop.edu.tr
Prejem rokopisa – received: 2014-11-14; sprejem za objavo – accepted for publication: 2015-01-16
doi:10.17222/mit.2014.282
This paper presents the optimization of the surface roughness using the Taguchi technique to assess the machinability of the
AISI 316Ti steel with PVD coated carbide inserts under different cooling conditions such as dry, conventional (wet) and
cryogenic cooling with liquid nitrogen (LN2). Based on the Taguchi L9 (33) orthogonal-array design, the machinability tests
were made utilizing a CNC lathe machine. Test parameters including the cutting speed, the cooling condition and the feed rate
were taken and then the surface roughness (Ra) was measured to obtain the machinability indicator. An analysis of variance was
performed to determine the importance of the input parameters for the surface roughness. The process parameters were
optimized by taking the Taguchi technique into consideration. The Taguchi signal-to-noise ratio was employed with the
smaller-the-better approach to obtain the best combination. On the basis of the first-order model, a mathematical model was
created using the regression analysis to predict the Ra model. The results indicate that the feed rate is the parameter with the
highest effect on the surface roughness and that the other parameters also have a statistical significance. In addition, cryogenic
cooling is an alternative method for increasing the surface quality of machined parts.
Keywords: AISI 316Ti, cryogenic cooling, machinability, optimization, surface roughness, Taguchi method
^lanek obravnava optimiranje hrapavosti povr{ine z uporabo Taguchi-jeve metode za oceno obdelovalnosti jekla AISI 316Ti s
karbidnimi vlo`ki s PVD-nanosom v razli~nih razmerah ohlajanja, kot je suho, navadno (mokro) in kriogensko hlajenje s
teko~im du{ikom (LN2). Preizkusi obdelovalnosti so bili izvr{eni s CNC-stru`nico na osnovi Taguchi-jevega ortogonalnega niza
L9 (33). Izbrani so bili parametri preizkusov, hitrost rezanja, razmere pri ohlajanju in hitrost podajanja, nato pa je bila izmerjena
hrapavost povr{ine (Ra) kot pokazatelj obdelovalnosti. Izvr{ene so bile analize variance, da bi ugotovili pomembnost vhodnih
parametrov na hrapavost povr{ine. Procesni parametri so bili optimirani z upo{tevanjem Taguchi-jeve tehnike. Uporabljeno je
bilo Taguchi-jevo razmerje signal – hrup s pribli`kom ~im manj{e tem bolj{e za doseganje najbolj{e kombinacije. Na osnovi
modela prvega reda je bil postavljen z uporabo regresijske analize matemati~ni model za napovedovanje Ra. Rezultati ka`ejo, da
je hitrost podajanja parameter z najve~jim u~inkom na hrapavost povr{ine, vsi drugi parametri imajo statisti~no zna~ilnost.
Dodatno je kriogensko ohlajanje alternativna metoda za pove~anje kvalitete povr{ine stru`enih delov.
Klju~ne besede: AISI 316Ti, kriogensko ohlajanje, obdelovalnost, optimizacija, hrapavost povr{ine, Taguchi-jeva metoda
1 INTRODUCTION
Stainless steels were developed to obtain a better
corrosion resistance compared to traditional carbon
steels and they allow us to work at higher temperatures.
There is a lot of stainless steel in the industry, but
austenitic and ferritic stainless steels are commonly used
in the manufacturing industry.1 As a type of the AISI 316
steel, austenitic stainless steel AISI 316Ti contains low
amounts of titanium (Ti), approximately 0.5 %. This
steel type has the advantage of enduring higher tempera-
tures for a longer time compared to the other stainless
steels.2 The physical and mechanical properties of the
AISI 316Ti steel are similar to those of the other types of
316, but the corrosion resistance of 316Ti is better than
those of the standard grades.2 In recent years, due to its
different properties, this steel has been extensively used
for certain applications such as boat and ship parts,
medical and chemical handling equipment, heat exchan-
gers, fastening tools, and in nuclear and construction
industries where a low thermal conductivity, good heat
resistance and corrosion resistance and a high strength
are required in the high-temperature working conditions.
However, the machining of this austenitic stainless steel
is very difficult since it contains a high amount of
strength-enhancing elements such as chromium, nickel
and molybdenum.1 One of the major problems is the heat
generation at the cutting region during the machining of
difficult-to-cut metals. The machining process requires
more energy, so high temperatures occur throughout the
deformation process and the friction at the tool-chip and
tool-workpiece interfaces.3 Recently, the machining
technology has been quickly improved to increase the
processing productivity and machining performance in
the cases of difficult-to-cut steels. An increase in the pro-
ductivity can be achieved by decreasing the temperature
Materiali in tehnologije / Materials and technology 49 (2015) 6, 941–948 941
UDK 621.7.01/.09:519.233:621.941 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)941(2015)
at the tool-chip and tool-workpiece interfaces thanks to
the cooling/lubrication methods. As the cutting velocity
and the feed rate increase during a machining process,
due to an improvement in the coating technology, the
cutting temperature increases as well. Thus, the use of
cooling/lubrication is necessary during the metal-cutting
operations. In recent years, certain cooling/lubrication
methods such as cryogenic cooling, solid coolants/lubri-
cants, wet cooling (traditional cooling), minimum-quan-
tity lubrication, high-pressure coolants, compressed
air/gases have been employed and these technologies
have considerably increased the machining productivity.4
However, the use of mineral- or syntactic-based cutting
fluids has led to certain problems like health risks and
environmental pollution.5,6 In order to eliminate all the
cutting fluids from the metal-cutting process, cryogenic
cooling or high-pressure cooling with compressed air
can be applied to protect the health and the environment.
Surface roughness is one of the most critical quality
indicators of the machined surfaces of engineering mate-
rials used for important applications and the producers
believe that it determines the degree of surface quality of
the manufactured parts.6 A low surface roughness ob-
tained from machining experiments contributes to some
properties of workpiece including fatigue strength, corro-
sion and wear resistance, friction, etc.6,7 Surface rough-
ness is affected by many parameters such as machined
material, depth of cut, cutting-tool material, cutting
speed, tool-nose radius, feed rate, coating type and
cooling/lubrication conditions. Modern industry aims at
producing high-quality parts, reducing the costs in a
short time. To manufacture a product with a desired
quality of the machining, the optimum process parame-
ters should be chosen. Therefore, recently, certain statis-
tical methods like the Taguchi technique, response-sur-
face methodology (RSM), desirability function analysis,
ANOVA and grey relational analysis (GRA) have been
implemented to optimize and analyze process para-
meters.8–12 In the engineering applications and academic
studies of experimental design, the Taguchi method is
very useful thanks to the orthogonal array that signifi-
cantly reduces the number of the tests and, in addition, it
attempts to eliminate the influence of uncontrollable
factors on the test results. The main purpose of the
Taguchi technique is to provide quality during the design
stage. In this way, the cost and the test time decrease in a
shorter period.12,13 Therefore, in this study, the Taguchi
method with the L9 orthogonal array was employed.
In some studies, the machinability of austenitic stain-
less steel was investigated by the researchers. For exam-
ple, Kayir et al.1 studied the effect of the tool geometry
and the cutting parameters on the surface roughness in
machining AISI 316Ti under dry cutting conditions.
Their results demonstrated that the main parameters
were the feed rate with a 73.97 % effect and the radius of
the edge with a 13.26 % effect on the surface roughness.
Xavior and Adithan14 explored the effects of cutting
fluids, cutting speed, depth of cut and feed rate on the
tool wear and surface roughness in the turning of the
AISI 304 austenitic stainless steel using a carbide tool. It
was seen that the most important parameter was the feed
rate having a 61.54 % effect on the surface roughness,
while the cutting speed had a 46.49 % effect on the tool
wear. Further, according to the ANOVA analysis, it was
found that the cutting fluid had a considerable effect on
both the surface roughness and the tool wear. Ciftci15
investigated the influence of the cutting speed and the
tool coating on the surface roughness and the cutting
force in the turning of the AISI 304 and AISI 316 auste-
nitic stainless steels under dry cutting conditions. It was
reported that the cutting speed considerably affected the
surface roughness. Korkut el al.16 determined the best
cutting parameters in the turning of the AISI 304 auste-
nitic stainless steel with cemented carbide inserts. Their
results showed that the surface roughness decreased with
the increasing cutting speed. Tekýner and Yeºýlyurt17
investigated the influences of the cutting parameters on
the basis of the process noise in the turning of the AISI
304 austenitic stainless steel. It was found that the
cutting speed of 165 m/min and the feed rate of 0.25
mm/r gave the best results.
The literature survey indicates that there are very few
studies dealing with the turning of the AISI 316 stainless
steel. When these studies are examined, it is seen that the
surface roughness has not been evaluated with respect to
different cutting conditions like dry, wet and cryogenic
cooling procedures used during the turning of the AISI
316Ti stainless steel. In the light of the above informa-
tion, this study can be summarized in three points:
Firstly, the influences of the cutting parameters on the
surface roughness in the turning of the AISI 316Ti stain-
less steel with a PVD coated carbide insert were inve-
stigated under dry, wet and cryogenic cooling conditions.
Secondly, a mathematical model was formed to estimate
the result of different levels of input parameters using a
regression analysis. In the next process, an analysis of
variance (ANOVA) was applied to determine the influ-
ences of the machining parameters. Lastly, the process
parameters were optimized using the Taguchi technique.
To achieve its goals, this paper employed a Taguchi L9
M. SARÝKAYA: OPTIMIZATION OF THE SURFACE ROUGHNESS BY APPLYING THE TAGUCHI TECHNIQUE ...
942 Materiali in tehnologije / Materials and technology 49 (2015) 6, 941–948
Figure 1: General flow diagram of the study
Slika 1: Prikaz poteka {tudije
(33) orthogonal array for planning the experiments. An
experimental design including three parameters (feed
rate, cutting speed and cooling condition) with three
levels was organized.
2 EXPERIMENTAL PROCEDURE
The workflow diagram of this study is illustrated in
Figure 1. It shows the sequence of the performed study.
Table 1: Chemical composition of the material in mass fractions, w/%
Tabela 1: Kemijska sestava materiala v masnih dele`ih, w/%
C Mn Si P S Cr Ni Mo Cu Ti
0.021 1.775 0.495 0.036 0.019 16.74 10.92 2.15 0.536 0.318
2.1 Material, machine tool, cutting tool and measure-
ment
The AISI 316Ti workpiece material was used in the
turning experiments and its chemical composition is
given in Table 1. Recently, because of its unique pro-
perties including good heat resistance and corrosion
resistance, a low thermal conductivity and a high
strength at higher temperatures, this material has been
used in many engineering operations involving boat and
ship parts, medical and chemical handling equipment,
heat exchangers, fastening tools, and in the nuclear and
construction industry. The dimensions of the test mate-
rial were Ø 60 mm × 200 mm. All the turning tests were
conducted using a Falco Fl-8 model (Taiwan) CNC lathe
machine with the maximum spindle speed of 4800 r/min
and a 15 kW drive motor. An assembly produced by
Sandvik including a PVD coated carbide insert of type
SNMG 12 04 08-QM and a PSBNR 2020K-12 tool
holder was utilized as the main tool arrangement with the
following tool geometry: a rake angle of –6 °, a clear-
ance angle of 0 °, the major cutting-edge angle of 75 °, a
cutting-edge inclination angle of –6 ° and a nose radius
of 0.8 mm. The same type of cutting insert was em-
ployed for each test parameter. In engineering applica-
tions, surface quality is one of the most important quality
indicators. For this reason, the average value of the
surface roughness (Ra) was measured using a TIME TR
100 profilometer tester. Before the measurements of the
surface roughness, the measuring device was calibrated
with a special calibration. Each surface was machined by
using a new cutting insert and after each test measure-
ments were carried out on the workpiece.
2.2 Cutting conditions and design of the experiments
The cutting speed (Vc), the feed rate (f) and the cool-
ing condition (C) were taken as the cutting parameters.
The values of the cutting parameters were chosen from
the plot experiments and the manufacturer’s handbook.
During the machining tests, a constant depth of cut (ap =
1.6 mm) was used; the other cutting parameters and their
levels are given in Table 2. In this study, on the basis of
the control factors and their levels from Table 2, the
Taguchi L9 orthogonal array (OA) from the Minitab soft-
ware was used, as shown in Table 3 indicating the design
of the experiments. It has nine rows and three columns.
The rows correspond to the number of the tests; the
columns correspond to the process parameters with three
levels. In this array, the first, second and third columns
represent the cutting speed, feed rate and cutting condi-
tion, respectively. The tests were conducted under diffe-
rent cutting conditions such as dry cutting, conventional
wet cooling (flood coolant) and cryogenic cooling inside
the tool with liquid nitrogen (LN2). For wet cooling, a
solution with boron oil and water (the ratio of boron
oil/water = 1/20) was prepared.
Table 2: Process parameters and their levels
Tabela 2: Procesni parametri in njihovi nivoji
Code Controlparameter Notation
Levels of factors
Level 1 Level 2 Level 3
A Coolingcondition C Dry Wet Cryogenic
B Feed rate f/(mm/r) 0.1 0.16 0.25
C CuttingSpeed
Vc/
(m/min) 90 126 176
Table 3: Experimental design
Tabela 3: Na~rt eksperimentov
Exp.
no.
Coded values Actual values
A B C C f/(mm/r) Vc/(m/min)
1 1 1 1 Dry 0.1 90
2 1 2 2 Dry 0.16 126
3 1 3 3 Dry 0.25 176
4 2 1 2 Wet 0.1 126
5 2 2 3 Wet 0.16 176
6 2 3 1 Wet 0.25 90
7 3 1 3 Cryogenic 0.1 176
8 3 2 1 Cryogenic 0.16 90
9 3 3 2 Cryogenic 0.25 126
For the cryogenic cooling, liquid nitrogen was deli-
vered directly from the liquid-nitrogen pressure tank to
the tool holder at a pressure of 1.5 bar as shown in
Figure 2. Three holes were drilled into the tool holder.
The diameter of the first hole on the tool holder was 6
mm and it provided a connection between the tool holder
M. SARÝKAYA: OPTIMIZATION OF THE SURFACE ROUGHNESS BY APPLYING THE TAGUCHI TECHNIQUE ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 941–948 943
Figure 2: Modified tool holder and adaptor
Slika 2: Prirejen nosilec orodja in adapter
and the liquid-nitrogen container with the help of a hose
and an adaptor. The liquid nitrogen accumulated inside
the tool holder was released to the environment as a gas
vapor with the help of the other two holes, taking the
heat from the insert. The diameter of the gas exit holes
was made to be 1.5 mm. The modified tool holder and
the connection adapter are seen in Figure 2, while Fig-
ure 3 shows the experimental set-up for cryogenic cool-
ing.
3 RESULTS AND DISCUSSION
3.1 Analysis of the experimental results
Surface roughness is one of the most important qual-
ity criteria for engineering materials. During the turning
operations, the surface roughness can be controlled with
the machining parameters. In this study, the surface
roughness was evaluated using 3D surface plots in the
graphs given in Figure 4. This figure shows that the sur-
face roughness increased significantly with the increas-
ing feed rate. The reason for this can be the fact that an
increase in the feed rate leads to a vibration and in-
creases the heat at the tool-chip interface; thereby a
higher surface roughness occurs.18 To calculate the
theoretical surface roughness, the abbreviated formula is
expressed as follows:
R
f
ra
=
⋅
2
32
(1)
According to Equation (1), in order to improve the
surface quality, the feed rate can be decreased or,
alternatively, the nose radius of the cutting insert can be
increased since the surface roughness is a function both
the nose radius and feed rate. The results obtained from
the experiments are similar to this formula. In the lite-
rature, it is pointed out that surface roughness is affected
negatively by an increase in feed rate and in order to
obtain the better surface quality, the feed rate is reduced
usually in machining proceses.6,8,13 In present work, a
similar result was detected when the surface roughness
decreased with the increasing of feed rate.
According to Figure 4, the surface roughness showed
a decreasing tendency with an increase in the cutting
speed. An improvement in the surface quality was
observed with the increasing cutting speed because the
increasing temperature during the cutting process made
the plastic deformation and the chip flow easier.6,18 Fur-
ther, it is thought that because of a reduction in built-up
edge (BUE) and built-up layer (BUL) formations, tool
wear was affected positively, and so this situation gives
rise to an improvement in the surface quality.6
During a manufacturing process, physical and che-
mical properties of the coolants allow a reduction in
thermal/mechanical-based damages. When coolants are
used efficiently, the dimensional accuracy and a better
surface quality may occur; also, a longer life of the cutt-
ing tool may be obtained. Figure 4 shows a significant
change in the surface-roughness values, depending on
the use of different cooling methods. It can be seen that
the surface roughness is minimum when using cryogenic
M. SARÝKAYA: OPTIMIZATION OF THE SURFACE ROUGHNESS BY APPLYING THE TAGUCHI TECHNIQUE ...
944 Materiali in tehnologije / Materials and technology 49 (2015) 6, 941–948
Figure 4: Effects of machining parameters on the surface roughness
Slika 4: Vpliv parametrov obdelave na hrapavost povr{ine
Figure 3: Experimental set-up
Slika 3: Eksperimentalni sestav
cooling. This may be due to a lower cutting temperature,
a lower adhesion between the cutting insert and the
machined-workpiece surface and a lower tool-wear rate
compared to dry and wet cooling conditions.19 In addi-
tion, a reduction in the surface roughness due to wet
cooling was determined in comparison with dry machin-
ing.
3.2 Signal-to-noise (S/N) analysis
The surface roughness (Ra) was evaluated with an
orthogonal array for each combination of the test para-
meters using the Taguchi technique and an optimization
of the process parameters was achieved with signal-to-
noise (S/N) ratios. Here, the signal data includes the
desired influence on the test results and the noise data
includes the undesired influence on the test results.
Therefore, the maximum S/N ratio provides the optimum
results. There are three different ways of calculating the
S/N ratios. These are the nominal-is-best, the smaller-
the-better and the larger-the-better approaches. In the
present study, the smaller-the-better option of the S/N
quality characteristic was utilized to obtain the best
combination for the surface roughness with respect to the
desired low Ra. The smaller-the-better approach is ex-
pressed as follows:7
Smaller-the-better (minimize):
S
N n
o
R
i
r
n
a
=−
⎡
⎣⎢
⎤
⎦⎥=
∑10 1 2
1
log (2)
In Equation (2), oi is the response of the output cha-
racteristic for the rth test and n is the number of the out-
puts of the test.
The experimental results and their S/N ratios were
calculated using Equation (2) as given in Table 4. From
this table, the mean surface roughness and the mean S/N
ratio were calculated as 2.53 μm and –7.39 dB, respec-
tively. The analysis of the process parameters like the
cutting speed, feed rate and cooling condition was made
using an S/N response table obtained with the Taguchi
method as seen in Table 5. The S/N response table of the
results gives the optimum points of the process para-
meters for the best surface roughness. Figure 5 was
plotted to determine the optimum control factor of a
machining parameter using the S/N response table. As
seen in Figure 5, for the highest S/N ratio, the optimum
parametric combination was found to be factor A (level
3, S/N = –5.985 dB, mean: 2.283 μm), factor B (level 1,
S/N = –3.450 dB, mean: 1.533 μm) and factor C (level 3,
S/N = –5.975 dB, mean: 2.267 μm).
Under cryogenic cooling, the cutting speed was 176
m/min and the feed rate was 0.1 mm/r.
Table 4: Experimental results and their S/N values
Tabela 4: Rezultati eksperimentov in njihove S/N vrednosti
Test
no.
Control parameters
Surface
roughness
Ra/μm
Signal to
noise
(S/N)/dB
A
Cooling
condition
B
Feed rate
f/(mm/r)
C
Cutting
speed
Vc/
(m/min)
1 Dry 0.1 90 1.90 –5.5751
2 Dry 0.16 126 3.55 –11.0046
3 Dry 0.25 176 3.75 –11.4806
4 Wet 0.1 126 1.65 –4.3497
5 Wet 0.16 176 2.00 –6.0206
6 Wet 0.25 90 3.20 –10.1030
7 Cryogenic 0.1 176 1.05 –0.4238
8 Cryogenic 0.16 90 2.15 –6.6488
9 Cryogenic 0.25 126 3.50 –10.8814
Table 5: Response table
Tabela 5: Tabela odgovorov
Levels
Control factors Control factors
S/N ratios Means
A B C A B B
Level 1 –9.353 –3.450 –7.442 3.067 1.533 2.417
Level 2 –6.824 –7.891 –8.745 2.283 2.567 2.900
Level 3 –5.985 –10.822 –5.975 2.283 3.483 2.267
Delta 3.369 7.372 2.770 0.833 1.950 0.633
Rank 2 1 3 2 1 3
3.3 Analysis of variance
Analysis of variance (also known as ANOVA) is a
statistical method and the significance of the machining
parameters was identified with its help. The ANOVA
analysis was performed with a 95 % confidence level and
M. SARÝKAYA: OPTIMIZATION OF THE SURFACE ROUGHNESS BY APPLYING THE TAGUCHI TECHNIQUE ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 941–948 945
Figure 5: Main effect plots for: a) means and b) S/N ratios
Slika 5: Diagram u~inka za: a) sredstva in b) razmerje S/N
5 % significance level. The F values of the control fac-
tors indicated the significance of the control factors
determined with the ANOVA analysis. The percentage
contribution of each parameter is shown in the last co-
lumn of the ANOVA table. The column shows the effect
rates of the input parameters on the outputs.6
In the present work, the ANOVA results are given in
Table 6 and, in addition, these results are graphically
presented in Figure 6. The ANOVA results indicate that
the cooling condition, the feed rate and the cutting speed
influenced the surface roughness by 17 %, 74.1 % and
8.5 %, respectively. Therefore, the feed rate (factor B) is
the most important factor affecting the surface rough-
ness. According to Table 6, it can be said that the
cooling condition, the feed rate and the cutting speed had
a statistical and physical significance with regard to the
surface roughness at the reliability level of 95 % because
their P values are lower than 0.05.
Table 6: ANOVA analysis
Tabela 6: Analiza ANOVA
Factors
Degree
of free-
dom
Sum of
squares
Mean of
squares
F ratio,
 = 0.05 P
Contri-
bution
(%)
Cooling
method 2 1.3106 0.6553 48.14 0.020 17
Feed
rate 2 5.7106 2.8553 209.78 0.005 74.1
Cutting
speed 2 0.6572 0.3286 24.14 0.040 8.5
Error 2 0.0272 0.0136 0.35
Total 8 7.7056 100
3.4 Regression analysis
In many studies, a regression analysis was used to
determine the relationship between the control factors
and experimental results. In the present work, the control
factors are the cutting speed (Vc), the feed rate (f) and the
cooling condition (C) and the surface roughness (Ra) is
the response. On the basis of the first-order model, a ma-
thematical model was created using a regression analysis
for predicting Ra. The first-order model can be expressed
with Equation (3):
y v v v= + ⋅ + ⋅ + ⋅   0 1 1 2 2 3 3 (3)
In this equation, y is the corresponding output, and v1,
v2, and v3 are the values of the variable. The term  is the
regression coefficient. The first-order model can be
written as a function of the cooling condition (C), the
feed rate (f) and the cutting speed (Vc). The relationship
between the output and the turning parameters from
Equation (3) was adapted as given in following Equation
(4):
R C f Va cpre = + ⋅ + ⋅ + ⋅   0 1 2 3 (4)
According to the above equations, a mathematical
model for the surface roughness with coded values
(Table 3) can be written in the following way:
R C f Va cpre = − ⋅ + ⋅ − ⋅156111 0 416667 0 975 0 075. . . . (5)
R2 = 87.98 %
The determination coefficient, expressed as R2, shows
the reliability of the predicted model. It was recom-
mended that R2 should be between 0.8 and 1.20 In this
study, the value of the determination coefficient is R2 =
0.8798 and it is high enough, demonstrating a high
significance of the predicted model. In order to evaluate
the contents of the residual of the model, a graphical
technique was employed. The sufficiency of the models
was investigated by examining the residuals. The nor-
mal-probability plot of the residuals for the surface
roughness is seen in Figure 7. It is seen that the residual
rather appropriately tend towards a straight line, meaning
that errors are normally delivered. This demonstrates that
the predictive model is satisfactory.
3.5 Determining the optimum surface roughness
In the last phase of the Taguchi method, a verification
experiment has to be made to check the reliability of the
optimization.21 The verification experiment was con-
ducted at the optimum levels of the variables determined
as seen in Figure 5. A3-B1-C3 and their values from this
figure were employed to calculate the estimated opti-
mum surface roughness. The equation for estimating the
optimum result (Raopt) was expressed as follows:
R A T B T C T Ta R R R Ropt a a a a= − + − + − +( ) ( ) ( )3 1 3 (6)
M. SARÝKAYA: OPTIMIZATION OF THE SURFACE ROUGHNESS BY APPLYING THE TAGUCHI TECHNIQUE ...
946 Materiali in tehnologije / Materials and technology 49 (2015) 6, 941–948
Figure 7: Normal-probability plot of the residuals
Slika 7: Diagram normalne verjetnosti preostankov
Figure 6: Graphical representation of the ANOVA results
Slika 6: Grafi~en prikaz rezultatov ANOVA
In Equation (6), A3, B1, and C3 are the mean values of
the surface roughness at the optimum level as seen in
Table 5. TRa is the mean of all the Ra values obtained
from the experimental results (Table 4). According to
Equation (6), Raopt is 1.023 μm.
In order to verify the result of the estimated surface
roughness, the confidence interval (CI) was calculated
using following equations:22
CI F V
n R
= ⋅ ⋅ +
⎛
⎝
⎜
⎞
⎠
⎟
 ,1, Ve ep
eff
1 1
(7)
n
N
Teff dof
=
+1
(8)
In Equation (7), F,1,Ve is the F ratio at the 95 % confi-
dence level,  is the significance level, Ve is the degree of
freedom of the error, Vep is the error variance, neff is the
effective number of replications, R is the number of
replications for the verification test. In Equation (8), N is
the total number of tests and Tdof is the total main factor
of the degree of freedom.
According to the F test table, F,1,2 is 18.51. Further,
Vep = 0.0136, R = 3, N = 9, Tdof = 6 and, according to
Equation (8), neff is 1.285. The confidence interval (CI) is
found to be 0.528 using Equations (7) and (8). The
predicted optimum surface roughness with the 95 %
confidence interval is:
[Raopt – CI] < Raexp < [Raopt + CI], i.e., [1.023 – 0.528] <
1.05 < [1.023 + 0.528] = 0.702 < 1.05 < 1.551.
The Raexp, which was found with the experiments, was
within the confidence interval limit. Therefore, the
system optimization was successfully achieved using the
Taguchi method at a significance level of 0.05 in the
turning of the AISI 316Ti stainless steel under different
cutting conditions.
3.6 Experimental validation
Verification experiments of the process parameters
were performed for the best result and the predictive
model at the optimum and at random points. Table 7
shows a comparison of the experimental results and the
estimated results obtained with the Taguchi technique
and mathematical model (Equation (5)). It was seen that
the estimated results and the test results are quite close.
The errors of the statistical analysis must be below 20 %
for the reliability of the analysis.20 Therefore, the results
found in the verification experiment showed that the
optimization was successful.
4 CONCLUSIONS
This study focused on the influences of the process
parameters such as the cooling condition, the feed rate
and the cutting speed on the surface roughness (Ra) in the
turning of the AISI 316Ti stainless steel and an
optimization was achieved on the basis of the Taguchi
method. Cryogenic cooling using liquid nitrogen (LN2)
was applied from within a modified tool holder. The
Taguchi S/N ratio was utilized with the smaller-the-better
approach to obtain the optimum values. An analysis of
variance was performed to define the importance of the
process parameters for the outputs. Based on the first-
order model, a mathematical model was created, namely
Rapre, using the regression analysis. The results obtained
from this study can be summarized as follows:
The best parameter levels were found to be A3-B1-C3
(i.e., cutting condition = cryogenic cooling, feed rate =
0.1 mm/r and cutting speed = 176 m/min). Cryogenic
cooling with LN2 and a modified tool holder provided a
better performance than dry and wet conventional cool-
ing in terms of the surface roughness and may be
recommended for use in the turning of the AISI 316Ti
stainless steel.
Although the surface quality decreased with an in-
crease in the feed rate, it showed an improvement
tendency with an increase in the cutting speed and with
the use of the cryogenic cooling and wet (traditional)
cooling.
Using ANOVA, it was found that the feed rate is the
dominant factor affecting the surface roughness, with a
fraction of 74.1 %, followed by the cooling method and
the cutting speed. Further, it was seen that the cooling
condition, the feed rate and the cutting speed had statis-
tical and physical significance for the surface roughness,
with a reliability level of 95 %.
The regression model showed a high correlation bet-
ween the experimental and predicted values. Further, the
normal-probability plot of the residuals for the surface
roughness showed that the residuals quite appropriately
tended to a straight line, meaning that errors were nor-
mally delivered. This proved that Rapre was satisfactory
and quite reliable. In addition, the value of the determi-
nation coefficient was high enough.
In the verification experiment, the measured values
were within the 95 % confidence interval (CI).
Future work may deal with analyzing the effects of
some cooling/lubrication methods like the minimum-
quantity lubrication (MQL), high-pressure cooling with a
coolant, high-pressure cooling with compressed air, and
external cryogenic cooling during the machining of the
AISI 316Ti stainless steel. Further, other process para-
M. SARÝKAYA: OPTIMIZATION OF THE SURFACE ROUGHNESS BY APPLYING THE TAGUCHI TECHNIQUE ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 941–948 947
Table 7: Verification of the test results
Tabela 7: Preverjanje rezultatov preizkusov
Level
Taguchi technique First-order model
Exp. Pre-dicted
Error
(%) Exp.
Pre-
dicted
Error
(%)
A3B1C3
(optimum) 1.05 1.023 1.8 1.05 1.06 0.9
A3B2C1
(random) 2.15 2.19 1.82 2.15 2.16 0.4
A2B1C2
(random) 1.65 1.55 6.06 1.65 1.66 0.6
meters like the cutting-tool geometry, depth of cut, CVD
coated inserts, uncoated carbide inserts, nose radius, and
chip-breaker geometry may be considered by the
researchers to define their influences on the tool life and
surface quality in future academic studies.
5 REFERENCES
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22 A. Dvivedi, P. Kumar, Surface quality evaluation in ultrasonic drill-
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M. SARÝKAYA: OPTIMIZATION OF THE SURFACE ROUGHNESS BY APPLYING THE TAGUCHI TECHNIQUE ...
948 Materiali in tehnologije / Materials and technology 49 (2015) 6, 941–948
T. KIVAK, U. ªEKER: EFFECT OF CRYOGENIC TREATMENT APPLIED TO M42 HSS DRILLS ...
949–956
EFFECT OF CRYOGENIC TREATMENT APPLIED TO M42 HSS
DRILLS ON THE MACHINABILITY OF Ti-6Al-4V ALLOY
VPLIV PODHLAJEVANJA SVEDROV M42 HSS NA
OBDELOVALNOST ZLITINE Ti-6Al-4V
Turgay Kývak1, Ulvi ªeker2
1Düzce University, Faculty of Technology, Department of Manufacturing Engineering, Düzce, Turkey
2Gazi University, Faculty of Technology, Department of Manufacturing Engineering, Ankara, Turkey
turgaykivak@duzce.edu.tr
Prejem rokopisa – received: 2014-11-16; sprejem za objavo – accepted for publication: 2014-12-16
doi:10.17222/mit.2014.283
This study investigated the effects of deep cryogenic treatment applied to M42 HSS drills on the tool wear, the tool life and the
surface roughness during the drilling of a Ti-6Al-4V alloy under dry and wet cutting conditions. Drilling tests were carried out
using untreated, cryogenically treated, cryogenically treated and tempered, and multi-layered TiAlN/TiN-coated HSS drills.
Four different cutting speeds ((6, 8, 10, 12) m/min) and a constant feed rate of 0.06 mm/r were used as the cutting parameters
and holes with a depth of 15 mm were drilled. At the end of the drilling tests, it was seen that the use of a coolant increased the
tool life and decreased the surface roughness. Among the four tools, the best results in terms of the tool life and surface
roughness were obtained with the multi-layered TiAlN/TiN-coated tool. The cryogenically treated and tempered drills exhibited
an increase of 87 % in the tool life compared to the untreated drills. Scanning electron microscope (SEM) and X-ray diffraction
(XRD) analyses showed that by reducing the size of the carbide particles in the microstructure, cryogenic treatment resulted in a
more uniform carbide distribution and in the transformation of retained austenite to martensite. This played an important role in
the increase in the hardness and wear resistance of the cutting tools.
Keywords: cryogenic treatment, microstructure, M42 HSS, drilling, tool life, surface roughness
V tej {tudiji je bil preiskovan vpliv globokega podhlajevanja svedrov M42 HSS na njihovo obrabo, zdr`ljivost in hrapavost
povr{ine med suhim in mokrim vrtanjem zlitine Ti-6Al-4V. Preizkusi vrtanja so bili izvr{eni z uporabo HSS neobdelanih,
podhlajenih, podhlajenih in popu{~anih ter svedrov z ve~plastnim nanosom TiAlN/TiN. Uporabljene so bile {tiri razli~ne hitrosti
rezanja ((6, 8, 10, 12) m/min) in konstantno podajanje 0,06 mm/r pri vrtanju 15 mm globokih izvrtin. Na koncu preizkusov
vrtanja se je pokazalo, da uporaba hlajenja s teko~ino pove~a zdr`ljivost orodja in zmanj{a hrapavost povr{ine. Med {tirimi
orodji je bil glede na njihovo zdr`ljivost in hrapavost povr{ine najbolj{i rezultat dose`en z orodji z ve~plastnim nanosom
TiAlN/TiN. Podhlajeni in popu{~ani svedri so imeli pove~ano zdr`ljivost za 87 % v primerjavi z neobdelanimi svedri. Analize
na vrsti~nem elektronskem mikroskopu (SEM) in rentgenska difrakcija (XRD) sta pokazali, da z zmanj{anjem velikosti
karbidnih zrn v mikrostrukturi pri podhlajevanju dobimo bolj enakomerno razporeditev karbidov, preostali avstenit pa se
pretvori v martenzit. To ima pomembno vlogo pri pove~anju trdote in odpornosti orodja za rezanje proti obrabi.
Klju~ne besede: obdelava s podhlajevanjem, mikrostruktura, M42 HSS, vrtanje, zdr`ljivost orodja, hrapavost povr{ine
1 INTRODUCTION
With the rapid development of technology, the past
few years have witnessed a rise in the expectations for
the products made of resistant, lightweight materials and
their production methods. In particular, the need for such
materials in the electronics, computer, automotive and
aerospace industries is increasing. Titanium and its
alloys meet a great many of these expectations due to
their low density, high resistance, and heat and corrosion
resistance.1–3 Among these alloys, all having different
properties, Ti-6Al-4V is the most widely used and it is
found in 60 % of industrial applications. This alloy has
the properties of high resistance to fatigue and corrosion
along with high strength and biocompatibility, and co-
vers a wide application field, primarily in the aerospace
industry.4 However, the Ti-6Al-4V alloy belongs to a
group of materials which are difficult to machine be-
cause of their high chemical reactivity and high tendency
to weld to the cutting tool,5 low heat conductivity, main-
tenance of strength at high temperatures and a low elasti-
city module. Furthermore, the production cost of these
materials is high and errors during machining can cause
serious increases in the cost of machining.6–10
The life of cutting tools plays a major role in in-
creasing the productivity and, consequently, it is an
important economic factor. In order to increase the life of
cutting tools, a common approach in the past was to
heat-treat the tool materials, thus providing a greater
control over the range of the properties that a given tool
material might have. In order to increase the life of
cutting tools and improve their properties, the con-
ventional heat treatment, applied especially to tool steel
and high-speed steel (HSS), has been a widely used me-
thod for many years.11 Cryogenic treatment is generally a
complementary treatment to the heat treatment applied to
increase the wear resistance of the materials exposed to
high wear conditions. It is also known as the cold or
subzero treatment.12 It is cheap and permanent; it is done
once and, unlike coatings, it affects the whole piece.13
Cryogenic treatment, depending on the temperatures
Materiali in tehnologije / Materials and technology 49 (2015) 6, 949–956 949
UDK 621.95:621.7.01:539.538 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)949(2015)
applied to the material, is classified as shallow cryogenic
treatment (between –50 °C and –100 °C) and deep cryo-
genic treatment (lower than –125 °C).
After the conventional heat treatment, materials are
first held at the temperatures of shallow or deep cryo-
genic treatment (generally for 24 h) and then brought
gradually up to room temperature.14 In this way, the
formation of fine carbide precipitates, a uniform carbide
distribution and a conversion of the retained austenite to
martensite are achieved. Thus, significant improvements
are obtained in the mechanical properties of the materials
such as the hardness and wear resistance.15–17 Cryogenic
treatment, having been previously applied to tool/die
steels, is now applied to the cutting tools in the machin-
ing and, as a result, important developments have been
obtained in the tool wear, tool life and recovery of
cutting conditions. Studies on cryogenically treated
high-speed-steel tools show microstructural changes in
the material that can considerably influence the tool life
and productivity. In the literature, results showed
tool-life improvements ranging from 92 % to 817 % for
the cryogenically treated HSS tools in the industrial
use.13
The influence of deep cryogenic treatment on the
wear resistance and the tool life of M42 HSS drills with
a high-speed dry-drilling configuration of carbon steels
was studied and the experimental results indicated tool-
life improvements of 77 % and 126 % for cryogenically
treated and cryogenically treated + tempered drills, res-
pectively.16 The improvement in the wear resistance and
the significance of the treatment parameters for different
materials were investigated in another study. It was
found that cryogenic treatment provided an improvement
in the tool life of nearly 110 %. The tool-life improve-
ment was even higher with the use of TiN coatings.18
Compared to the other material-removal processes,
the drilling process has quite a wide application field.
Especially in the aerospace industry, drilling constitutes
a large portion of the material-removal processes, up to
40–60 %.19 In the literature, studies investigating the
machinability of titanium alloys are generally focused on
turning and milling while those dealing with drilling are
very limited.
In this study, cryogenic-treatment-induced changes in
the microhardness and microstructure of M42 HSS tools
and the effects of these changes on the tool wear, tool
life and surface roughness during the drilling of a
Ti-6Al-4V alloy were investigated. Furthermore, the ma-
chining performance of cryogenically treated tools was
examined in comparison with that of the untreated and
multi-layered TiAlN/TiN-coated tools.
2 EXPERIMENTAL METHODS
2.1 Drilling experiments
Drilling tests were carried out using a JOHNFORD
VMC 550-7.5 kW CNC vertical machining center under
dry and wet cutting conditions. Four different cutting
speeds (6, 8, 10 and 12) m/min were used for the experi-
ments and the hole depth and feed rate were kept con-
stant at 15 mm and 0.06 mm/r, respectively. The experi-
mental set-up is shown in Figure 1. For the workpiece
material, 100 mm × 80 mm × 15 mm blocks of the
Ti-6Al-4V alloy were used. Before the experiments, the
sample blocks were ground to eliminate the adverse
effects of any surface defects. The chemical composition
and mechanical properties of the workpiece material are
shown in Tables 1 and 2, respectively.
Table 1: Chemical composition of Ti-6Al-4V alloy (w/%)
Tabela 1: Kemijska sestava zlitine Ti-6Al-4V (w/%)
Ti Al V Fe O C N H
89.85 5.90 4.00 0.08 0.14 0.01 0.01 0.002
Table 2: Mechanical properties of Ti–6Al-4V
Tabela 2: Mehanske lastnosti zlitine Ti-6Al-4V
Tensile
strength (MPa) Yield strength(MPa)
Elongation 5D
(%) Hardness (Rc)
900-1100 830 10 36
In the wet cutting experiments, a 6 % concentration
of a semi-synthetic emulsion was used as the coolant.
For the elimination of the twisting effect, the distance
from the tool holder to the drill tip was determined as 30
mm. This value was kept constant in all the experiments
in order to validate the obtained values. Three holes were
drilled under each machining condition for the compa-
rison of the surface-roughness measurements. The ave-
rage of each set of three measurements was used for the
comparison. As the initial condition of each test, a new
drill was used for each experiment. The surface rough-
ness of the machined holes was measured using a Mohr
Perthometer M1 portable surface-roughness tester for
each machining condition and the average values of the
surface roughness (Ra) were determined. In order to
measure the surface roughness, the Ti-6Al-4V alloy
blocks were sliced parallel to the hole axes and the
measurements were taken at three different points. The
average of these three measurements was used in the
evaluations.
For the tool-wear experiments, as various wear
mechanisms and types were formed on the drills and the
T. KIVAK, U. ªEKER: EFFECT OF CRYOGENIC TREATMENT APPLIED TO M42 HSS DRILLS ...
950 Materiali in tehnologije / Materials and technology 49 (2015) 6, 949–956
Figure 1: Details of the experimental set-up
Slika 1: Podrobnosti eksperimentalnega sestava
cost of the workpiece was high, the following were
determined as the tool-life criteria:20
• the average non-uniform flank wear Vb = 0.15 mm
• the maximum flank wear Vbmax = 0.2 mm
• the chipping = 0.2 mm
• the out-of-corner wear = 0.2 mm
• the fracture or catastrophic failure
As soon as one of the criteria mentioned above was
realized, it was accepted that the tool was worn. A
professional hand-held digital microscope (Dino-Lite,
AM413ZT) and a JEOL JSM-6060 LV scanning electron
microscope (SEM) were used to determine the wear
mechanisms and types. Due to the limited amount and
high cost of the workpiece material, the tool wear
experiments were performed only at the cutting speed of
10 m/min. The tool life time was obtained by multi-
plying the total number of holes drilled by the drilling
time at which the tool reached one of the wear criteria.
2.2 Cryogenic treatment and tempering
A number of uncoated and coated drills (Guhring) of
a 5 mm diameter were cryogenically treated in order to
observe the effect of cryogenic treatment on the drilling
of the Ti-6Al-4V alloy with M42 HSS twist drills. Three
types of uncoated drills were used: untreated drills (U),
cryogenically treated drills (CT), and cryogenically
treated and tempered (at 200 °C for 2 h) drills (CTT).
Cryogenic treatment was not applied to the TiAlN/TiN-
coated drills in order to compare the performance of the
cryogenically treated material to that of the coated
material. The chemical composition and properties of the
M42 HSS twist drills used in the experiments are given
in Tables 3 and 4, respectively.
Table 3: Chemical composition of M42 HSS drills (w/%)
Tabela 3: Kemijska sestava svedrov M42 HSS (w/%)
C Cr Co Mo W V
1.1 4.2 8.0 10 1.8 1.2
Table 4: Properties of M42 HSS drills
Tabela 4: Lastnosti svedrov M42 HSS
Uncoated HSS
(U, CT, CTT) Coated HSS (U)
Tool material M42 M42
Tool reference DIN 338 DIN 338
Point angle 135 ° 135 °
Helix angle 35 ° 35 °
Diameter 5 mm 5 mm
Coating – Multi-layer TiAlN/TiN
Coating thickness – 4 μm
Hardness – 3600 (HV0.05)
The cryogenic treatment of the M42 HSS drills was
performed by gradually lowering the temperature from
room temperature to –145 °C at the cooling rate of about
1–2 °C/min, holding the drills at this cryogenic tempe-
rature for 24 h, and then raising the temperature back to
room temperature at the heating rate of 1–2 °C/min.
Figure 2 schematically illustrates the cryogenic treat-
ment applied to the M42 HSS drills. To verify the forma-
tion of fine and homogeneous carbide particles and the
transformation of the retained austenite to martensite, the
microstructures of the untreated, cryogenically treated,
and cryogenically treated and 2 h tempered drills were
observed via SEM photographs and X-ray diffraction
(XRD) profiles. The microstructure and phase distribu-
tion were characterized with SEM and the volume
fraction of the retained austenite was determined using a
GE-SEIFERT X-ray diffraction instrument with a Cr-K1
X-ray source. From the X-ray diffractograms, the con-
tents of the retained austenite and martensite in the alloy
after different treatments were measured using the
ASTM E975-84 standard.21
3 RESULTS AND DISCUSSION
3.1 Evaluation of cryogenic treatment of drills
In the HSS tools the main alloying elements which
change the microstructure and the properties are C, Cr,
Mo, V, W and Co. Except for Co, these elements precipi-
tate in the microstructure and create carbides. Generally,
seven groups of carbides precipitate in high-speed steels:
(1) E carbide, Fe2.4C (hcp); (2) 
-carbide, M3C (Fe3C);
(3) MC or M4C3, (V4C3); (4) M2C, (W2C or Mo2C); (5)
-carbide, M7C3 (Cr7C3); (6) -carbide, M23C6 (Cr23C6);
(7) -carbide, M6C (Fe3W3C or Fe4W2C), as expressed
in22. The M6C carbides were originally known as
high-speed steels and they are similar to the complex
surface-centered cubical carbides that are rich in tung-
sten and molybdenum and give red hardness to steel. The
distribution of these carbide particles in the microstruc-
ture, their size, their amount and the distances between
them affect the mechanical properties of the material.23
In this study, profiles were obtained from the XRD
analyses. These analyses were made to determine the
differences in the amount of the carbide present in the
cryo-treated (CT) and cryo-treated and tempered (CTT)
drills compared to the untreated (U) tool (Figure 3). For
the purpose of determining the residual-austenite vol-
Materiali in tehnologije / Materials and technology 49 (2015) 6, 949–956 951
T. KIVAK, U. ªEKER: EFFECT OF CRYOGENIC TREATMENT APPLIED TO M42 HSS DRILLS ...
Figure 2: Details of the cryogenic-treatment process: a) schematic
configuration of the cryogenic-treatment system, b) cryogenic treat-
ment and tempering cycle used for M42 HSS drills
Slika 2: Podrobnosti postopka podhlajevanja: a) shematski prikaz
sistema za podhlajevanje, b) cikel podhlajevanja in popu{~anja, upo-
rabljenega pri svedrih M42 HSS
ume, the peaks in the austenite (A200 and A220) and
martensite (M200 and M211) planes were used. With the
cryogenic treatment and cryogenic treatment + tem-
pering, the austenite peaks in the A200 and A220 planes
decreased, while the martensite peaks in the M200 and
M211 planes increased. The volume proportion of the
retained austenite in the untreated tool was measured as
6.5 % and in the cryo-treated and cryo-treated + tem-
pered tools this proportion was 2.4 % and 1.8 %, respec-
tively. Therefore, the cryogenic and tempering treatments
played an important role in the transformation of the
austenite (retained in the structure after the conventional
heat treatment) into martensite. It is believed that the
transformation of the retained austenite into martensite
due to cryogenic treatment can provide significant im-
provements in the mechanical properties of cutting tools
such as hardness. This was verified with the positive
variations that occurred in the hardness and the micro-
structure.
After the cryogenic treatment, a Leica WMHT MOT
microhardness tester was used to measure the Vickers
HV microhardness on one cryogenically treated sample,
one cryogenically treated and 2 h tempered sample and
one untreated sample, with a minimum of eight inden-
tations in each sample and the average used for the com-
parison. In Table 5, the differences in the microhardness
values depending on the treatment applied to the M42
HSS tools are seen. On the untreated tool, the initial
hardness was 703 HV and immediately after the cryo-
genic treatment it became 742 HV. After the cryogenic
treatment, tempering was applied and the hardness was
measured as 718 HV. With the cryogenic treatment and
cryogenic treatment + tempering the percentages of the
increase in the hardness were 5.5 and 2.1, respectively. It
is thought that the increase in the hardness after the
cryogenic treatment is the result of the transformation of
the austenite retained after the conventional heat treat-
ment to martensite.15,24 Moreover, the influence of cryo-
genic treatment on the increasing hardness can be found
in the literature as well.13,25,26 Although hardness values
differ depending on the material type and application
method, increases in the hardness values of 1–3 HRC
can be obtained with the cryogenic treatment.16 The tem-
pering treatment caused some decrease in the hardness
compared to the cryogenic treatment; however, it was
observed that the hardness value was still higher than
that of the untreated tool. It was assumed that this de-
crease in the hardness was the result of the dissociation
T. KIVAK, U. ªEKER: EFFECT OF CRYOGENIC TREATMENT APPLIED TO M42 HSS DRILLS ...
952 Materiali in tehnologije / Materials and technology 49 (2015) 6, 949–956
Figure 4: Microstructures of the drills: a) untreated, b) cryo-treated, c) cryo-treated and tempered
Slika 4: Mikrostruktura svedrov: a) neobdelano, b) podhlajeno, c) podhlajeno in popu{~ano
Figure 3: XRD profiles of HSS tools: a) untreated, b) cryo-treated, c) cryo-treated and tempered
Slika 3: Rentgenogram HSS-orodij: a) neobdelano, b) podhlajeno, c) podhlajeno in popu{~ano
Table 5: Microhardness and retained-austenite volume after different
treatment cycles
Tabela 5: Mikrotrdota in volumen zaostalega avstenita po razli~nih
ciklih obdelave
Cutting tools Retained austenite(/%)
Microhardness
(HV0.2)
Untreated 6.5 703
Cryo-treated 2.4 742
Cryo-treated and
tempered 1.8 718
of some MC (Mo, V, W, Cr) carbides and precipitate
phases.
In order to specify the changes in the microstructure
caused by the cryogenic treatment and cryogenic treat-
ment + tempering applied to the HSS tools compared to
the untreated tools, SEM microstructure photographs
were taken. The purpose of the microstructure examina-
tion was to explain the increasing hardness values and
the improved tool life. The cutting-tool performance
depends on the carbide properties in the microstructure.
The microstructures of the U, CT and CTT HSS tools are
shown in Figure 4. From the photograph of the untreated
HSS-tool microstructure (Figure 4a), it is clear that the
carbide particles in the matrix are large. As a result of
the cryogenic treatment and cryogenic treatment and
tempering, the carbide particles decreased in size and
exhibited a much better distribution (Figures 4b and 4c).
Cryogenic treatment and tempering may cause a further
increase in the particle volume fraction. Compared with
Figures 4a and 4b, in Figure 4c, it is possible to see a
decrease in the size of the particles and a more uniform
distribution of these particles due to the dissolution of
the precipitates and the fracture of large particles.
With the decrease in the size of the carbide particles
and their uniform distribution, the interior stresses in the
martensite structure are relieved and the micro-cracking
sensitivity is minimized, thus providing a significant
improvement in the hardness and wear resistance. The
precipitation of fine carbides as a result of cryogenic
treatment is responsible for the improvement in the wear
resistance.15 The increase in the microhardness as a
result of cryogenic treatment seems to confirm this
thesis. Uygur27 showed that there was a strong relation-
ship between the microstructure hardness and the wear
properties of steel. Cryogenic treatment provides not
only a carbide formation but also a uniform carbide
distribution.15,16 The tempering after the cryogenic
treatment provides the second carbide precipitation and
plays an effective role in relieving interior stresses.16
Thus, in this study, it was thought that non-uniform
carbides of different sizes were subjected to a size
reduction by the cryogenic treatment, and then the
tempering relieved the interior stresses.
3.2 Evaluation of the tool wear
A series of wear experiments was carried out in order
to compare the performances of the U, TiAlN/TiN-
coated, CT and CTT tools under dry and wet cutting
conditions. In the wear experiments, holes with a depth
of 15 mm were drilled at a cutting speed of 10 m/min
and a feed rate of 0.06 mm/r. Under dry cutting condi-
tions, the wear curve could not be obtained because the
tool was subjected to a catastrophic failure due to an
excessive adhesion without showing a regular wear ten-
dency. Figure 5 shows SEM images of the four different
tools tested under dry cutting conditions. It was clearly
observed that the high temperatures generated at the
cutting area due to a lower coefficient of the heat con-
ductivity of titanium in all of the drills caused a built-up
edge (BUE).28 Another BUE formation was observed at
the outer corners, in particular where the cutting speed
was at its maximum. This was because the outer-corner
area was subjected to the extensive heat and chemical
loads due to a greater heat generation. The adhesion ten-
dency was higher with the untreated tool although fewer
holes were drilled with it (Figure 5a). At the specified
cutting parameters, the uncoated (U) tool completed its
life after the drilling of four holes, the CT tool completed
it after five holes, the CTT tool after six holes and the
TiAlN/TiN-coated tool after seven holes. It was observed
that the lower heat conductivity and friction coefficient
of the coating reduced the friction at the tool-chip inter-
face and decreased the BUE formation.
Under wet-cutting conditions, a more regular wear
tendency was observed compared to dry cutting condi-
T. KIVAK, U. ªEKER: EFFECT OF CRYOGENIC TREATMENT APPLIED TO M42 HSS DRILLS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 949–956 953
Figure 6: Number of holes and flank wear for drills under wet cutting
conditions, at 10 m/min cutting speed and 0.06 mm/r feed rate
Slika 6: [tevilo lukenj in obraba bokov pri svedrih v mokrih razmerah
rezanja: hitrost rezanja 10 m/min, hitrost podajanja 0,06 mm/r
Figure 5: SEM images of drills tested under dry cutting conditions, at
10 m/min cutting speed and 0.06 mm/r feed rate: a) U, b) TiAlN/TiN,
c) CT, d) CTT
Slika 5: SEM-posnetki svedrov, preizku{enih v suhih razmerah,
hitrost rezanja 10 m/min in hitrost podajanja 0,06 mm/r: a) U, b)
TiAlN/TiN, c) CT, d) CTT
tions, and with the specified tool-wear criteria, the
flank-wear curve was obtained, depending on the num-
ber of holes. In Figure 6, under wet cutting conditions,
the differences in the number of holes and flank wear
among the four different tools, at the cutting speed of 10
m/min and the feed rate of 0.06 mm/r, are given. The
best performance in terms of the tool wear was obtained
with the multi-layered TiAlN/TiN-coated tool, followed
by the CTT, CT, and U tools, respectively. The U tool
reached a flank wear value of 0.15 mm at the 24th hole,
the CT tool at the 42nd hole, the CTT tool at the 44th hole,
and the TiAlN/TiN-coated tool at the 51st hole. At this
stage it is possible to state that cryogenic and tempering
treatments created a change in the microstructure and
provided an increase in the hardness, which had an
important influence on the tool life. In general, while the
TiAlN/TiN-coated, CT and CTT tools exhibited similar
flank-wear values up to the 27th hole, after this hole, a
wear difference began to be seen. In particular, the CT
and CTT tools reached the tool wear criterion with a
difference of a couple of holes.
In Figure 7, SEM images of the four different cutting
tools tested under wet-cutting conditions are shown.
From these images it can be seen that here the tool wear
is more regular in comparison with the dry cutting con-
ditions. The tendency to form an excessive BUE forma-
tion that was observed with dry cutting conditions was
minimized with the use of a coolant. It is known that
during cutting operations, the coolant forms a thin film
layer at the tool-chip interface and decreases the friction,
at the same time making the chip removal easier and
decreasing the temperatures at the cutting area,29 thus
delaying the wear of the cutting tool, compared to dry
cutting. It was seen that the effective wear type of the U
and TiAlN/TiN-coated tools was the outer-corner wear
(Figures 7a and 7b), while with the CT and CTT tools it
was the flank wear (Figures 7c and 7d). Furthermore, it
was observed that with the CT and CTT tools, the wear
at the outer corner did not form as rapidly as with the
untreated tool, but instead it followed a more uniform
feed along the cutting edges. This was thought to be the
result of the improvement in the wear resistance of the
cutting tool provided with the cryogenic treatment.
During the drilling of the Ti-6Al-4V alloy, a tool-life
curve was prepared to determine the effects of dry and
wet cutting conditions on the tool life (Figure 8). From
the curve, it can be seen that wet cutting conditions pro-
vided a significant increase in the tool life in comparison
with the dry cutting conditions. As the thermal properties
of Ti-6Al-4V are poor, the use of cutting fluids (or
coolants) is very important to improve the tool life3. The
highest tool life obtained under dry cutting conditions
was 2.7 min, while this value reached 19.9 min under
wet cutting conditions. At this stage, the importance of
using a coolant in the drilling of the Ti-6Al-4V alloy was
once more confirmed. Under dry cutting conditions, the
CT, CTT and TiAlN/TiN-coated tools exhibited life
increases of (25, 50 and 68) % compared to the U tool.
These values were (76, 87 and 112) % in the case of wet
cutting conditions. It is believed that cryogenic treatment
and cryogenic treatment + tempering have important
roles in reducing the size of carbide particles, providing
a uniform carbide distribution, transforming the retained
austenite to martensite and increasing the hardness and
wear resistance of cutting tools. The tool wear experi-
ments confirmed this as well.
The CTT tools provided an increase in the tool life of
20 % under dry cutting conditions and 7 % under wet
cutting conditions in comparison with the CT tools. The
TiAlN/TiN-coated drills exhibited the best performance
among the tested drills. It is thought that the TiAlN/TiN
coating has a multi-layer structure and a lower friction
coefficient and makes the chip flow more easily during
the cutting; due to these properties it has an important
influence on the increase in the tool life. Apart from that,
its high hardness and lower friction coefficient, com-
pared with the uncoated tool, affected the increase in the
tool life. It is interesting to note that the TiAlN/TiN-
coated tool had a longer tool life under wet cutting con-
ditions. This was believed to be a result of the solid
T. KIVAK, U. ªEKER: EFFECT OF CRYOGENIC TREATMENT APPLIED TO M42 HSS DRILLS ...
954 Materiali in tehnologije / Materials and technology 49 (2015) 6, 949–956
Figure 7: SEM images of drills tested under wet cutting conditions, at
10 m/min cutting speed and 0.06 mm/r feed rate: a) U, b) TiAlN/TiN,
c) CT, d) CTT
Slika 7: SEM-posnetki svedrov, preizku{enih pri mokrem rezanju:
hitrost rezanja 10 m/min in hitrost podajanja 0,06 mm/r: a) U, b)
TiAlN/TiN, c) CT, d) CTT
Figure 8: Tool-life differences depending on cutting conditions for the
four tools at 10 m/min cutting speed and 0.06 mm/r feed rate
Slika 8: Razlike v zdr`ljivosti orodja v odvisnosti od razmer pri
rezanju pri {tirih orodjih pri hitrosti rezanja 10 m/min in hitrosti
podajanja 0,06 mm/r
lubricating property of the coating acting together with
the lubricating and cooling property of the coolant.
3.3 Evaluation of the surface roughness
Surface finish is also an important index of machi-
nability or grindability because the performance and
service life of the machined/ground components are
often affected by their surface finish, the nature and
extent of residual stresses and the presence of surface or
subsurface microcracks, if any. This is particularly
relevant when this component is to be used under dyna-
mic loading or in conjugation with some other mating
part(s).30 Figure 9 shows the differences in the surface
roughness (Ra) among the four different tools, depending
on the cutting speed and cutting conditions. For all four
tools, with the increase in the cutting speed, the Ra values
decreased up to the cutting speed of 10 m/min, but with
further increases in the cutting speed, some increase in
Ra was observed. It is thought that the decrease in Ra
values with the increase in the cutting speed was due to
the reduction in the BUE size with the temperature inc-
rease at the tool-workpiece interface.31 Moreover, it is
also believed that improved surface quality is influenced
by the reduced friction resulting from the higher tempe-
ratures at the contact area at the tool-workpiece inter-
face.32 The lowest Ra values among the four cutting tools
were obtained for the holes drilled with the TiAlN/TiN-
coated tools, followed by the CTT, CT and U tools,
respectively. This sequence shows a parallelism with the
one for the tool life. The average decreases in the Ra
values of (6.5, 10.4 and 15) % were obtained with the
CT, CTT and TiAlN/TiN-coated tools compared to the U
tools. With the TiAlN/TiN-coated tool, compared to the
other tools, less BUE formation and wear led to an im-
proved surface quality (Figures 5 to 7). It was seen that
the combination of the cryogenic and tempering treat-
ments was the second most effective procedure with
respect to improving the surface quality due to the
positive microstructural changes and the increase in the
hardness and wear resistance, surpassed only by the
improvement achieved with the TiAlN/TiN coating. Wet
cutting conditions provided important improvements for
all four tools regarding the trend toward reduced Ra
values. Under dry cutting conditions, the Ra values were
between 1.4 μm and 1.94 μm whereas under wet cutting
conditions, the values ranged from 0.95 μm to 1.4 μm.
The main reasons for the difficult machinability of the
Ti-6Al-4V alloy are its low heat-conductivity coefficient
and high chemical reactiveness. The use of a coolant
makes the chip removal easier, inhibits the heat forma-
tion in the cutting area, and decreases the BUE formation
to a large extent. Taken altogether, these factors increase
the surface quality in parallel with the increase in the
tool life.
4 CONCLUSIONS
From the observed performance of the U, CT, CTT
and TiAlN/TiN-coated M42 HSS drills in the machining
of Ti-6Al-4V, the following conclusions were drawn:
• Cryogenic treatment significantly improved the wear
resistance and tool life of M42 HSS drills under dry
and wet conditions in the drilling of the Ti-6Al-4V
alloy. Cryogenic treatment and tempering increased
the performance of the cutting tools.
• By reducing the size of the carbide particles, cryo-
genic and tempering treatment enabled their uniform
distribution and increased the concentration as well.
Furthermore, the treatment had an important influ-
ence on the transformation of the retained austenite to
martensite, a process which contributes to the abra-
sive-wear resistance as a result of the increased
hardness.
• The CT and CTT tools, unlike the U tools, exhibited
a performance approaching that of the TiAlN/TiN-
coated tools. It was seen that the use of a coolant also
had a significant influence on the increase in the tool
life and surface roughness. In dry cutting conditions,
T. KIVAK, U. ªEKER: EFFECT OF CRYOGENIC TREATMENT APPLIED TO M42 HSS DRILLS ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 949–956 955
Figure 9: Different Ra values for the four tools, depending on cutting
condition and cutting speed: a) dry, b) wet
Slika 9: Razli~ne vrednosti Ra pri {tirih orodjih v odvisnosti od raz-
mer pri rezanju in hitrosti rezanja: a) suho, b) mokro
the CT, CTT and TiAlN/TiN-coated tools exhibited
an increase of (25, 50, and 68) % in the tool life,
compared to the U tools. Under wet cutting condi-
tions, these values were (76, 87 and 112) %. Under
dry cutting conditions, the effective wear types were
BUE and catastrophic failure, whereas under wet
cutting conditions, flank wear and outer-corner wear
were effective.
• The biggest advantage of cryogenic treatment com-
pared to coatings is its cheapness and its influence on
the whole piece of the material. In this study, the
results that were obtained showed that cryogenic
treatment, with some improvement, can serve as an
alternative to coatings.
Acknowledgments
The authors wish to express their sincere thanks to
the Gazi University Scientific Research Project Division
for its financial support for Project No: 07/2010-38.
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M35MHV as a function of the processing route, Journal of Materials
Processing Technology, 143–144 (2003), 555–560, doi:10.1016/
S0924-0136(03)00359-5
23 J. Jeleñkowski, A. Ciski, T. Babul, Effect of deep cryogenic treat-
ment on substructure of HS6-5-2 high speed steel, Journal of
Achievements in Materials and Manufacturing Engineering, 43
(2010) 1, 80–87
24 S. Zhirafar, A. Rezaeian, M. Pugh, Effect of cryogenic treatment on
the mechanical properties of 4340 steel, J. Mater. Process. Technol.,
186 (2007), 298–303, doi:10.1016/j.jmatprotec.2006.12.046
25 Y. Dong, X. Lin, H. Xiao, Deep cryogenic treatment of high-speed
steel and its mechanism, Heat Treatment of Metals, 3 (1998), 55–59
26 A. Çiçek, T. Kývak, Ý. Uygur, E. Ekici, Y. Turgut, Performance of
cryogenically treated M35 HSS drills in drilling of austenitic stain-
less steels, Journal of Advanced Manufacturing Technology, 60
(2012) 1–4, 65–73, doi:10.1007/s00170-011-3616-8
27 I. Uygur, Microstructure and wear properties of AISI 1038H steel
weldments, Lubr. Tribol. Ind., 58 (2006), 303–311, doi:10.1108/
00368790610691383
28 L. N. Lüpez de Lacalle, J. Pérez, J. I. Llorente, J. A. Sánchez, Ad-
vanced cutting conditions for the milling of aeronautical alloys,
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29 S. Palanisamy, S. D. McDonald, M. S. Dargusch, Effects of coolant
pressure on chip formation while turning Ti6Al4V alloy, Int. J.
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j.ijmachtools.2009.02.010
30 N. R. Dhar, M. Kamruzzaman, M. Ahmed, Effect of minimum quan-
tity lubrication (MQL) on tool wear and surface roughness in turning
AISI-4340 steel, J. Mater. Process. Technol., 172 (2006), 299–304,
doi:10.1016/j.jmatprotec.2005.09.022
31 C. Ibrahim, Machining of austenitic stainless steels using CVD
multi-layer coated cemented carbide tools, Tribol. Int., 39 (2006),
565–569, doi:10.1016/j.triboint.2005.05.005
32 E. M. Trent, Metal cutting, Butterworths Press, London 1989, 1–171
T. KIVAK, U. ªEKER: EFFECT OF CRYOGENIC TREATMENT APPLIED TO M42 HSS DRILLS ...
956 Materiali in tehnologije / Materials and technology 49 (2015) 6, 949–956
J. KRYSTEK, R. KOTTNER: LOAD-CAPACITY PREDICTION FOR THE CARBON- ...
957–960
LOAD-CAPACITY PREDICTION FOR THE CARBON- OR
GLASS-FIBRE-REINFORCED PLASTIC PART OF A WRAPPED
PIN JOINT
NAPOVED NOSILNOSTI PLASTI^NIH DELOV ZATI^NEGA
SPOJA, OJA^ANEGA Z OGLJIKOVIMI ALI STEKLASTIMI
VLAKNI
Jan Krystek, Radek Kottner
University of West Bohemia in Pilsen, NTIS – New Technologies for the Information Society, Technická 8, 306 14 Plzeò, Czech Republic
krystek@kme.zcu.cz, kottner@kme.zcu.cz
Prejem rokopisa – received: 2014-11-28; sprejem za objavo – accepted for publication: 2015-01-09
doi:10.17222/mit.2014.289
A joint using a metal pin is one possibility of how to achieve a removable joint of composites. The load capacity of a wrapped
pin joint depends on many parameters, especially on the types of fibres and resin, and geometric properties of the joint. The
composite part of a wrapped pin joint is exposed to a combination of the tension in the longitudinal direction and local com-
pression in the transverse direction. The values of the compressive stress in the transverse direction can exceed several times the
uniaxial compressive strength. In this work, CFRP (carbon-fibre-reinforced plastic) and GFRP (glass-fibre-reinforced plastic)
parts of wrapped pin joints were tested. Experimental specimens with different geometries were exposed to a quasi-static
loading. A Zwick/Roell Z050 testing machine was used for the tensile tests. Moreover, the load capacities of the carbon or glass
composite parts were determined using a finite-element analysis. A new measure based on the LaRC04 criterion was proposed
for the prediction of the load capacity. The numerical and experimental results were compared.
Keywords: composite, finite-element method, load capacity, loop criterion, wrapped pin joint
Spoji z uporabo kovinskega zati~a so ena od mo`nosti, kako dose~i odstranljivo kompozitno povezavo. Nosilnost zavite sti~ne
povezave je odvisna od mnogih parametrov, posebno od vrste vlaken in smole ter geometrijskih lastnosti povezave. Kompozitni
del zavite zati~ne povezave je izpostavljen kombinaciji napetosti v vzdol`ni smeri in lokalnim tlakom v pre~ni smeri. Vrednosti
tla~nih napetosti v pre~ni smeri lahko ve~krat prese`ejo enoosno tla~no trdnost. V tem delu so bili preizku{eni zaviti zati~ni
spoji z deli iz CFRP (plastika, oja~ana z ogljikovimi vlakni) in GFRP (plastika, oja~ana s steklenimi vlakni). Preizkusni vzorci z
razli~no geometrijo so bili izpostavljeni kvazistati~ni obremenitvi. Za natezne preizkuse je bila uporabljena naprava
Zwick/Roell Z050. Poleg tega je bila nosilnost kompozitnih delov z ogljikovimi ali steklastimi vlakni dolo~ena z uporabo anali-
ze kon~nih elementov. Novo merilo, ki temelji na merilu LaRC04, je bilo predlagano za napovedovanje nosilnosti. Primerjani so
numeri~ni in eksperimentalni rezultati.
Klju~ne besede: kompozit, metoda kon~nih elementov, nosilnost, merilo zanke, zavit spoj s kovinskim zati~em
1 INTRODUCTION
Joints are often the critical parts of constructions.
This work focuses on wrapped pin joints. The main
principle of manufacturing the wrapped pin joints is to
place the wrapping fibres of a composite directly around
the metal pin, precisely following its shape. It allows
creating a joint without any cutting fibres, which results
in a high load capacity of the joint.1
The curved composite part (loop – Figure 1) of a
wrapped pin joint is exposed to a combination of the
tension in the longitudinal direction (1 in Figure 2) and
the compression in the transverse direction (3 in Figure
2) during a tensile loading of the joint. The tensile and
compressive stresses in the loop reach significantly high
values compared with the ultimate strengths in the
principal material directions. No standard criterion for a
correct failure prediction of the loop was found.1,2
Therefore, the LaRC043 criterion was adjusted2 so that
the failure of the loop was also described. In the case of
freely fastened (FF) loops, a matrix failure reduces the
loop’s load capacity (if a matrix failure occurs). The
matrix failure results in a separation of the wrapping’s
cross-section. In the case of tightly fastened (TF) loops,
a matrix failure does not influence the loop’s load
capacity. The difference between the fastenings is
explained in Figure 3.
However, the parameters of the adjusted LaRC04 cri-
terion2 depend on the types of the composite and fasten-
ing.4 Therefore, a new criterion was proposed in this
work.
Experimental and numerical investigations of the
loop’s load capacity for a wrapped pin joint and a com-
parison of the experimental and numerical results were
the aims of this work. Two types of composite fibres
were used in these analyses.
2 EXPERIMENTS
Experimental specimens (unidirectional composite
loops – Figure 1) were manufactured using the fila-
Materiali in tehnologije / Materials and technology 49 (2015) 6, 957–960 957
UDK 621.792.6:624.046:519.61/.64 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)957(2015)
ment-winding technology. Carbon or glass fibres and
epoxy resin were used. Mechanical properties of the
composites are presented in Table 1.
The joints were exposed to a tensile quasi-static load-
ing (the loading speed was 0.5 mm/min) at a temperature
of +22 °C. A Zwick/Roell Z050 testing machine was
used. The description of geometric parameters is obvious
from Figure 2: the diameter of pins (D) is 8 mm, the
thickness of the loops (H) is 3 mm and the width (Q) is
1–9 mm. FF and TF loops were tested.
Table 1: Mechanical properties of the tested composites
Tabela 1: Mehanske lastnosti preizku{enih kompozitov
carbon fibres + epoxy (LG120 + EM100) (Vf = 0.65)
E1
(GPa)
118.9
E2
(GPa)
4.7
G12
(GPa)
2.0
v12
–
0.35
XT
(MPa)
3264
YC
(MPa)
92
SL
(MPa)
48
glass fibres + epoxy (LH298 + H512) (Vf = 0.73)
E1
(GPa)
52.6
E2
(GPa)
8.6
G12
(GPa)
4.7
v12
–
0.30
XT
(MPa)
2000
YC
(MPa)
50
SL
(MPa)
48
A comparison of force-displacement curves of the
specimens with the same geometry (Q = 5.5 mm) and of
different materials and types of fastening is shown in
Figure 4. Experimental dependencies of the load capa-
city Fmax on the width Q are presented in Figure 5 (TF
J. KRYSTEK, R. KOTTNER: LOAD-CAPACITY PREDICTION FOR THE CARBON- ...
958 Materiali in tehnologije / Materials and technology 49 (2015) 6, 957–960
Figure 5: Load capacity of TF loops – experiment
Slika 5: Nosilnost TF-zank – preizkus
Figure 2: Geometric parameters of the loop and the principal material
directions
Slika 2: Geometrijski parametri zanke in osnovne smeri materiala Figure 4: Comparison of force-displacement curves
Slika 4: Primerjava krivulj sila – raztezek
Figure 3: Types of loop fastening
Slika 3: Vrsta pritrditve zankeFigure 1: Loops with different geometries
Slika 1: Zanke z razli~no geometrijo
loops) and Figure 6 (FF loops). The diameter D was
8 mm and the thickness H was 3 mm. Typical failures of
the FF and TF loops are presented in Figure 7.
2.1 Numerical simulations
The finite-element system MSC. Marc was used for
the numerical simulation. Linear hexahedral elements
with eight nodes (SOLID elements) were used in a para-
metrically created model. Due to the symmetry of the
loop, only one eighth of the loop was modelled. The pro-
perties of transversely isotropic material were assigned
to the elements considering the orientations of the fibres.
The loading was controlled with a displacement of the
rigid surface, which simulated the pin (Figure 8). The
friction was neglected.
The loop criterion was proposed in this work. It is
based on the LaRC04 criterion3 and respects the specific
combination of the stresses in a loop.4 The failure index
in the case of the fibre-failure mode (1 > 0) of the loop
criterion is:
FI
X X P
Y X P
P X
F T T
f
C T
m
f
T
=
− ⋅
+ ⋅
⋅ ⋅ +
≤


1
3
1 (1)
where Pf and Pm are the criterion parameters (the loop
parameters), X T is the tensile strength in the longitudi-
nal direction and Y C is the compressive strength in the
transverse direction. The failure index in the case of the
matrix-failure mode (3 < 0, 1  0) of the loop criterion
is:
FI
S P S PM
T
T T
n 1 m
L
L L
n 1 m
=
− +
⎛
⎝
⎜
⎞
⎠
⎟ +
− +
⎛
⎝
⎜
⎞
⎠
⎟ ≤

  

  
2 2
1
(2)
where T and L are the stresses in the plain of the fail-
ure (T – transverse direction, L – longitudinal direction),
T and L are the coefficients of friction, ST is the trans-
verse shear strength and SL is the longitudinal shear
strength. The loop criterion was implemented into the
finite-element system MSC. Marc.
3 RESULTS
The dependencies of the loop’s load capacity Fmax on
the width Q, based on both the experimental and nume-
rical results are presented in Figures 9 and 10. It is
obvious that an increase in the width Q did not have a
significant influence on the FF loop’s load capacity. The
load capacity of the TF loops increased with the width
Q, but only up to the determinate width (Q = 5 mm).
The difference between the experimental and nume-
rical results was minimized in the process of identifying
the loop parameter Pf. The loop parameter Pf = 0.87 was
identified for the carbon composite and it was Pf = 0.65
for the glass composite (the value of the loop parameter
Pf does not depend on the fastening type). The influence
of the loop parameter Pm on the load capacity was not
investigated. Pm was assumed to be 0.05.2
J. KRYSTEK, R. KOTTNER: LOAD-CAPACITY PREDICTION FOR THE CARBON- ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 957–960 959
Figure 7: Failures of a: a) FF loop and b) TF loop
Slika 7: Poru{itev: a) FF-zanke in b) TF-zanke
Figure 6: Load capacity of FF loops – experiment
Slika 6: Nosilnost FF-zank – preizkus
Figure 8: Mesh of the finite-element model
Slika 8: Mre`a modela kon~nih elementov
Very good agreement between the experimental and
numerical results was achieved in the case of the FF car-
bon loops (Figure 9). In the case of the TF carbon loops,
the difference between the experimental and numerical
results was lower than 17 %. In the case of the glass
composite, sufficient agreement between the experimen-
tal and numerical results was achieved for both types of
loop fastening (Figure 10). The difference between the
experimental and numerical results was lower than 25 %.
4 CONCLUSION
Experimental dependency of a loop’s load capacity
on the loop’s width was investigated. Carbon loops had a
higher load capacity than glass loops. Irrespective of the
type of fibres, the TF loops exhibited a higher load capa-
city than the FF loops. In the case of the FF loops, an
increase in the width Q did not have a significant influ-
ence on the loop’s load capacity. In the case of the TF
loops, the influence of the width Q was significant only
up to its determinate value.
The loop criterion for determining the loop’s load
capacity was proposed using the finite-element method.
The values of the loop parameter Pf were identified for
both analysed materials. With regard to the proposed
loop criterion, this parameter does not depend on the
type of the loop fastening. The maximum differences
between the experimental and numerical results were
17 % for the carbon composite and 25 % for the glass
composite.
Acknowledgement
This publication was supported by the project
LO1506 of the Czech Ministry of Education, Youth and
Sports.
5 REFERENCES
1 T. Havar, E. Stuible, Design and testing of advanced composite load
introduction structure for aircraft high lift devices, ICAF 2009,
Bridging the Gap between Theory and Operational Practice, Pro-
ceedings of the 25th Symposium of the International Committee on
Aeronautical Fatigue, Rotterdam, 2009, 365–374, doi:10.1007/978-
90-481-2746-7_21
2 R. Kottner, J. Krystek, R. Zem~ík, J. Lomberský, R. Hynek, Strength
analysis of carbon fiber-reinforced plastic coupling for tensile and
compressive loading transmission, 52nd AIAA/ASME/ASCE/AHS/
ASC Structures, Structural Dynamics and Materials, Denver, USA,
2011, doi:10.2514/6.2011-1982
3 S. T. Pinho, C. G. Dávila, P. P. Camanho, L. Iannucci, P. Robinson,
Failure Models and Criteria for FRP under In-Plane or Three-Dimen-
sional Stress States Including Shear Non-Linearity, Research report,
NASA/TM-2005-213530, NASA Langley Research Center, 2005
4 J. Krystek, Damage of composite components under various types of
loading, Ph.D. thesis, University of West Bohemia, 2014 (in Czech)
J. KRYSTEK, R. KOTTNER: LOAD-CAPACITY PREDICTION FOR THE CARBON- ...
960 Materiali in tehnologije / Materials and technology 49 (2015) 6, 957–960
Figure 10: Load capacity of glass loops
Slika 10: Nosilnost zank s steklenimi vlakni
Figure 9: Load capacity of carbon loops
Slika 9: Nosilnost zank z ogljikovimi vlakni
K. MRAMOR et al.: A MESHLESS MODEL OF ELECTROMAGNETIC BRAKING FOR THE CONTINUOUS CASTING ...
961–967
A MESHLESS MODEL OF ELECTROMAGNETIC BRAKING FOR
THE CONTINUOUS CASTING OF STEEL
BREZMRE@NI MODEL ELEKTROMAGNETNEGA ZAVIRANJA PRI
KONTINUIRANEM ULIVANJU JEKLA
Katarina Mramor1, Robert Vertnik2, Bo`idar [arler1,2
1University of Nova Gorica, Vipavska 13, 5000 Nova Gorica, Slovenia
2Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
katarina.mramor@ung.si, bozidar.sarler@imt.si, robert.vertnik@imt.si
Prejem rokopisa – received: 2015-04-21; sprejem za objavo – accepted for publication: 2015-06-08
doi:10.17222/mit.2015.084
The application of magnetohydrodynamics in the continuous casting of steel enables improved control of the quality of the
strand. The most common applications are electromagnetic braking (EMBR) and electromagnetic stirring (EMS). The former
slows the flow by applying a static magnetic field and thus improves the steel flow pattern, reduces the velocity and the turbu-
lence of the flow, increases the cleanliness of the material, improves the surface quality and reduces the number of inclusions,
whereas the latter stirs the flow by applying an alternating magnetic field and thus improves the quality of the strand, reduces
the surface and subsurface defects, enhances the solidification and reduces the number of breakouts.
In this contribution EMBR in a continuous-casting process is considered. The local radial basis function collocation method
(LRBFCM) is used for the solution of coupled mass, energy, turbulent fluid flow, species and magnetic field equations. The
explicit Euler time-stepping scheme and the collocation with multiquadrics radial basis functions on the five-noded overlapping
influence domains are used to obtain the solution of the partial differential equations. The Abe-Kondoh-Nagano low Reynolds
turbulence model is used to describe the turbulent fluid flow, whereas the fractional step method is used to solve the pressure-
velocity coupling. The method has been thoroughly tested in several test cases. In the present article the influence of the
application of electromagnetic braking on the macro-segregation in the continuous-casting process for carbon steel is presented.
Keywords: LRBFCM, continuous casting of steel, turbulent flow, magnetic field, macro-segregation
Uporaba magnetohidrodinamike pri kontinuiranem ulivanju jekla omogo~a izbolj{ano kontrolo kakovosti `ile. Najpogostej{i
aplikaciji sta elektromagnetno zaviranje (EMBR) in elektromagnetno me{anje (EMS). Prva zavira tok z uporabo stati~nega
magnetnega polja in tako izbolj{a tokovni vzorec jekla, zmanj{a hitrost in turbulenco toka, pove~a ~istost materiala, izbolj{a
kvaliteto povr{ine in zmanj{a {tevilo vklju~kov, medtem ko druga me{a tok z uporabo izmeni~nega magnetnega polja in tako
izbolj{a kvaliteto `ile, zmanj{a nepravilnosti na povr{ini in pod njo, pospe{i strjevanje in zmanj{a {tevilo prodorov.
V tem prispevku je obravnavano EMBR pri kontinuiranem ulivanju jekla. Lokalna kolokacijska metoda z radialnimi baznimi
funkcijami (LRBFCM) je uporabljena za re{evanje sklopljenih ena~b za maso, energijo, turbulenten tok teko~ine, koncentracijo
sestavin in magnetno polje. Eksplicitna Euler-jeva ~asovna shema in kolokacija z multikvadri~nimi radialnimi baznimi funk-
cijami na petto~kovnih prekrivajo~ih se poddomenah sta uporabljeni za re{itev parcialnih diferencialnih ena~b. Abe-Kondoh-
Naganov turbuletni model za nizka Reynolds-ova {tevila je uporabljen za opis turbulentnega toka, medtem ko je metoda delnih
korakov uporabljena za re{itev tla~no-hitrostne sklopitve. Metoda je bila izdatno preizku{ena na ve~ preizkusnih primerih.
V tem ~lanku je predstavljen vpliv elektromagnetnega zaviranja na makroizcejanje pri kontinuiranem ulivanju oglji~nega jekla.
Klju~ne besede: LRBFCM, kontinuirano ulivanje jekla, turbulentni tok, magnetno polje, makroizcejanje
1 INTRODUCTION
In the manufacturing of steel,1 which has in recent
years greatly expanded, continuous casting is one of the
most common processes in steel production.2 The de-
mand for cast steel of high quality fuels the need to
further improve the casting process. One way to do that
is to introduce either a static or an alternating electro-
magnetic (EM) field. In general, EM devices in the con-
tinuous casting of steel are divided into electromagnetic
brakers (EMBR), which use a direct current to produce a
static EM field, and into electromagnetic stirrers, which
use an alternating current to produce an alternating
magnetic field. The EM force, which is a result of the
applied magnetic field in both cases, affects the velocity,
temperature and concentration fields. By adjusting the
magnetic field, the amount of defects, inclusions and air
bubbles in the material can be significantly reduced. In
the present contribution, the application of an EMBR
system and its effects on velocity, temperature and
concentration fields are presented.
The quality of the final product depends on the
magnitudes of velocity, temperature, concentration and
magnetic fields. In the continuous-casting process the
velocity, temperature, concentration and magnetic fields
are difficult, if not impossible, to measure. The nume-
rical models are therefore applied in order to help us
better understand and further improve the process. The
problem under consideration has already been con-
sidered with several different numerical models, among
which are the Finite Volume Method (FVM),3–8 the Fini-
te Element Method (FEM),9 and some more advanced
meshless methods, like the Local Radial Basis Function
Materiali in tehnologije / Materials and technology 49 (2015) 6, 961–967 961
UDK 519.61/.64:621.74.047:532 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)961(2015)
Collocation Method (LRBFCM),10 which is used in the
present case as well. The purpose of this article is to
present the results obtained for the application of EMBR
in the CC process for carbon steel. The results are pre-
sented with and without magnetic field for the velocity,
temperature and concentration fields.
2 GOVERNING EQUATIONS
The system of governing equations that describes the
heat transfer, turbulent fluid flow, species concentration,
and magnetic field in the continuous casting of steel, is
based on the Reynolds time-averaging approach for
modeling the turbulent flow11 and mixture continuum
formulation, first introduced by Bennon and Incropera.12
Our model consists of six time-averaged equations:
∇⋅ =v 0 (1)
( )∂ ∂
( )
( ) ( )

  


v
vv v v
t
p l
T+ ∇⋅ = −∇ + ∇ +
⎛
⎝
⎜⎜
⎞
⎠
⎟⎟ ∇ + ∇
⎡
⎣
⎢
⎤
⎦
L
L
⎥ −
− ∇ − − − +
+ − +
2
3
10
2
( )
( )
( )
( ) (
 
  
k
K f
f
T T C
L
L
L
3 S
T ref C
v v
g( )− + ×Cref ) j B
(
2)
( )
∂
∂
( )
( ) ( )
( )( )

 


h
t
h T
f h h f
v
+∇⋅ = ∇⋅ ∇ +
+∇⋅ − − +∇
v
v vS L S S L
L t
t
L 
∇
⎛
⎝
⎜
⎞
⎠
⎟ +h
j
2 (3)
∂
∂
( )
( ) ( )
( )(

  

C
t
C f D C f D C
C C
+∇⋅ = ∇⋅ ∇ + ∇ +
+∇⋅ −
v
v
S S S L L L
L( )− +∇ ∇
⎛
⎝
⎜
⎞
⎠
⎟v S
L t
C
L)
f
C


(4)
∂
∂
( )
( )

 




k
t
k k
P G
+∇⋅ = ∇⋅ +
⎛
⎝
⎜
⎞
⎠
⎟∇
⎡
⎣⎢
⎤
⎦⎥
+
+ + −
v L
L
t
k
k k   


− +
−
D
K f
f
kk – L
L
L
3
L
0
21( )
(5)
∂
∂
( )
( )

  




 

t
E+∇⋅ = ∇⋅ +
⎛
⎝
⎜
⎞
⎠
⎟∇
⎡
⎣⎢
⎤
⎦⎥
+ −
−
v L
L
t
k –
[ ]     L L
L
3 k k
K f
f
c f P c G c f
k
0
2
1 1 3 2 2
1( )
( )
−
+ −
(6)
Where v is the velocity of the mixture,  = S = L is
the density (assumed to be constant and equal in both
phases). t stands for time and p for pressure. μt is the
turbulent viscosity and μL is the dynamic viscosity, k
represents the turbulent kinetic energy and K0 is the
permeability constant. vS, T, C, g, T, Tref, C, and Cref
represent the velocity of the solid phase, the thermal
expansion coefficient, the solute expansion coefficient,
the gravitational acceleration, the temperature, the refe-
rence temperature, the species concentration and the
reference species concentration, respectively. j×B stands
for the Lorentz force, h for the enthalpy, and  for the
thermal conductivity. fS, fL, hS, and hL represent the solid
volume fraction, the liquid volume fraction, the enthalpy
of the solid phase and the enthalpy of the liquid phase, vt
is the turbulent kinematic viscosity and  is the electrical
conductivity. DS, and DL are diffusion coefficients for the
solid and liquid phases, respectively.  stands for the
dissipation rate, t, C, k, , c, f1, c and f2 are closure
coefficients. Pk, Gk, Dk-, and Ek- are the shear pro-
duction of turbulent kinetic energy, the generation of
turbulence due to the buoyancy force, the source term in
the k equation and the source term in the  equation, res-
pectively. In this contribution, only the porous zone is
considered and is modelled using Darcy’s law and the
Kozeny-Carman relation. The closure relations, defined
by Abe-Kondoh-Nagano,13 are used to set the turbulence
closures. A detailed description of the closure coeffi-
cients, source terms and damping functions are given in
the paper of [arler et al.14 The Lorentz force is defined
as:
Fm = j×B (7)
where j and B are the current density and the magnetic
flux density. The Maxwell’s equations are used to calcu-
late the current density:
j = (–ø+v×B) (8)
where ø is the fluid’s electric potential. The assumption
of a low magnetic Reynold’s number (Rem << 1) is
made.
2.1 Boundary and initial conditions
The governing equations for velocity, species concen-
tration and temperature in the continuous-casting process
K. MRAMOR et al.: A MESHLESS MODEL OF ELECTROMAGNETIC BRAKING FOR THE CONTINUOUS CASTING ...
962 Materiali in tehnologije / Materials and technology 49 (2015) 6, 961–967
Figure 1: Simplified 2D model for continuous casting of steel
Slika 1: Poenostavljen 2D-model za kontiunirano ulivanje jekla
are strongly coupled. Although the coupling between the
magnetic field and the rest of the governing equations is
weak, it is still very important for the solution to the
problem of how the initial and boundary conditions are
chosen. In present case, five different boundaries are
chosen: inlet, free surface, wall, outlet and symmetry.
The model of the domain is presented in Figure 1, the
computational domain is depicted in Figure 2, and the
initial and boundary conditions are given in Figure 3.
3 SOLUTION PROCEDURE
The explicit Euler time stepping and LRBFCM are
used to solve the governing equations of the EMBR in
the continuous-casting process. The pressure-velocity
coupling is solved with the Fractional Step Method
(FSM).15
The first step in the solution procedure is the calcu-
lation of the initial Lorentz force (Equation (7)). The
procedure begins by solving the Poisson’s equation for
electric potential:
ø  v×B (9)
the solution of which is then inserted into Equation (7).
The Lorentz force is inserted into the equation for the
intermediate velocity v*, which is calculated from the
momentum equation by omitting the pressure-gradient
term. The pressure is then calculated from the Poisson’s
equation by solving the pressure sparse matrix.14 The
calculated pressure gradient is then used to correct the
intermediate velocities of the final velocity field. After
the solution of the velocity field, the equations for tur-
bulent kinetic energy and dissipation rate are solved.
This is followed by the solution of the enthalpy and
species concentration equations. The enthalpy-tempe-
rature14 constitutive relation is used to calculate the
temperature from the enthalpy. Finally, the turbulent
viscosity, velocity, temperature, species concentration,
turbulent kinetic energy and dissipation rate are updated
and the solution is ready for the next step.
The spatial discretization is solved using LRBFCM
by constructing the approximation function 
, that is re-
presented on each of the subdomains as a linear combi-
nation of the radial basis functions (RBFs) as:

  ( ) ( )l n l i l n
i
M
l ip p=
=
∑
1
(10)
where M, li, and li represent a number of shape func-
tions, an expansion coefficient, and RBF shape func-
tions, centred at points lpn, respectively. The most
commonly used RBFs are Multiquadric RBF16,17:
l i l ir c ( ) ( )p p= +
2 2 (11)
where c stands for a dimensionless shape parameter,
which is in our case set to 32, and:
l i
i
l i
i
l i
r
x x
x
y y
y
( )
max max
p =
−⎛
⎝
⎜
⎞
⎠
⎟ +
−⎛
⎝
⎜
⎞
⎠
⎟
2 2
(12)
is scaled by lximax, and lyimax, the scaling parameters in
the subdomains in the x and y directions, respectively
(Figure 3).
A subdomain consists of the lM – 1 nodes nearest to
the node lpn and is formed around each of the calculation
points. In this contribution, five-nodded overlapping sub-
domains are used. A linear system of equations is
obtained by considering the collocation condition:
l
 (lpn) = 
i(l,n) (13)
To construct the Partial Differential Equation (PDE)
derivatives, originating from the governing equations, the
first and the second derivatives of the function l
(p) have
to be calculated:
∂
∂
∂
∂
j
j l
j
j l i
i
M
l i



 ( ) ( )p p=
=
∑
1
(14)
where the index j is used to denote the order of the
derivative and  = x, y. A detailed explanation of the
solution procedure is given in.18,19 The discretization
scheme is shown in Figure 3.
4 RESULTS
The numerical procedure has so far been tested on
the following benchmark test cases: lid driven cavity,
natural convection in a cavity with a magnetic field, and
a backward facing step with a transverse magnetic field
and a test case for a simplified magnetic field in the
continuous-casting process. As the results in all of the
test cases are in good agreement with the reference
results, both those calculated with the commercial code
and those obtained from the literature, the method is now
applied to the electromagnetic braking problem for the
continuous casting of steel. The results of the EMBR
K. MRAMOR et al.: A MESHLESS MODEL OF ELECTROMAGNETIC BRAKING FOR THE CONTINUOUS CASTING ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 961–967 963
Figure 2: Computational domain scheme
Slika 2: Shema ra~unske domene
problem with continuous casting have been published in
several articles.19–21
4.1 EM field calculations
The magnetic field for the EMBR is calculated
analytically. The EMBR device consists of two coils, as
shown in Figure 4. The magnetic fields in these coils
can either face one another or point in a parallel direc-
tion, as shown in Figures 5 and 6. The magnetic field of
both coil configurations is shown in Figures 7 and 8.
The parallel coil configuration is chosen, as this is the
default configuration for EMBR. As the coils in the
EMBR device have iron cores, the magnetic field is
enhanced due to the magnetization, as shown in Figure
9.
In general, steel is ferromagnetic, and thus the pos-
sibility of the influence of magnetization needs to be
checked. The operating temperature in the strand is well
above the Curie temperature, where steel is paramag-
netic. The magnetic field in the molten steel in the strand
is therefore not further influenced by the external mag-
netic field. The temperature dependence of the perme-
ability of molten steel is shown in Figure 10.
The computational domain presents half of the
longitudinal section of the billet, which is 1.8 m long and
14 cm wide. The SEN diameter is 3.5 cm, the mold
K. MRAMOR et al.: A MESHLESS MODEL OF ELECTROMAGNETIC BRAKING FOR THE CONTINUOUS CASTING ...
964 Materiali in tehnologije / Materials and technology 49 (2015) 6, 961–967
Figure 6: Scheme of the magnetic fields facing in the same direction
Slika 6: Shema magnetnega polja tuljav, obrnjenih v enako smer
Figure 4: Scheme of EMBR
Slika 4: Shema EMBR
Figure 5: Scheme of the magnetic field for the coils facing each other
Slika 5: Shema magnetnega polja tuljav, obrnjenih druga proti drugi
Figure 7: Magnetic field of coils facing each other
Slika 7: Magnetno polje tuljav, obrnjenih druga proti drugi
Figure 3: a) Boundary and initial conditions for velocity, pressure,
turbulent kinetic energy and dissipation rate, b) boundary and initial
conditions for temperature, magnetic field and species concentration
Slika 3: a) Robni in za~etni pogoji za hitrost, tlak, turbulentno kine-
ti~no energijo in hitrost disipacije, b) robni in za~etni pogoji za
temperaturo, magnetno polje in koncentracijo
(a) (b)
height is 0.8 m and the coil height is 10 cm. The mag-
netic field is calculated for a coil configuration with 11
windings in the y direction and 25 windings in the x
direction for coils placed 0.05 m away from the strand.
The coils are placed just below the mold. A direct
current with an amplitude of 50 A runs through the coils.
Normally, the material properties of steel are temperature
dependent. However, for the purpose of this simplified
model, constant values are used for each of the phases.
The values are given in Table 1.
Table 1: Material properties of steel
Tabela 1: Snovne lastnosti jekla
property value
 7200 kg/m3
 30 W/(m K)
cp 700 J/(kg K)
TS 1680 K
TL 1760 K
hm 250000 J/kg
μ 0.006 Pa s
T 1·10–4 1/K
C 4·10–3 1/%
K0 6.25·109 m–1
Cref 0.008
DS 1.6·10–11 m2/s
DL 1.0·10–8 m2/s
 0.59·106/(Ω m)
K. MRAMOR et al.: A MESHLESS MODEL OF ELECTROMAGNETIC BRAKING FOR THE CONTINUOUS CASTING ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 961–967 965
Figure 11: Magnetic field of parallel coils configuration with iron
core: a) contour plot, b) top: vertical cross-section of the magnetic
field at 0.07 m, 0.125 m and 0.14 m, b) bottom: horizontal cross-sec-
tions at –0.7 m, –0.8 m, –0.85 m, –1.0 m, –1.4 m and –1.8 m
Slika 11: Magnetno polje paralelne postavitve tuljav z `eleznim
jedrom: a) konture magnetnega polja, b) zgoraj: navpi~ni prerez
magnetnega polja pri 0,07 m, 0,125 m in 0,14 m, b) spodaj: vodoravni
prerez magnetnega polja pri –0,7 m, –0,8 m, –0,85 m, –1,0 m, –1,4 m
in –1,8 m
Figure 9: Magnetic field of parallel coils configuration (coils facing
in the same direction) with an iron core
Slika 9: Magnetno polje paralelne postavitve tuljav (tuljave obrnjene v
enako smer) z `eleznim jedrom
Figure 8: Magnetic field of coils facing in the same direction
Slika 8: Magnetno polje tuljav, obrnjenih v enako smer
Figure 10: Temperature dependence of permeability for steel
Slika 10: Temperaturna odvisnost permeabilnosti jekla
(a) (b)
4.2 EMBR for the continuous casting of steel
First the magnetic field in the strand for a default coil
configuration with 25 windings in the x direction, 11
windings in the y direction, an electric current of 50 A,
and a span distance of 0.05 m is calculated. The results
are shown in Figure 11. The effect of this magnetic field
is then investigated for the velocity, temperature and con-
centration fields. To better present the effect of the
magnetic field the results are compared to the example
without a magnetic field, as can be seen in Figures 12 to
14.
The effect of magnetic field on temperature is shown
in Figure 12. The application of EMBR to the conti-
nuous casting of steels lowers the temperature through-
out the mold. The lowering of the temperatures is the
most apparent in the mold region.
In Figure 13 a contour plot of the velocity field is
shown together with several representative, both vertical
and horizontal, cross-sections. The calculations confirm
that the application of a magnetic field affects the velo-
city field. In the case of a parallel coil arrangement of the
EMBR, the magnetic field slows down the velocity and
diminishes the recirculation zones.
Finally, the effect of applying the EMBR to the con-
tinuous casting of steel is investigated for the con-
centration field. The results of the calculations are shown
in Figure 14, from which it can be confirmed that the
K. MRAMOR et al.: A MESHLESS MODEL OF ELECTROMAGNETIC BRAKING FOR THE CONTINUOUS CASTING ...
966 Materiali in tehnologije / Materials and technology 49 (2015) 6, 961–967
Figure 14: a) The concentration contour plots for configuration
without (left) and with (right) magnetic field, b) top: comparison of
vertical cross-sections for concentration with and without magnetic
field at 0.07 m, 0.125 m and 0.14 m, b) bottom: comparison of
horizontal cross-sections for concentration with and without magnetic
field at 0.8 m, 0.9 m and 1.8 m
Slika 14: a) Konture koncentracijskega polja za konfiguracijo z
magnetnim poljem (levo) in brez njega (desno), b) zgoraj: primerjava
navpi~nih prerezov koncentracije z magnetnim poljem in brez njega
pri 0,07 m, 0,125 m in 0,14 m, b) spodaj: primerjava vodoravnih
prerezov koncentracije z magnetnim poljem in brez njega pri 0,8 m,
0,9 m in 1,8 m
Figure 13: a) The temperature contour plots for configuration without
(left) and with (right) magnetic field, b) top: comparison of vertical
cross-sections for temperatures with and without magnetic field at
0.07 m, 0.125 m and 0.14 m, b) bottom: comparison of horizontal
cross-sections for temperatures with and without magnetic field at 0.8
m, 0.9 m and 1.8 m
Slika 13: a) Konture temperaturnega polja za konfiguracijo z mag-
netnim poljem (levo) in brez njega (desno), b) zgoraj: primerjava
navpi~nih prerezov temperature z magnetnim poljem in brez njega pri
0,07 m, 0,125 m in 0,14 m, b) spodaj: primerjava vodoravnih prerezov
temperature z magnetnim poljem in brez njega pri 0,8 m, 0,9 m in 1,8
m
Figure 12: a) The velocity contour plots for configuration without
(left) and with (right) a magnetic field, b) top: comparison of vertical
cross-sections for velocities with and without magnetic field at 0.07
m, 0.125 m and 0.14 m, b) bottom: comparison of horizontal
cross-sections for velocities with and without magnetic field at 0.8 m,
0.9 m and 1.8 m
Slika 12: a) Konture hitrostnega polja za konfiguracijo z magnetnim
poljem (levo) in brez njega (desno), b) zgoraj: primerjava navpi~nih
prerezov hitrosti z magnetnim poljem in brez njega pri 0,07 m, 0,125
m in 0,14 m, b) spodaj: primerjava vodoravnih prerezov hitrosti z
magnetnim poljem in brez njega pri 0,8 m, 0,9 m in 1,8 m
magnetic field affects the concentration. In present case,
the binary mixture of carbon and iron is investigated for
the mass fraction 0.08 % of carbon. It is shown that the
magnetic field affects the pattern of the segregation in
such a way that the levels of carbon are slightly de-
creased in the outer layers of the strand and increased in
the middle of the strand.
5 CONCLUSIONS
In this paper, numerical calculations for electromag-
netic braking in the continuous casting of steel are pre-
sented. The results, calculated with LRBFCM method,
confirm that the application of a magnetic field affects
the velocity of the fluid flow, as well as the temperature
and species concentration. The present configuration of
the coils produces a magnetic field that effectively slows
down the velocity of the flow and decreases the
temperature. It also affects the pattern of segregation in
such a way that the concentration of carbon is decreased
in the middle of the strand. In the future, an alternating
magnetic field will be applied in order to calculate the
electromagnetic stirring.
Acknowledgements
The research in this paper was sponsored by Centre
of Excellence for Biosensors, Instrumentation and
Process Control (COBIK) and Slovenian Grant Agency
under Programme group P2-0357: Modelling and Simu-
lation of Materials and Processes and applied project
L2-6775 Modelling of Industrial Solidification Systems
under Influence of Electromagnetic Field. This paper
forms a part of the doctoral study of the first author that
is partly co-financed by the European Union and by the
European Social Fund respectively. The co-financing is
carried out within the Human resources development
operational programme for years 2007–2013, 1. Deve-
lopmental priorities: Encouraging entrepreneurship and
adaptation; Preferential directives 1.3: Scholarship
schemes.
6 REFERENCES
1 W. R. Irwing, Continuous Casting of Steel, The Institute of Mater-
ials, London 1993
2 Steel Statistical Yearbook 2013, World Steel Association, Brussels
2013
3 H. K. Versteeg, An Introduction to Computational Fluid Dynamics:
the Finite Volume Method, Pearson Education, India 1995
4 B. Zhao, B. G. Thomas, S. P. Vanka, R. J. O’Malley, Metallurgical
and Materials Transactions B, 36 (2005) 6, 801–823, doi:10.1007/
s11663-005-0083-3
5 M. R. Aboutalebi, M. Hasan, R. I. L. Guthrie, Metallurgical and Ma-
terials Transactions B, 26 (1995) 4, 731–744, doi:10.1007/
bf02651719
6 D. S. Kim, W. S. Kim, K. H. Cho, Iron and Steel Institute of Japan
International, 40 (2000) 7, 670–676, doi:10.2355/isijinternational.
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461–481, doi:10.1080/10407780590911639
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9 C. H. Moon, S. M. Hwang, International Journal for Numerical Me-
thods In Engineering, 57 (2003), 315–339, doi:10.1002/nme.679
10 B. [arler, R. Vertnik, Computers and Mathematics with Applications,
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11 D. C. Wilcox, Turbulence modeling for CFD, DCW Industries, Inc.,
California 1993
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Mass Transfer, 37 (1994), 139–151, doi:10.1016/0017-9310(94)
90168-6
14 B. [arler, R. Vertnik, K. Mramor, IOP Conference Series: Materials
Science and Engineering, 33 (2012), 12012–12021, doi:10.1088/
1757-899x/33/1/012012
15 A. Chorin, Mathematical Computation, 22 (1968), 745–762
16 M. D. Buchmann, Radial Basis Function: Theory and Implementa-
tions, Cambridge University Press, Cambridge 2003, doi:10.1017/
cbo9780511543241
17 R. Franke, Mathematics of Computation, 38 (1982), 181–200,
doi:10.1090/s0025-5718-1982-0637296-4
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neering & Science, 92 (2013) 4, 327–352, doi:10.3970/cmes.2013.
092.327
19 K. Mramor, Modelling of Continuous Casting of Steel under the In-
fluence of Electromagnetic Field with Meshless Method, disser-
tation, UNG, Nova Gorica, 2014, p. 244
20 K. Mramor, R. Vertnik, B. [arler, Engineering Analysis with Boun-
dary Elements, 49 (2014), 37–47, doi:10.1016/j.enganabound.
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K. MRAMOR et al.: A MESHLESS MODEL OF ELECTROMAGNETIC BRAKING FOR THE CONTINUOUS CASTING ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 961–967 967

Q. LIU, B. [ARLER: NON-SINGULAR METHOD OF FUNDAMENTAL SOLUTIONS FOR THREE–DIMENSIONAL ...
969–974
NON-SINGULAR METHOD OF FUNDAMENTAL SOLUTIONS
FOR THREE–DIMENSIONAL ISOTROPIC ELASTICITY
PROBLEMS WITH DISPLACEMENT BOUNDARY CONDITIONS
NESINGULARNA METODA FUNDAMENTALNIH RE[ITEV ZA
DEFORMACIJO TRIDIMENZIJSKIH ELASTI^NIH PROBLEMOV Z
DEFORMACIJSKIMI ROBNIMI POGOJI
Qingguo Liu1, Bo`idar [arler1,2
1University of Nova Gorica, Vipavska 13, 5000 Nova Gorica, Slovenia
2Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
Qingguo.Liu@ung.si, bozidar.sarler@imt.si
Prejem rokopisa – received: 2015-04-23; sprejem za objavo – accepted for publication: 2015-10-09
doi:10.17222/mit.2015.086
The purpose of the present paper is to develop the Non-Singular Method of Fundamental Solutions (NMFS) based on the
boundary-distributed source method for three-dimensional elasticity problems with displacement boundary conditions. In the
NMFS, the source points and the collocation points coincide and both are positioned on the boundary of the problem domain. In
this case, the fundamental solution is singular. In order to remove the singularities of the fundamental solution, the concentrated
point sources are replaced by the distributed sources over the sphere around the singularity. The values of the distributed sources
are calculated directly in the case of displacement boundary conditions for isotropic problems. The performance of the novel
approach is shown on two three-dimensional elastic problems with displacement boundary conditions. The method requires the
discretization of the boundary only and shows excellent accuracy. It represents an efficient alternative to the classic numerical
methods. The developments lead to the possibility of modelling micromechanical problems without the discretization of the
interor of each of the grains, like required in classic numerical methods.
Keywords: linear isotropic elasticity, non-singular method of fundamental solutions, boundary meshless method
Namen ~lanka je razvoj nesingularne metode fundamentalnih re{itev (NMFS) na podlagi robno distribuirane metode izvirov za
tridimenzijske probleme linearne elasti~nosti z deformacijskimi robnimi pogoji. V NMFS se izvirne in kolokacijske to~ke
skladajo in so pozicionirane na robu obravnavanega obmo~ja. V tem primeru je fundamentalna re{itev singularna. Za odstranitev
singularnosti fundamentalne re{itve so koncentrirani izviri nadome{~eni s porazdeljenimi izviri po krogli okoli singularnosti.
Vrednosti porazdeljenih izvirov so neposredno izra~unane pri Dirichletovih robnih pogojih za izotropne probleme. Zna~ilnosti
novega na~ina so prikazane na dveh primerih tridimenzijskih problemov z deformacijskimi robnimi pogoji. Metoda zahteva
zgolj diskretizacijo roba in prikazuje odli~no natan~nost. Pomeni tudi u~inkovito alternativo klasi~nim numeri~nim metodam.
Opisani razvoj vodi do mo`nosti simulacije mikromehanskih problemov brez diskretizacije notranjosti zrn, kot je to potrebno pri
klasi~nih numeri~nih metodah.
Klju~ne besede: linerna izotropna elasti~nost, nesingularna metoda fundamentalnih re{itev, robna brezmre`na metoda
1 INTRODUCTION
The main idea of MFS1 consists of approximating the
solution of the partial differential equation by a linear
combination of fundamental solutions, defined in source
points. The expansion coefficients are calculated by
collocation or a least-squares fit of the boundary
conditions. The fundamental solution is usually singular
in the source points and this is the reason why the source
points are located outside the domain in the MFS. In this
case, the original problem is reduced to determining the
unknown coefficients of the fundamental solutions and
the coordinates of the source points by requiring the
approximation to satisfy the boundary conditions and
hence solving a non-linear problem. If the source points
are a priori fixed, then the coefficients of the MFS appro-
ximation are determined by solving a linear problem.
The MFS has become very popular in recent years
because of its simplicity2–5 and for 3D problems.6,7
In the traditional MFS, a fictitious boundary, po-
sitioned outside the problem domain, is required to place
the source points. This is very impractical or even impos-
sible, particularly when solving muti-body problems. In
recent years, various efforts have been made, with the
aim being to remove this barrier in the MFS, so that the
source points can be placed on the real boundary
directly8–12 In the present paper, we use a Non-Singular
MFS based on8 to deal with the three-dimensional iso-
tropic elasticity problems with displacement boundary
condition. The application of a non-singular method of
fundamental solutions (NMFS) in two-dimensional iso-
tropic and anisotropic linear elasticity has been origi-
nally developed.13–15 We respectively used area-distri-
buted sources covering the source points to replace the
concentrated point sources. This NMFS approach also
does not require any information about the neighboring
points for each source point, thus it is a truly a meshfree
Materiali in tehnologije / Materials and technology 49 (2015) 6, 969–974 969
UDK 519.61/.64:539.3 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)969(2015)
boundary method. The present develoments are dedi-
cated to enabling NMFS for solving three-dimensional
micromechanical elasticity problems. This is of utmost
importance in the simulation of an effective Young’s
modulus and Poisson’s ratio for multigrain systems that
appear in many engineering systems.
The rest of the paper is structured as follows. The
governing equations are shown in matrix form. The
solution procedure is given for MFS and NMFS. A
three-dimensional example in two cases, translation and
deformation, is given, followed by the conclusions and
future research.
2 GOVERNING EQUATIONS
Consider a 3D domain  with the boundary  filled
with isotropic elasticity materials. Let us introduce a 3D
Cartesian coordinate system with the orthonormal base
vectors ix, iy and iz and the coordinates px, py and pz of the
position vector p, i.e., p = px ix+ py iy+ pz iz. To simplify
the calculations we shall assume that (i) the solid is free
of body forces and (ii) the thermal strains can be ne-
glected. Under these conditions the general equation of
elasticity16 is:
C
u
p p
x y z !"
"
 
 # !# "#  # #
∂
∂ ∂
2
0
( )
,
p
= (1)
where u" are the displacements, C !" are the elastic
stiffnesses and the components of a fourth rank stiffness
tensor:17
C =C
C C C C C C
C C C
xxxx xxyy xxzz xxyz xxxz xxxy
xxyy yyyy yy
 !"
zz yyyz xzyy xyyy
xxzz yyzz zzzz yzzz xzzz xyzz
xx
C C C
C C C C C C
C yz yyyz yzzz yzyz xzyz xyyz
xxxz xzyy xzzz xzyz xz
C C C C C
C C C C C xz xyxz
xxxy xyyy xyzz xyyz xyxz xyxy
C
C C C C C C
⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
⎥
⎥
=
=
c c c c c c
c c c c c c
c c c
11 12 13 14 15 16
12 22 23 24 25 26
13 23 33 c c c
c c c c c c
c c c c c c
c c
34 35 36
14 24 34 44 45 46
15 25 35 45 55 56
16 26 36 46 56 66c c c c
⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
⎥
⎥
(2)
In subsequent discussions, it will be convenient to
write the equilibrium Equation (1) in matrix form as:
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
p p p
p p p
p p p
x y z
y x z
z x y
0 0 0
0 0 0
0 0 0
⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
⎥
⎥
⎥
⋅
⋅
c c c c c c
c c c c c c
11 12 13 14 15 16
12 22 23 24 25 26
13 23 33 34 35 36
14 24 34 44 45 46
15 25 35 45
c c c c c c
c c c c c c
c c c c c55 56
16 26 36 46 56 66
c
c c c c c c
u / px x⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
⎥
⎥
⎥
⋅
∂ ∂
∂u / p
u / p
u / p u / p
u / p u / p
u / p
y y
z z
y z z y
x z z x
x y
∂
∂ ∂
∂ ∂ ∂ ∂
∂ ∂ ∂ ∂
∂ ∂
+
+
+
⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
⎥
⎥
⎥
=
∂ ∂u / py x
0
(3)
The stresses  ! are related to the strains through the
generalized Hooke’s law:
s = Ce (4)
where C !" satisfy the fully symmetrical conditions:
C C !" ! "= , C C !"  !"= , C C !" " != (5)
e is the strains vector:
e ≡
⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
⎥
⎥
=






xx
yy
zz
yz
xz
xy
x x
y
u / p
u
∂ ∂
∂ / p
u / p
u / p u / p
u / p u / p
u / p
y
z z
y z z y
x z z x
x y
∂
∂ ∂
∂ ∂ ∂ ∂
∂ ∂ ∂ ∂
∂ ∂ ∂
+
+
+ u / py x∂
⎡
⎣
⎢
⎢
⎢
⎢
⎢
⎢
⎤
⎦
⎥
⎥
⎥
⎥
⎥
⎥
(6)
3 SOLUTION PROCEDURE
The fundamental solution for the isotropic elasticity
is given18 in three dimensions (3D) by:
U
v r
v
p s p s
r
 !
 !
  ! !


$
( , )
( )
( )
( )( )
p s =
−
⋅
− +
− −⎧
⎨
⎩
1
16 1
3 4
π
⎫
⎬
⎭
,  # ! # #x y z
(7)
where U !(p,s) represents the displacement in the direc-
tion  at point p due to a unit point force acting in the direc-
tion ! at point s. r = [(px – sx)2 + (py – sy)2 + (pz – sz)2]1/2
is the distance between the point p and the source point
s. Equation (7) is expanded as follows:
Q. LIU, B. [ARLER: NON-SINGULAR METHOD OF FUNDAMENTAL SOLUTIONS FOR THREE–DIMENSIONAL ...
970 Materiali in tehnologije / Materials and technology 49 (2015) 6, 969–974
U
v r
v
p s
r
U
xx
x x
yy
=
−
− +
−⎡
⎣⎢
⎤
⎦⎥
=
−
1
16 1
3 4
1
16 1
2
π
π


( )
( )
( )
( v r
v
p s
r
U
v r
v
p
y y
zz
)
( )
( )
( )
( )
(
3 4
1
16 1
3 4
2
− +
−⎡
⎣
⎢
⎤
⎦
⎥
=
−
− +

π
z z
xy yx
x x y y
s
r
U U
v r
p s p s
r
U
−⎡
⎣⎢
⎤
⎦⎥
= =
−
− −
)
( )
( )( )
2
1
16 1

π
xz zx
x x z z
yz zy
U
v r
p s p s
r
U U
= =
−
− −
= =
−
1
16 1
1
16 1
π
π


( )
( )( )
( v r
p s p s
r
y y z z
)
( )( )− −

(8)
It can be shown that the following ux, uy and uz satisfy
the governing Equations (3):
u U U Ux xx xy xz( ) ( , ) ( , ) ( , )p p s p s p s= +   (9)
u U U Uy yx yy yz( ) ( , ) ( , ) ( , )p p s p s p s= +   (10)
u U U Uz zx zy zz( ) ( , ) ( , ) ( , )p p s p s p s= +   (11)
where ,  and  represent arbitrary constants. The fun-
damental solution U !(p,s) is singular when p = s. We
use the desingularization technique, proposed by Liu8
for evaluating the singular values. We modify his
approach in a sense of preserving the original funda-
mental solution at all the points except the singularity,
and by scaling the singularity with the area of the sphere
over which the desingularization integration is per-
formed. This allows us to treat the MFS and the NMFS
in formally the same way. The desingularization (trans-
formation of U !(p,s) into
~
U !(p,s)) is thus performed in
the following way:
~
( , )
( , )
( , )
( , )
U
U r R
R
U A r R
A R
 !
 !
 !
p s
p s
p s
s
=
>
≤
⎧
⎨
⎪
⎩⎪
∫
1
2π
d (12)
where A(s,R) represents a sphere with radius R, centered
around s. The involved integrals can be calculated as
follows (by using the integration in polar coordinates
px – sx = r sin cos
, py – sy = r sin sin
 and pz – sz =
r cos, Figure 1):
~
( , )
~
( , )
~
( , )
( )
~
(
U U U
v
v R
U
xx yy zz
xy
p p p p p p
p
= = =
−
−
5 6
16 1π
, )
~
( , )
~
( , )
~
( , )
~
( , )
~
p p p
p p p p
p p
= =
= =
=
U
U U
U U
yx
xz zx
yz zy
0
0
( , )p p = 0
(13)
It can also be shown that the following ux, uy and uz
satisfy the governing Equations (3):
u U U Ux xx xy xz( )
~
( , )
~
( , )
~
( , )p p s p s p s= +   (14)
u U U Uy yx yy yz( )
~
( , )
~
( , )
~
( , )p p s p s p s= +   (15)
u U U Uz zx zy zz( )
~
( , )
~
( , )
~
( , )p p s p s p s= +   (16)
The solution of the problem is sought in the form:
u U U
U
x xx n
n
N
n xy n
n
N
n
xz
( )
~
( , )
~
( , )
~
( ,
p p p p p
p
= +
= =
∑ ∑
1 1
  
 p n
n
N
n)
=
∑
1

(17)
u U U
U
y yx n
n
N
n yy n
n
N
n
yz
( )
~
( , )
~
( , )
~
( ,
p p p p p
p
= +
= =
∑ ∑
1 1
  
 p n
n
N
n)
=
∑
1

(18)
u U U
U
z zx n
n
N
n zy n
n
N
n
zz
( )
~
( , )
~
( , )
~
( ,
p p p p p
p
= +
= =
∑ ∑
1 1
  
 p n
n
N
n)
=
∑
1

(19)
The coefficients n, n and n are calculated from a
system of 3N algebraic equations:
Ax = b (20)
where A stands for a 3N × 3N matrix with the entries
Aij, x is a 3N × 1 vector with the entries xi, and b is a 3N
× 1 vector with entries bi:
A U A U
A U
ij xx i j i N j xy i j
i N j
= =
=
+
+
~
( , )
~
( , )
~
( )
( )
p p p p,
2 xz i j N i j yx i j
N i N j yy
A U
A U
( , )
~
( , )
~
(
( )
( )( )
p p p p, +
+ +
=
= p p p p
p
i j N i N j yz i j
N i j zx i
A U
A U
, )
~
( , )
~
( ,
( )( )
( )
, + +
+
=
=
2
2 p p pj N i N j zy i j
N i N j zz
A U
A U
)
~
( , )
~
( )( )
( )( )
, 2
2 2
+ +
+ +
=
= ( , )p pi j i,j= , , ..., N, 12
(21)
x x x i= , , ..., Ni i N i i N i i= = =+ +  , ,( ) ( ) ,2 1 2 (22)
b u b u b u
i N
i x i N i y i N i z i= = =
=
+ +( ), ( ), ( ),
, , ...,
( ) ( )p p p2
1 2
(23)
Q. LIU, B. [ARLER: NON-SINGULAR METHOD OF FUNDAMENTAL SOLUTIONS FOR THREE–DIMENSIONAL ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 969–974 971
Figure 1: Distributed source on a sphere A(s,R) with radius R
Slika 1: Porazdeljeni izviri na krogli A(s,R) z radijem R
By knowing all the elements Aij and bi of the system
(20), we can determine the values of xi (i.e., n, n and
n). Afterwards, we can calculate the solution of the
governing equation from:
u U U
U
x n
n
N
n y n
n
N
n
z
   
 
 ( )
~
( , )
~
( , )
~
( ,
p p p p p
p
= + +
+
= =
∑ ∑
1 1
p n
n
N
n x y z) , ,
=
∑ =
1
  ,
(24)
where p is any point inside the domain or on the boun-
dary.
4 NUMERICAL EXAMPLES
We consider a cube with the side length a = 2 m
centered around px = 0 m, py = 0 m, pz = 0 m. The elastic
media is defined by E = 1 N/m2, v = 0.3.
4.1 Translation
We consider a solution of the governing equations in
this cube subject to the boundary conditions ux = 2 m, uy
= 2 m, uz = 2 m. The analytical solution is:
ux = 2 m, uy = 2 m, uz = 2 m, (25)
A plot of the translation, obtained with the analytical
solution and the numerical solutions with MFS and
NMFS, is shown in Figure 2 for the case with 150 nodes
(25 nodes on each side of he cube). The distance of the
fictitious boundary from the true boundary for the MFS
is set RM = 5d, where d is the smallest distance between
two nodes on the boundary. The radius of the sphere for
the distributed area source covering each node is set to R
= d/3.
The solution of the points on a square with the side
length a = 1 m centered around px = 0 m, py = 0 m, pz = 0
m on the plane pz = 0 are computed and compared with
the analytical solutions. The root-mean-square (RMS)
errors of the numerical solution are defined as:
e
N
u u x yn n
n
N
   
 = − =
=
∑1 2
1
( ) ,, (26)
Q. LIU, B. [ARLER: NON-SINGULAR METHOD OF FUNDAMENTAL SOLUTIONS FOR THREE–DIMENSIONAL ...
972 Materiali in tehnologije / Materials and technology 49 (2015) 6, 969–974
Figure 2: The analytical solution and the numerical solution of MFS
and NMFS for the translation case with N = 150, R = d/3, RM = 5d
(•: collocation points, +: analytical solution, x: MFS solution, :
NMFS solution)
Slika 2: Analiti~na in numeri~na re{itev z MFS in NMFS za transla-
cijski primer z N = 150, R = d/3, RM = 5d (•: kolokacijske to~ke, +:
analiti~na re{itev, x: MFS re{itev, : NMFS re{itev)
Figure 3: The relationship between the RMS errors and the number of
boundary nodes for translation case, calculated by NMFS. R = d/3 (+:
ex, x: ey, : ez).
Slika 3: Odvisnost med RMS-napakami in {tevilom robnih to~k za
translacijski primer, izra~unan z NMFS. R = d/3 (+: ex, x: ey, : ez).
Table 1: RMS errors of NMFS solutions for the translation case with
R = d/3
Tabela 1: RMS-napake NMFS-re{itev za translacijski primer z R =
d/3
Num. of boun-
dary nodes (N) ex(× 10
–3) ey(× 10–3) ez(× 10–3)
150 1.2200 1.2200 0.8390
216 0.8769 0.8769 0.6109
294 0.6570 0.6570 0.4658
384 0.5112 0.5112 0.365
486 0.4091 0.4091 0.2950
600 0.3348 0.3348 0.2428
726 0.2791 0.2791 0.2033
864 0.2363 0.2363 0.1727
1014 0.2026 0.2026 0.1485
1176 0.1757 0.1757 0.1291
1350 0.1538 0.1538 0.1132
1536 0.1357 0.1357 0.1001
1734 0.1207 0.1207 0.0891
1944 0.1080 0.1080 0.0799
2166 0.0973 0.0973 0.0720
2400 0.0880 0.0880 0.0652
2646 0.0800 0.0800 0.0594
2904 0.0731 0.0731 0.0543
3174 0.0670 0.0670 0.0498
3456 0.0617 0.0617 0.0459
where u k and u k( = x, y) are the analytical and the
numerical solutions, respectively. The number of boun-
dary nodes used is from 150 to 3 456.
Figure 3 shows the RMS errors of the results ob-
tained using the NMFS. The errors are already less than
10–3 with N = 216 and the solution converges to the
analytical solution with an increasing number of nodes
(Table 1). The MFS result is shown in Table 2 for RM =
5d. Here it should be noted that the MFS solution error is
relatively small; however, the convergence is not
uniform. This fact is due to the choice of the artificial
boundary position, which was for all node arrangements
RM = 5d and thus most probably not optimally varying.
Table 2: RMS errors of MFS solutions for the translation case with
RM = 5d
Tabela 2: RMS-napake MFS-re{itev za translacijski primer z RM = 5d
Num. of boun-
dary nodes (N) ex(× 10
–14) ey(× 10–14) ez(× 10–14)
150 0.2204 0.2204 0.5548
216 1.7907 1.7907 4.8854
294 0.0364 0.0364 0.0794
384 0.1590 0.1590 0.1617
486 0.0348 0.0348 0.0017
600 0.0058 0.0058 0.0015
726 0.1103 0.1102 0.1631
864 0.0004 0.0004 0.0001
1014 0.0475 0.0445 0.0478
1176 0.0033 0.0050 0.0025
1350 0.0005 0.0003 0.0010
1536 0.0038 0.0295 0.0238
1734 0.0000 0.0000 0.0000
1944 0.0000 0.0000 0.0000
2166 0.0004 0.0004 0.0006
2400 0.0000 0.0000 0.0000
2646 0.0000 0.0000 0.0000
2904 0.0000 0.0000 0.0000
3174 0.0000 0.0001 0.0000
3456 0.0000 0.0000 0.0000
4.2 Deformation
We consider a solution of the governing equations in
this cube subject to the boundary conditions ux = px, uy =
py, uz = pz. The analytical solution is:
ux = px, uy = py, uz = pz (27)
A plot of the deformation, obtained with the analyti-
cal solution and the numerical solutions with MFS and
NMFS, is shown in Figure 4 for the case with 150 nodes
Q. LIU, B. [ARLER: NON-SINGULAR METHOD OF FUNDAMENTAL SOLUTIONS FOR THREE–DIMENSIONAL ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 969–974 973
Figure 4: The analytical solution and the numerical solution of MFS
and NMFS for the deformation case with N = 150, R = d/3, RM = 5d
(•: collocation points, +: analytical solution, x: MFS solution, :
NMFS solution)
Slika 4: Analiti~na in numeri~na re{itev z MFS in NMFS za defor-
macijski primer z N = 150, R = d/3, RM = 5d (•: kolokacijske to~ke, +:
analiti~na re{itev, x: MFS re{itev, : NMFS re{itev)
Figure 5: The relationship between the RMS errors and the number of
boundary nodes for the deformation case, calculated by NMFS. R=d/3
(+: ex, x: ey, : ez).
Slika 5: Odvisnost med RMS-napakami in {tevilom robnih to~k za de-
formacijski primer, izra~unan z NMFS. R = d/3 (+: ex, x: ey, : ez).
Table 3: RMS errors of the NMFS solutions for the deformation case
with R=d/3
Tabela 3: Odvisnost med RMS-napakami in {tevilom robnih to~k za
deformacijski primer, izra~unan z NMFS, R = d/3
Num. of boun-
dary nodes (N) ex(× 10
–3) ey(× 10–3) ez(× 10–3)
150 4.1487 4.1487 0.0000
216 3.1826 3.1826 0.0000
294 2.4837 2.4837 0.0000
384 1.9972 1.9972 0.0000
486 1.6395 1.6395 0.0000
600 1.3703 1.3703 0.0000
726 1.1623 1.1623 0.0000
864 0.9983 0.9983 0.0000
1014 0.8667 0.8667 0.0000
1176 0.7596 0.7596 0.0000
1350 0.6711 0.6711 0.0000
1536 0.5973 0.5973 0.0000
1734 0.5350 0.5350 0.0000
1944 0.4820 0.4820 0.0000
2166 0.4365 0.4365 0.0000
2400 0.3971 0.3971 0.0000
2646 0.3629 0.3629 0.0000
2904 0.3329 0.3329 0.0000
3174 0.3064 0.3064 0.0000
3456 0.2830 0.2830 0.0000
(25 nodes on each side of the cube). The same R and RM
as with example 4.1 are used.
Figure 5 shows the RMS errors of the results ob-
tained using the NMFS and the solution converges to the
analytical solution with an increasing number of nodes
(Table 3). The MFS results are shown in Table 4 for RM
= 5d.
Table 4: RMS errors of the MFS solutions for the deformation case
with RM = 5d
Tabela 4: RMS-napake MFS-re{itev za deformacijski primer RM = 5d
Num. of boun-
dary nodes (N) ex(× 10
–11) ey(× 10–11) ez(× 10–11)
150 0.1410 0.1410 0.0000
216 0.0256 0.0256 0.0000
294 0.0018 0.0018 0.0000
384 0.0012 0.0012 0.0000
486 0.0014 0.0014 0.0000
600 2.1088 2.1088 0.0000
726 0.0010 0.0010 0.0000
864 0.0001 0.0001 0.0000
1014 0.0000 0.0000 0.0000
1176 0.0000 0.0000 0.0000
1350 0.0010 0.0008 0.0019
1536 0.0000 0.0002 0.0001
1734 0.0000 0.0000 0.0000
1944 0.0002 0.0001 0.0001
2166 0.0005 0.0005 0.0008
2400 0.0000 0.0000 0.0000
2646 0.0000 0.0000 0.0000
2904 0.0000 0.0000 0.0000
3174 0.0000 0.0000 0.0000
3456 0.0000 0.0000 0.0000
5 CONCLUSION
A new, non-singular method of fundamental solu-
tions13 is extended in the present paper to solve 3D linear
elasticity problems. In this approach, the singular values
of the fundamental solution are integrated over a small
sphere, so that the coefficients in the system of equations
can be evaluated analytically and consistently, leading to
an extremely simple computer implementation of this
method. The method essentially gives similar results as
the classic MFS. It has the advantage that the artificial
boundary is not present; however, the problems with the
traction boundary condition have not yet been solved.
The main advantage of the method is that the discre-
tisation is performed only on the boundary of the domain
and no polygonisation is needed, like in the finite-ele-
ment method. The NMFS, presented in this paper, can be
adapted or extended to handle many related problems,
such as anisotropic elasticity, and multi-body problems,
which all represent directions for our further investi-
gations. The advantage of not having to generate the
artificial boundary is particularly welcome in these types
of problems. The method will be used in the future for
the calculation of 3D enginering deformation problems
in steel and aluminium alloys, with realistic grain shapes,
obtained from microscope images. The developed
method is believed to represent the simplest state-of-
the-art way to numerically cope with these types of
problems.
Acknowledgement
This paper forms a part of the project L2-6775 Simu-
lation of industrial solidification proceesses under
influence of electromagnetic fieleds. This work was par-
tially performed within the Creative Core program
(AHA-MOMENT) contract no. 3330-13-500031, co-
supported by RSMIZS and European Regional Develop-
ment Fund Research.
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974 Materiali in tehnologije / Materials and technology 49 (2015) 6, 969–974
M. FEIZPOUR et al.: SOLID-STATE SINTERING OF (K0.5Na0.5)NbO3 SYNTHESIZED FROM ...
975–982
SOLID-STATE SINTERING OF (K0.5Na0.5)NbO3 SYNTHESIZED
FROM AN ALKALI-CARBONATE-BASED LOW-TEMPERATURE
CALCINED POWDER
SINTRANJE V TRDNEM KERAMIKE (K0,5Na0,5)NbO3,
SINTETIZIRANE IZ NIZKOTEMPERATURNO KALCINIRANEGA
PRAHU, PRIPRAVLJENEGA NA OSNOVI ALKALIJSKIH
KARBONATOV
Mahdi Feizpour1, Touradj Ebadzadeh1, Darja Jenko2
1Department of Ceramic Engineering, Materials and Energy Research Center, P.O. Box 31787-316, Karaj, Alborz, Iran
2Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
m-feizpour@merc.ac.ir, iusteng@gmail.com
Prejem rokopisa – received: 2015-10-10; sprejem za objavo – accepted for publication: 2015-10-19
doi:10.17222/mit.2015.315
Potassium sodium niobate K0.5Na0.5NbO3 (KNN) was synthesized by the double calcination of a homogenized mixture of
potassium and sodium carbonates and niobium pentoxide for 4 h at 625 °C. The calcination temperature was chosen on the basis
of the thermal analyses of the mixture of precursors, where the weight loss being the function of the temperature reaches the
plateau. The calcined powder was investigated by X-ray Diffraction (XRD) and Transmission Electron Microscopy (TEM) and
was found to be without unreacted materials or secondary phases. Before sintering, the powder compacts were annealed for 4 h
at 450 °C, while the sintering was carried out for 2 h at 1115 °C using two different configurations: 1) in a closed crucible where
the KNN pellets were in close physical proximity to, but not in direct contact with, the KNN packing powder, and 2) in a
completely open crucible without any packing powder. The Archimedes’ density of the sintered samples was 91.5 % of
theoretical density for the first configuration, while it was 93.4 % for the second configuration. The Field-Emission Scanning
Electron Microscopy (FE-SEM) and XRD analyses of the sintered ceramics showed that by using a calcination temperature as
low as 625 °C a typical sintered microstructure of KNN could be achieved with both sintering configurations.
Keywords: potassium sodium niobate, low-temperature calcination, solid-state synthesis, sintering, microstructure
Keramiko kalij-natrijevega niobata K0,5Na0,5NbO3 (KNN) smo sintetizirali z dvojno kalcinacijo homogenizirane zmesi
kalijevega in natrijevega karbonata ter niobijevega oksida 4 h pri temperaturi 625 °C. Temperaturo kalicinacije smo izbrali glede
na termi~no analizo me{anice prekurzorjev, kjer se izguba mase v odvisnosti od temperature ni ve~ spreminjala in je dosegla
plato. Kalciniran prah smo preiskovali z rentgensko fazno analizo (XRD) in presevno elektronsko mikroskopijo (TEM), ki sta
pokazali, da prah ne vsebuje nezreagiranih materialov ali sekundarnih faz. Stisnjene tabletke kaciniranega prahu so bile pred
sintranjem segrevane 4 h pri 450 °C, samo sintranje pa je potekalo 2 h pri 1115 °C z uporabo dveh razli~nih konfiguracij: 1) v
zaprtem lon~ku, kjer so bile tabletke KNN v neposredni fizi~ni bli`ini, a ne v neposrednem stiku z zasipom KNN in 2) v
popolnoma odprtem lon~ku brez zasipa. Povpre~na Arhimedova gostota sintranih vzorcev je bila za prvo konfiguracijo 91,5 %
teoreti~ne gostote, medtem ko je bila za drugo konfiguracijo 93,4 %. Vrsti~na elektronska mikroskopija z emisijo polja
(FE-SEM) in analiza XRD sintrane keramike sta pokazali, da lahko `e pri ni`ji temperaturi kalciniranja 625 °C dose`emo
tipi~no sintrano mikrostrukturo KNN za obe konfiguraciji sintranja.
Klju~ne besede: kalij-natrijev niobat, nizkotemperaturna kalcinacija, sinteza v trdnem stanju, sintranje, mikrostruktura
1 INTRODUCTION
Pb-based piezoceramics are the main group of piezo-
electric materials with high electromechanical proper-
ties.1 However, with the beginning of the 21st Century,
some legislation was imposed to limit the fabrication and
usage of substances that contain lead, due to its toxicity,
and to develop environmentally friendlier replaceable
materials.2 During the past 15 years, different families of
known lead-free piezoelectrics were restudied and seve-
ral attempts were focused on obtaining functional pro-
perties close to those of the lead-based piezoelectrics.3–5
Among the different groups of lead-free piezoelectric
ceramics, the potassium sodium niobate K0.5Na0.5NbO3
[KNN] family is a widely investigated candidate for the
replacement of Pb(Zr,Ti)O3 and the other lead-based
piezoelectric ceramics, mainly due to its fair electrome-
chanical properties, high Curie point and its
compatibility with base-metal electrodes like Ni.6–14
Dealing with KNN-based materials involves some diffi-
culties, mainly concerning the preservation of the stoi-
chiometry during the powder synthesis and the sintering
of the powder compact, because of the high potential of
the alkali elements for evaporation, especially at elevated
temperatures.15,16
Up to now, numerous studies have been conducted on
maximizing the piezoelectric response of pure and/or
doped KNN-based lead-free piezoelectric ceramics.
However, since the major and cheapest source for
providing alkali elements for the synthesis of KNN is
alkali carbonates of Na2CO3 and K2CO3, the calcination
of the homogenized KNN precursors is an important step
Materiali in tehnologije / Materials and technology 49 (2015) 6, 975–982 975
UDK 666.3/.7:621.762.5 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)975(2015)
in the solid-state synthesis of KNN. The most widely
used temperatures for the calcination of KNN from the
alkali carbonates are in the range between 750 °C and
950 °C.9 But as Popovic et al.16 have thermodynamically
shown, the vapor pressure of alkali elements over the
respective niobates increases by 4 to 5 orders of mag-
nitudes when increasing the temperature from 627 °C to
927 °C. This makes the low-temperature calcination
process an easy solution to overcome the off-stoichio-
metry problem for the synthesized KNN powder due to
the evaporation of the alkali elements at elevated
temperatures. On the other hand, using lower calcination
temperatures may lead to powders with a lower
crystallinity and still-unreacted precursors. Based on the
literature and the best of our knowledge, up to now, the
lowest temperature of synthesis for KNN powder from a
Na2CO3, K2CO3 and Nb2O5 mixture was 700 °C.17 The
aim of this study is to explore whether this temperature
can be even more reduced and the sintered properties of
samples still remain good enough or not.
2 MATERIALS AND METHODS
Na2CO3 (99.95–100.05 % purity), K2CO3 (99+ %
purity) and Nb2O5 (99.9 % purity), all from Sigma-
Aldrich, Germany, were dried over night at 200 °C and
weighed according to the K0.5Na0.5NbO3 stoichiometry in
a dry-box and homogenized in a PM-400 Retsch plane-
tary mill for 4 h at 175 r/min using 125 mL grinding
zirconia jar and zirconia balls (3 mm in diameter) and
99.8 % purity acetone as the medium. The homogenized
mixture of KNN precursors (hereafter, HOM powder)
was dried for 1 h at 90 °C and at least 2 h at 200 °C and
stored in a desiccator.
Differential Thermal Analysis (DTA) and Thermo-
Gravimetry (TG) of the HOM powder were recorded
from 200 °C to 750 °C with a heating rate of 10 °C/min
by means of a Netzsch STA 409 C/CD under a constant
flow rate of 100 cm3/min of dried synthetic air. Prior to
the measurement, the sample was isothermally held for 3
h at 200 °C. In addition, the dimensional changes vs. the
temperature of the HOM and calcined powder compact
(100 MPa, uniaxial press) were recorded from 200 °C to
1200 °C with a heating rate of 5 °C/min under synthetic
air using a Leitz optical dilatometry. The X-ray Diffrac-
tion (XRD) patterns of the calcined powders and the
crushed sintered pellets were recorded at room tempera-
ture using a PANalytical X’Pert PRO MPD diffracto-
meter with Cu-K1 radiation of 0.15406 nm in the 2

range 10–90 ° with a step of 0.017 ° and an integration
time of 200 s.
The particle size distribution of the calcined powder
was analyzed with a Microtrac S3500 static light-scatter-
ing particle size analyzer. The specific surface area of the
calcined powder was analyzed by nitrogen adsorption/
desorption at –196 °C (BET method, Belsorp-mini II)
using an automated gas-adsorption analyzer. Prior to the
measurement, the powders were degassed under vacuum
for 2 h at 250 °C. The morphology, crystal structure and
elemental composition of the calcined powder were
observed and analyzed with a High-Resolution Trans-
mission Electron Microscope (HR-TEM – JEOL
JEM-2100) using an accelerating voltage 200 kV with an
attached Energy-Dispersive X-ray Spectrometer EDS
(JEOL JED-2300 Series) and a Scanning Transmission
Electron Microscope (STEM) unit with a Bright-Field
(BF) detector (EM-24511SIOD, JEOL).
After the calcination process, the powders were
compacted at 200 MPa using a cold isostatic press.
Before sintering, the powder compacts were annealed for
4 h at 450 °C in order to remove the adsorbed moisture
and CO2 from the powder before the sintering process
started. The sintering was carried out for 2 h at 1115 °C
using two different configurations: 1) in a closed,
double-crucible (high-purity alumina, total volume: 
 16
cm3) where the KNN pellets were put in a small Pt
crucible (volume: 
 2 cm3), which was surrounded by
KNN packing powder (2.5 g, after the first calcination at
800 °C), hereafter the CC/wPP configuration, and 2) in a
completely open crucible where the KNN pellets were
put in a small Pt crucible without using any packing
powder, hereafter the OC/woPP configuration. The
density of the sintered ceramics was measured based on
Archimedes’ principle using the ASTM C373 standard
and reported as an average value of three measurements.
The sintered pellets were cut, mounted, and polished
using standard ceramography techniques and finished
with 3 μm and ¼ μm diamond-paste polishes. The micro-
structure of the sintered ceramics was characterized
using a Field-Emission Scanning Electron Microscope
(FE-SEM – JEOL JSM-7600F). Prior to the analysis, the
samples were coated with a thin layer of carbon.
M. FEIZPOUR et al.: SOLID-STATE SINTERING OF (K0.5Na0.5)NbO3 SYNTHESIZED FROM ...
976 Materiali in tehnologije / Materials and technology 49 (2015) 6, 975–982
Figure 1: DTA/TG thermal analyses curves of the homogenized mix-
ture of KNN precursors (the HOM powder)
Slika 1: Krivulji termi~ne analize DTA/TG homogenizirane zmesi
prekurzorjev KNN (prah HOM)
3 RESULTS AND DISCUSSION
3.1 Synthesis of the KNN powder at a low calcination
temperature
The DTA/TG thermal analyses of the HOM powder
are shown in Figure 1. There is a pronounced exother-
mic peak around 500 °C, which is accompanied by a
sharp weight loss in this temperature region, which is
related to the decomposition of the alkali carbonates and
also to the formation of the (K,Na)2Nb4O11 intermediate
phase and the KNN perovskite phase from the pre-
cursors. However, there are still two minor exothermic
peaks at 590 °C and 625 °C, which can be attributed to
the completion of the synthesis and the final formation of
KNN.18,19 According to the TG curve, the amount of
weight loss is almost zero above 625 °C. In addition, this
temperature coincides with the last exothermic DTA
peak.
To further explore the possible low calcination tem-
perature, optical dilatometry of the HOM powder was
performed, which is shown in Figure 2. The expansion
started from 390 °C and it underwent a large expansion
of  26 % by heating the powder compact further from
495 °C to 625 °C. Again, the temperature of 625 °C be-
came an important temperature since the dimensional
change of the sample reached a plateau at 625 °C. The
dimension of the powder compact stayed constant up to
910 °C and underwent a shrinkage of  3 % during heat-
ing to 940 °C.
According to the thermal analyses of the HOM pow-
der and the dimensional changes of the HOM powder
compact, 625 °C was chosen as a possible temperature
for the calcination of the KNN precursors at low tempe-
rature. The HOM powder was calcined two times for 4 h
with a heating rate of 5 °C/min at 625 °C with an inter-
mediate and final milling step similar to the homogeni-
zation process.
XRD patterns of the KNN calcined powders after the
first and second calcinations at 625 °C are shown in Fig-
ure 3. For comparison, the patterns of the HOM powder
that was calcined for 4 h at 800 °C and the HOM powder
that was held for only 30 min at the sintering tempe-
rature of KNN, i.e., 1115 °C, are also shown in Figure 3.
It is clear that all the peaks in these patterns can be
indexed as the KNN perovskite phase and, based on the
detection limit of the XRD technique, there is no trace of
secondary phases or unreacted precursors in the patterns
of the first and second calcined KNN powders at 625 °C.
By increasing the calcination temperatures, the peaks
sharpened and split, which are indications of the more
homogenous structures and the growth of the crystallites.
The densification behavior of the double-calcined
KNN powder at 625 °C and subsequently milled is
shown in Figure 4. There is no change in the dimensions
of the powder compact until 895 °C, but similar to the
HOM powder, it again underwent a small shrinkage of 
2 % during heating from 895 °C to 925 °C. We tried to
decode the reason for such behavior at this temperature
region and it is also a topic of another of our papers that
will be published soon. However, the powder compact
underwent a large shrinkage due to the sintering process,
which started from  1090 °C and finally, it melted at
1150 °C. This narrow sintering window is typical for
KNN-based ceramics and it is not related to the
calcination temperature of the KNN.20,21 In the inset of
Figure 4, the particle size distribution of the double-cal-
cined KNN powder at 625 °C and subsequently milled is
shown. The d10, d50 and d90 of this powder are 0.19 μm,
0.41 μm and 0.82 μm, respectively, and the powder has a
unimodal distribution. The specific surface area of this
powder was 18.0 m2/g.
M. FEIZPOUR et al.: SOLID-STATE SINTERING OF (K0.5Na0.5)NbO3 SYNTHESIZED FROM ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 975–982 977
Figure 3: XRD patterns of KNN powder after: a) first calcination for
4 h at 625 °C, b) double calcination for 4 h at 625 °C, c) first
calcination for 4 h at 800 °C, d) holding for 30 min at the sintering
temperature of KNN (1115 °C)
Slika 3: Rentgenogram prahov KNN po: a) prvi kalcinaciji 4 h pri 625
°C, b) dvakratni kalcinaciji 4 h pri 625 °C, c) prvi kalcinaciji 4 h pri
800 °C in d) segrevanju 30 min pri temperaturi sintranja KNN (1115
°C)
Figure 2: Dimensional changes of the powder compact of the homo-
genized mixture of the KNN precursors (the HOM powder)
Slika 2: Spremembe dimenzij stisnjene tabletke iz homogenizirane
zmesi prekurzorjev KNN (prah HOM)
The morphology, crystallite size, crystal structure and
elemental composition of the double calcined KNN
powder at 625 °C and subsequently milled is shown in
Figure 5. The lower-magnification TEM BF image
revealed that the calcined powder was agglomerated
(Figure 5a). The powder was composed of agglomerates
of around 500–700 nm or smaller, and of individual
particles as small as 20 nm or larger (Figure 5c). A
Selected-Area Electron Diffraction (SAED) pattern,
taken from the dashed-line circled area and included as
the inset in Figure 5a, shows a characteristic diffraction
of a ring pattern with relatively continuous rings, which
means the crystallites are small, in the nm range, and in a
random orientation. The SAED pattern contains some
brighter and more distinct spots in the rings, which
indicates the presence of some larger crystallites. The
electron-diffraction spots can be described by a perov-
skite phase with indices as shown in the inset image of
Figure 5a.
M. FEIZPOUR et al.: SOLID-STATE SINTERING OF (K0.5Na0.5)NbO3 SYNTHESIZED FROM ...
978 Materiali in tehnologije / Materials and technology 49 (2015) 6, 975–982
Figure 4: Densification behavior of double-calcined KNN powder for
4 h at 625 °C and subsequently milled recorded by optical dilatometry.
Inset shows the particle size distribution of this powder.
Slika 4: Potek zgo{~evanja dvakrat kalciniranega in mletega prahu
KNN 4 h pri 625 °C, dobljen z uporabo opti~ne dilatometrije. Vstav-
ljena slika prikazuje porazdelitev velikosti delcev tega prahu.
Figure 5: a) Lower magnification TEM BF image of the morphology and crystallite size of the double-calcined KNN powder at 625 °C and
subsequently milled with the SAED pattern of the selected dashed-line circled area included as the inset, and b) typical EDS spectrum from the
same selected dashed-line circled area in Figure a. c) Higher magnification TEM BF and d) HR-TEM images of this powder. From the selected
dashed-line circled area of the TEM BF image, the HR-TEM image was taken. From the dashed-line circled area in the HR-TEM image, the 2D
FFT and simulation image were calculated and included as the insets.
Slika 5: a) Posnetek TEM v svetlem polju (BF) pri nizki pove~avi prikazuje morfologijo in velikost kristalitov dvakrat kalciniranega in mletega
prahu KNN pri 625 °C z vklju~eno sliko uklonskega posnetka (SAED) iz izbranega obkro`enega podro~ja in b) zna~ilno kemijsko analizo z
rentgensko spektroskopijo (EDS) iz istega izbranega obkro`enega podro~ja na sliki a; c) posnetek TEM BF in d) visokolo~ljivostna slika
(HR-TEM) pri vi{ji pove~avi istega prahu. Iz ozna~enega ~rtkano kro`nega mesta na sliki TEM BF je bil narejen posnetek HR-TEM. Iz izbranega
~rtkano kvadratnega podro~ja na sliki HR-TEM sta bili izra~unani vstavljeni sliki 2D hitre Fourierjeve transformacije (FFT) in simulacija.
The elemental composition was determined with
EDS. A typical EDS spectrum of a selected analysis of
KNN powder from the dashed-line circled area in Figure
5a is shown in Figure 5b. The analyses confirmed the
presence of Nb, Na, K and O in ratios close to the KNN
stoichiometry. An observation at higher magnification
(Figure 5c) revealed particles with cuboidal shapes,
which is typical for KNN powders.
The HR-TEM image with a 2D Fast Fourier Trans-
form (FFT) inset (Figure 5d) shows a small KNN
particle at an atomic resolution in the [00-1] zone axis.
The simulated image in the inset of Figure 5d shows the
position of the atoms in the [00-1] zone axis and corres-
ponds to the FFT image calculated from the experimen-
tal HR-TEM image. The measured distance between the
atoms from the HR-TEM image in the (100) direction is
0.40 nm, which is in very good agreement with the cell
parameters of the KNN.22
STEM/EDS elemental mapping of the double-cal-
cined KNN powder at 625 °C and subsequently milled is
shown in Figure 6, with a distribution of the important
comparable elements of Nb, Na and K and their overlay.
The analysis nicely showed a homogeneous distribution
of all three elements in the range of the detection limit of
the EDS. As shown in the EDS spectrum, in some areas
a small amount of Si was also detected. However, this
M. FEIZPOUR et al.: SOLID-STATE SINTERING OF (K0.5Na0.5)NbO3 SYNTHESIZED FROM ...
Materiali in tehnologije / Materials and technology 49 (2015) 6, 975–982 979
Figure 6: STEM/EDS elemental mapping of the double-calcined
KNN powder at 625 °C and subsequently milled with EDS spectrum
of the selected dashed-line circled area shown in the overly STEM
image. Unmarked peaks at 8.04 keV and 8.90 keV belong to the
Cu-K1 and K1 characteristic X-rays, respectively, which came from
the Cu TEM grid that was used as a sample holder.
Slika 6: Povr{inska kemijska analiza z rentgensko spektroskopijo
(STEM/EDS) dvakrat kalciniranega in mletega prahu KNN pri 625 °C
s spektrom EDS iz izbranega obkro`enega podro~ja, prikazanega na
prekrivni sliki STEM. Neozna~eni vrhovi pri 8,04 keV in 8,90 keV
pripadajo Cu-K1 in K1 karakteristi~nim rentgenskim `arkom, ki
izvirajo iz Cu-mre`ice TEM, in ki je bila uporabljena kot nosilec
vzorca.
Figure 7: Orientation-contrast SEM images of a polished surface of
KNN ceramics sintered under two different configurations: a) closed
crucible with KNN packing powder – CC/wPP and b) open crucible
without KNN packing powder – OC/woPP. The  sign in Figure a
marks a Si-containing grain with a stoichiometry close to
K6Nb6Si4O26.
Slika 7: Posnetki SEM v na~inu orientacijskega kontrasta poliranih
povr{in keramike KNN, sintrane pri dveh razli~nih konfiguracijah: a)
zaprt lon~ek z zasipom KNN – CC/wPP in b) odprt lon~ek brez zasipa
KNN – OC/woPP. Znak  na sliki a ozna~uje zrno, ki vsebuje Si,
katerega sestava je blizu K6Nb6Si4O26.
amount is very low for the exact characterization and its
determination.
3.2 Sintering of the 625 °C-calcined KNN powder
The relative density of the ceramics sintered under
the CC/wPP and OC/woPP configurations were 91.5 %
and 93.4 %, respectively. Figure 7 shows the orienta-
tion-contrast SEM images of the polished surface of the
KNN ceramics sintered under the two different configu-
rations mentioned above. The percentages of porosity,
which are visible as black areas in the images, were
evaluated using ImageTool version 3.00 software and
measured approximately as 17.5 % and 6.5 % for the
samples sintered under CC/wPP and OC/woPP configu-
rations, respectively. Some grains may have been pulled
out, which typically happens during the grinding and po-
lishing of KNN-based ceramics. Therefore, these values
are not necessarily representative of the amount of poro-
sity in the bulk ceramics. Apart from the percentages of
the porosity, the shapes of the grains are also different in
these two microstructures. The CC/wPP configuration
mainly resulted in cuboidal grains with plane grain
boundaries, whereas the OC/woPP configuration led to
almost non-polyhedral grain shapes with intragranular
porosities. According to Acker et al.23 more faceted
grains can be achieved for the KNN with the amount
fraction 2 % excess on the A-site. Later on, they
explained that for stoichiometric KNN the excess alkali
elements from the packing powder atmosphere drives the
ceramic composition towards the A-site excess regime
and consequently away from the ideal stoichiometric
conditions for sintering dense materials.24 This may
explain why using packing powder for the suppression of
the evaporation of alkali elements during sintering may
cause the density of the ceramic to be even lower than
the density of the unprotected sample.25 This contradicts
the popular belief about the sintering of KNN-based
ceramics and raises an interesting question on the use of
KNN packing powder, to what extent the addition of an
alkali-rich atmosphere during sintering is beneficial for
the properties of KNN-based ceramics? In addition,
based on our observations, the weight losses of samples
during the sintering process were significantly similar
for both sintering configurations.
Figure 8 shows the XRD patterns of ceramics sin-
tered under different sintering configurations of CC/wPP
and OC/woPP. Though the KNN powder was calcined
only at 625 °C and, thus, has a lower crystallinity and
homogeneity compared to the usual higher-calcination-
temperature powders, both sintered ceramics have highly
crystalline and homogenous structures. All the main
peaks can be indexed as the KNN perovskite phase with
a monoclinic structure.22 However, in both patterns, a
trace of a secondary phase was found at around 2 
30 °. These two extra visible peaks, which are not related
to the KNN perovskite phase, are in a very good accord-
ance with the main peaks of K6Nb6Si4O26 with JCPDS
M. FEIZPOUR et al.: SOLID-STATE SINTERING OF (K0.5Na0.5)NbO3 SYNTHESIZED FROM ...
980 Materiali in tehnologije / Materials and technology 49 (2015) 6, 975–982
Figure 8: XRD patterns of ceramics sintered under two different configurations: a) closed crucible with KNN packing powder – CC/wPP and b)
open crucible without KNN packing powder – OC/woPP
Slika 8: Rentgenogram keramike KNN, sintrane pri dveh razli~nih konfiguracijah: a) zaprt lon~ek z zasipom KNN – CC/wPP in b) odprt lon~ek
brez zasipa KNN – OC/woPP
card number 01-072-0558. The existence of a Si-contain-
ing phase was also proved in the compositional analyses
of the sintered samples from both sintering configura-
tions. For example, the grain which is marked with  in
Figure 7a represents a phase with the following atomic
stoichiometry for Na : K : Nb : Si : O as (0.47 : 13.34 :
15.09 : 9.11 : 62.00) %, which matches very well with
the nominal atomic stoichiometry of K6Nb6Si4O26, i.e.,
(14.29 : 14.29 : 9.52 : 61.90) %.
4 CONCLUSIONS
For the first time, 625 °C is suggested as the possible
low-calcination temperature for the solid-state synthesis
of K0.5Na0.5NbO3 from an alkali-carbonates source. A
microstructural analysis of the KNN sample processed
from double-calcined KNN powder at 625 °C shows that
a typical microstructure of KNN, otherwise processed
with higher calcination temperatures ( 800 °C), can be
achieved. The sintering of KNN in an open-crucible
setup without using packing powder contributed to a
higher density. A trace of K6Nb6Si4O26 secondary phase
was found in the SEM and XRD of sintered samples, but
its origin requires further investigation.
Acknowledgement
The authors would like to dedicate this paper to the
memory of the late Professor Marija Kosec on the third
anniversary of her death, 23rd Dec.
This work was financially supported by Ministry of
Science, Research and Technology of Iran with the grant
No. 38-1392-063 from Materials and Energy Research
Center. In addition, the support of Jo`ef Stefan Institute
– Electronic Ceramics Department (K5) and Institute of
Metals and Technology (IMT), both from Ljubljana, Slo-
venia, is greatly appreciated. The contributions of Mo-
hammad Ali Bahrevar, Alireza Aghaei, Andreja Ben~an
Golob, Tadej Rojac, Goran Dra`i}, Jurij Koruza, Kata-
rina Vojisavljevi}, Julian Walker, Jernej Pavli~, Jitka
Hre{~ak, Tina Bakari~, Marko Vrabelj, Andra` Bra-
de{ko, Maja Makarovi~, Jerca Praprotnik, Tina Rucigaj,
Jena Cilen{ek, Silvo Drnov{ek, Edi Kranjc and Maja
Koblar are much appreciated. Special thanks from Mahdi
go to Prof. Barbara Mali~, the head of K5, for her kind
and generous hosting at K5 during the visiting research
period of 2014, and to Assist. Prof. Matja` Godec, the
director of IMT, for his special attention to the TEM
investigations and our mutual cooperation.
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M. FEIZPOUR et al.: SOLID-STATE SINTERING OF (K0.5Na0.5)NbO3 SYNTHESIZED FROM ...
I. PAULIN: STABILITY OF CLOSE-CELL Al FOAMS DEPENDING ON THE USAGE OF DIFFERENT FOAMING AGENTS
983–988
STABILITY OF CLOSE-CELL Al FOAMS DEPENDING ON THE
USAGE OF DIFFERENT FOAMING AGENTS
STABILNOST ALUMINIJEVIH PEN Z ZAPRTO POROZNOSTJO
GLEDE NA UPORABO RAZLI^NIH PENILNIH SREDSTEV
Irena Paulin
Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
irena.paulin@imt.si
Prejem rokopisa – received: 2015-10-21; sprejem za objavo – accepted for publication: 2015-10-28
doi:10.17222/mit.2015.322
Close-cell Al foams produced by the powder-metallurgy (PM) route can be made with different foaming agents, regarding the
type of liberated gas. Most commonly, H2 gas is used as the liberation agent, but there is a huge improvement made with CO2
gas liberating agents, such as calcite and dolomite. In order to determine the benefits and/or disadvantages of foaming agents,
studies of the foam pores’ stability, depending on the type of liberated gas, were performed.
The stability of Al foams was studied by different analytical techniques, i.e., AES, expandometer, heating microscopy and
SEM/EDS. AlSi12 aluminium powder as the matrix material and TiH2 and CaCO3 as the foaming agents liberating different
gases – one based on H2 and another on CO2 – were used. Based on the obtained results, the mechanism of foam stability was
studied and a comparison of the two foaming agents was made and evaluated.
Keywords: Al foams, CaCO3, TiH2, stability of pores, oxide layers
Aluminijeve pene z zaprto poroznostjo, narejene po postopku metalurgije prahov, se lahko pripravi z razli~nimi penilnimi
sredstvi glede na plin, ki se pri penjenju spro{~a. Najbolj pogosto je uporabljeno penilno sredstvo na osnovi H2 plina, vendar se
v zadnjem ~asu veliko uporabljajo tudi penila na osnovi CO2, kot sta kalcit in dolomit. Da bi ugotovili prednosti in slabosti
razli~nih penilnih sredstev, so bile narejene {tudije stabilnosti por glede na tip izhajajo~ega plina.
Stabilnost aluminijevih pen je bila raziskovana z razli~nimi analitskimi tehnikami – z AES, ekspandometrom, segrevalnim
mikroskopom in SEM/EDS analizo. Za raziskavo smo uporabili aluminijev prah AlSi12 in TiH2 ter CaCO3 kot penilni sredstvi z
razli~nima tipoma spro{~enega plina – en na osnovi H2 in drugi na osnovi CO2. Glede na rezultate raziskave smo opisali
mehanizem za stabilizacijo aluminijevih pen in primerjali ter ocenili obe penilni sredstvi.
Klju~ne besede: Al pene, CaCO3, TiH2, stabilnost por, oksidne plasti
1 INTRODUCTION
Al foams as a promising class of materials with great
chances for applicative use1–3 due to their mechanical,
chemical and physical properties were analyzed from
different viewpoints. Most studies were performed on the
synthesis, characterization and mechanical testing of
foams, but few investigated the stability of foams by
adding ceramic particles into the matrix and no analysis
was made to explain the stability of the interior of the
pores and the properties of the material depending on
this stability.
The PM production process starts with mixing metal
powders – elementary-metal powders, or alloyed pow-
ders or metal powder blends – and a foaming agent, after
which the mixture is compacted to yield a dense,
semi-finished product.4,5 The method depends on the
preparation of the precursors6, which in principle can be
made by any technique that ensures the embedding of the
foaming agent into the metal matrix without any ob-
servable residual open porosity. In general, precursors
consist of a compacted metallic powder and a foaming
agent that are sintered at a pre-determined temperature.
Due to the high temperature of the thermal treatment, the
foaming agent decomposes into a solid component that is
incorporated into the matrix material, and a gas compo-
nent that causes foaming of the matrix material.7
The metal matrix is a semi-solid, thus the gas libe-
rated from the agent is forming pores that give a specific
shape and mechanical properties to the foams. The
obtained foam must be rapidly and properly cooled down
to retain the bloated structure. Pore formation depends
on the amount and the type of the liberated gas, the time
and temperature of the thermal treatment and the cooling
process. All those parameters influence the stability of
the pores and thus the stability of the material properties.
The quantitative expansion and collapsing behaviors of
the samples were characterized with a mechanical ex-
pandometer and heating microscopy.
This investigation is focused on the closed-cell Al
foam produced by the powder metallurgy (PM) process
using different gas-liberating foaming agents, such as
TiH2 and CaCO3. The aim of the study was to investigate
the mechanism of the stability of the Al foams to
understand the process of pore formation, the possibility
of predicting the foam properties and for the optimal
selection of the foaming agent, depending on the ex-
pected properties.
Materiali in tehnologije / Materials and technology 49 (2015) 6, 983–988 983
UDK 669.71:66.069.852:621.762 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 49(6)983(2015)
2 EXPERIMENTAL PROCEDURES
In the present study of Al-foam stability, the interior
surfaces of the pores were carefully studied. AlSi12 alu-
minium powder, TiH2 and CaCO3 as the foaming agents
liberating different gases – one based on H2 and the other
on CO2 – were used. The decomposition of both foaming
agents could be described by the two chemical reactions
given below, and the volume of liberated gas was
calculated from the ideal gas law:
pV = nRT (1)
where p is the pressure, V is the volume of gas, n is the
amount of substance (n = m/M, where m is the mass and
M is the molar mass), R is the gas constant and T is the
temperature.
The decomposition of:
TiH2: TiH2(s)  Ti(s) + H2(g) (2)
1 g of TiH2 liberates at T = 923 K (650 °C) a gas volume
of V = 1.51 L
CaCO3: CaCO3(s)  CaO(s) + CO2(g) (3)
1 g of CaCO3 liberates at T = 923 K (650 °C) a gas vo-
lume of V = 0.76 L
The calculation of the gas liberation at standard tem-
perature and pressure shows that the volume of liberated
gas is relatively high, but in the real foaming process,
where the pressure of gas in the material is higher, the
volume of gas is smaller. On the other hand, some of the
liberated gas escapes out of the material when the
pressure in the formed bubbles is too high.8,9 The process
of pore formation can be explained by the theory of soap
bubbles.
A calculation of the amount of liberated gas shows
that the largest amount of gas liberated under the same
conditions is obtained with the TiH2 decomposition,
while the amount of liberated gas with the decomposi-
tion of CaCO3 is just a half of that. A calculation of the
volumes of liberated gas gives how much foaming agent
is needed for similar results of foaming.
All the samples were prepared by the same pro-
cedure, i.e., mixing the AlSi12 powder with mass
fractions 1 % TiH2 and 3 % CaCO3, respectively, cold
compaction at 1200 MPa 6 into semi-products, called
precursors, and foaming at a temperature that is specific
for each foaming agent.10 After preparing the foams,
some in a lab furnace and some in the expandometer,
specimens were analyzed with different techniques:
SEM/EDS (SEM JEOL 6500F with Oxford INCA EDX
analyzer) was used for micro-chemical analyses of the
inclusions in the foam thin walls, while the surface of the
interior of the pores and the exterior of the foams were
investigated by AES (Microlab 310 F VG–Scientific).
AES depth profiles were performed by ion etching with a
velocity of 0.125 nm/min. The AES surface analyses
were also performed for the initial powder particles to
confirm the presence of an oxide layer on the surface.
The mechanism of foam stability was explained by
the investigation with heating microscopy and the
expandometer. Both methods were used to determine the
time and the temperature of foaming the material. In
heating microscopy, small samples of precursor material
with different foaming agents were put into the furnace
within the microscope and heated up to 800 °C with a
heating rate of 7 °C/min. Precursors were inserted at
room temperature into the microscope and heated up
until the foaming process did not stop and foams started
to collapse. During the heating, changes to the shape and
the size of the investigated material were observed. This
method enabled us to determine the temperature of the
foaming process start and the temperature of the foam
collapse.
The mechanical expandometer was used to measure
the expansion of the foamable precursors inside a cylin-
drical mould as a function of time and temperature.11 The
temperature of the expandometer furnace was held
constant at 750 °C for the TiH2 foaming agent and at 800
°C for the CaCO3 foaming agent.
3 RESULTS AND DISCUSSION
The foaming process was studied by observing pre-
cursors in the heating microscope during heating. Sam-
ples began to change their shape and volume at a certain
temperature, depending on the type of foaming agent.
The results obtained with the heating microscope
enabled us to determine the approximate time and tem-
perature of the beginning of the process. Changes of the
foam expansion (precursor was made with AlSi12 +
TiH2) are represented in Figure 1.
The heating microscope also enabled us to observe
the collapsing of the foams after extended foaming times
at individual temperatures. This resulted in gas escaping
from the bubbles that were formed inside the matrix
material. At a certain temperature, the matrix material
became viscous, which prevented the escape of gas from
the material. Higher temperatures caused a drop in the
material viscosity, and due to pressure in the gas bubbles
the gas escaped and the foam collapsed. That was the
reason that the foams had to be cooled down fast after
the desired expansion of material was achieved. Cooling
could be achieved either by immersing the foamed
product into water or with cold compressed air. In our
case cooling in water was applied.
Foams prepared with different foaming agents were
cut into smaller pieces and samples for the investigation
with SEM/EDS were prepared, using a standard metallo-
graphic procedure with grinding and polishing. Figures
2 and 3 show SE (secondary electron) and BE (back-
scattered) images of the polished foamed material, where
residuals of solid particles of foaming agents captured in
the walls of the foams after the gas liberation are visible.
In12,13 the authors confirmed that those ceramic par-
ticles helped to stabilize the Al foams. In analyses of
I. PAULIN: STABILITY OF CLOSE-CELL Al FOAMS DEPENDING ON THE USAGE OF DIFFERENT FOAMING AGENTS
984 Materiali in tehnologije / Materials and technology 49 (2015) 6, 983–988
I. PAULIN: STABILITY OF CLOSE-CELL Al FOAMS DEPENDING ON THE USAGE OF DIFFERENT FOAMING AGENTS
Materiali in tehnologije / Materials and technology 49 (2015) 6, 983–988 985
Figure 3: SEM images of thin foam-cell wall with residuals of CaCO3 foaming agent: a) SE image and b) BE image; the same spot
Slika 3: SEM posnetka tanke celi~ne stene z ostanki penila CaCO3: a) SE posnetek in b) BE posnetek; isto mesto
Figure 2: SEM images of thin foam-cell wall with Ti residuals of TiH2 foaming agent: a) SE image and b) BE image
Slika 2: SEM posnetka tanke celi~ne stene aluminijeve pene z ostanki Ti iz penilnega sredstva TiH2: a) SE posnetek in b) BE posnetek
Figure 1: Heating microscope – change of precursor AlSi12 + TiH2 size: a) 420 °C, no changes, b) 500 °C, first noticed change, c) 520 °C,
precursor starts rapidly to bloat, d) 525 °C, still rapidly bloating, e) 580 °C, maximum expansion, f) 600 °C, collapse of sample, keeping the
shape of a loaf
Slika 1: Segrevalna mikroskopija – sprememba velikosti prekurzorja iz AlSi12 + TiH2: a) 420 °C, {e nobenih sprememb velikosti, b) 500 °C,
prve spremembe, c) 520 °C, vzorec za~ne hitro nara{~ati, d) 525 °C, vzorec zelo hitro nara{~a, e) 580 °C, vzorec dose`e najve~jo velikost, f) 600
°C, vzorec se sesede, vendar obdr`i obliko hleb~ka
ceramic additions to Al foams during preparation, a spe-
cial emphasis was given to the positive effects of CaO
particles.
One of the important properties of aluminium foams
is also their stability during the foaming process. The
quantitative expansion and collapse behavior of the
samples were characterized with the mechanical expan-
dometer, in which the precursor was exposed to a con-
stant temperature of 750 °C for AlSi12 + TiH2 and 800
°C for AlSi12 + CaCO3. The expansion was measured
with seven samples prepared using TiH2 and CaCO3
agents, respectively, and it was approximately 245 %
with the CaCO3 and up to 285 % with the TiH2 precur-
sors. Nevertheless, the TiH2 precursors appeared to
collapse faster than the CaCO3 precursors.
Observations in the mechanical expandometer and in
the heating microscope revealed that the foams prepared
with a foaming agent based on the liberation of CO2 gas
were more stable than those that were prepared with the
liberation of H2. The obtained results were confirmed by
the AES analysis of the interior surfaces of the pores
where the oxide layers were analyzed and thus the sta-
bility can be explained.6,14,15 Two interior pore surfaces,
obtained by foaming AlSi12 + TiH2 and AlSi12 + CaCO3
mixtures, were analyzed, and the results were compared
(Figures 4 and 5). Also, the external surface of the foam
was analyzed to see the difference in the interior and
exterior surfaces of the foam (Figure 6).
The results showed differences in the thicknesses of
the oxide layers on the surface of the pore interior de-
pending on the applied foaming agent and on the surface
of the foam exterior. The thickness of the oxide layer on
the interior pore surface obtained by foaming with TiH2
was 10–15 nm, while the thickness of the oxide layer on
the external surface of the foam was 35–40 nm. On the
other hand, the thickness of the oxide layer on the inte-
rior pore surface obtained by foaming with CaCO3 was
90–125 nm, while the oxide layer on the external surface
of the foam was about 45–60 nm thick. The difference in
the oxide layer thicknesses was explained by the fact that
H2 represented a reduction atmosphere in the material,
and oxygen formed on the surface of the matrix material
only a thin oxide layer. Oxide was present in the material
as a thin oxide layer on the surface of the powder parti-
cles. On the other hand, the CaCO3 powder created a
CO2 atmosphere during the foaming process, so that the
oxide layer on the pore surface was easily formed. Thick
oxide layers enabled better stability of the foams pro-
duced by foaming agents based on CO2. However, in
both cases there was no great difference in the thick-
nesses of the oxide layers on the external surface.
The question appears, why CO2-based foaming
agents are not yet more widely used? And the answer is
that agents releasing CO2 need higher temperatures for
the liberation of the gas and the needed temperature is
approximately 100 °C above the melting point of pure
aluminium10. In such a case, the matrix material is
already molten and the viscosity of such material is not
high enough to prevent gas escaping out of the material.
I. PAULIN: STABILITY OF CLOSE-CELL Al FOAMS DEPENDING ON THE USAGE OF DIFFERENT FOAMING AGENTS
986 Materiali in tehnologije / Materials and technology 49 (2015) 6, 983–988
Figure 6: AES depth profiles of the exterior AlSi12 + TiH2 foam
surface
Slika 6: Globinski profili AES zunanje povr{ine pene AlSi12 + TiH2
Figure 4: AES depth profiles of the interior surface of the AlSi12 +
TiH2 foam pores
Slika 4: Globinski profili AES notranje povr{ine v pori pene AlSi12 +
TiH2
Figure 5: AES depth profiles of the interior surface of the AlSi12 +
CaCO3 foam pores
Slika 5: Globinski profili AES notranje povr{ine v pori pene AlSi12 +
CaCO3
The solution is to pre-heat the precursors for some time
before the foaming process begins. When foaming
agents in those precursors are then exposed to an ele-
vated temperature, decomposition has already started and
the agent did not need such a long time to decompose
completely, while the matrix aluminium has insufficient
time to melt completely.
The sizes and shapes of the pores in the foams pre-
pared by exposure to H2 or CO2 gas are different. H2 gas
forms spherical pores, while CO2 forms smaller pores of
not completely spherical shape; they are slightly elon-
gated in the horizontal direction (Figure 7). However,
the stability of foams is better when a CO2-based foam-
ing agent is used, but the pore shapes and consequently
the mechanical properties in the vertical direction are
better when using H2-based foaming agents. The final
application of the foamed material determines the selec-
tion of the foaming agent as well as the selection of the
matrix material. On the other hand, the use of a combina-
tion of both foaming agents in a well-defined proportion
can be an alternative.
4 CONCLUSIONS
CaCO3 is a good substitute for the expensive and
more widely used TiH2 foaming agent, though some
properties of foams, like the stabilization of pores, are in
this case even better. Natural oxidation of the pore
surfaces inside the material, as well as the oxidation of
external surfaces, helps to stabilize foams so that they do
not collapse after the completed foaming.
The oxide layers detected by the AES analysis on the
surface of aluminium powders helped to prevent prema-
ture melting of the matrix material and enabled foaming
agents to decompose before the matrix material melted
and lost sufficient viscosity. Similar case is observed
when using TiH2 foaming agent that decomposes at a
relatively low temperature, but the additional oxide layer
on the surface of the powder delays the beginning of the
decomposition as it was pre-studied and confirmed in the
literature.
According to12,13 and from our research the conclu-
sion can be made that a thicker oxide layer and the
residuals of the ceramic solid particles from the CaCO3
foaming agent give an opportunity to foaming agents
based on CO2 to be used more often in the production of
aluminium foams.
Nevertheless, the stability of the foams is better when
a CO2-based foaming agent is used, but the pore shapes
are more spherical and consequently the mechanical
properties, as known from the literature, are higher when
H2-based foaming agents are used, so the use of a com-
bination of both foaming agents in a well-defined
proportion can be an alternative.
Acknowledgements
This work was supported by the “Physics and
chemistry of porous aluminum for Al panels capable of
highly efficiently energy absorption” project L2-2410
(D), funded by the Slovenian Research Agency.
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I. PAULIN: STABILITY OF CLOSE-CELL Al FOAMS DEPENDING ON THE USAGE OF DIFFERENT FOAMING AGENTS
Materiali in tehnologije / Materials and technology 49 (2015) 6, 983–988 987
Figure 7: Images of Al foam, cut from loaf: a) AlSi12 + TiH2 and b)
AlSi12 + CaCO3
Slika 7: Posnetka Al pene, odrezan Al hleb~ek: a) AlSi12 + TiH2 in b)
AlSi12 + CaCO3
5 I. Duarte, J. Banhart, A study of aluminium foam formation – Kine-
tics and microstructure, Acta Materialia, 48 (2000) 9, 2349–2362,
doi:10.1016/S1359-6454(00)00020-3
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of aluminium foams by powder-metallurgy process: compacting of
precursors, Mater. Tehnol., 45 (2011) 1, 13–19
7 I. Duarte, P. Weigand, J. Banhart, Foaming kinetics of aluminium
alloys, In: J. Banhart, N. A. Ashby, N. A. Fleck (Eds.), Metal Foams
and Porous Metal Structures, MIT Verlag, Bremen, Germany 1999,
97-104
8 I. Jeon, T. Asahina, K. J. Kang, S. Im, T. J. Lu, Finite element simu-
lation of the plastic collapse of closed-cell aluminum foams with
X-ray computed tomography, Mechanics of Materials, 42 (2010) 3,
227–236, doi:10.1016/j.mechmat.2010.01.003
9 A. E. Markaki, T. W. Clyne, The effect of cell wall microstructure on
the deformation and fracture of aluminium-based foams, Acta Mate-
rialia, 49 (2001) 9, 1677–1686
10 I. Paulin, Synthesis and characterization of Al foams produced by
powder metallurgy route using dolomite and titanium hydride as a
foaming agents, Mater. Tehnol., 48 (2014) 6, 943–947
11 P. Weigand, Untersuchung der Einflussfaktoren auf die pulver-
metallurgische Herstellung von Aluminiumschäumen, Dissertation,
RWTH Aachen University, Aachen 1999
12 J. Baumeister, J. Weise, A. Jeswein, M. Busse, M. Haesche, Metalic
foam production from AlMg4.5Mn recycling machining chips by
means of thixocasting and the effects of diferent additions for sta-
bilization, MetFoam 2007, Montreal, Canada 2007
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tions on foam expansion and stability in compacted Al-TiH2 powder
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doi:10.1002/adem.200405145
14 J. Banhart, Metal foams: Production and stability, Adv. Eng. Mater.,
9 (2006) 8, 781–794, doi:10.1002/adem.200600071
15 A. Haibel, A. Rack, J. Banhart, Why are metal foams stable?,
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I. PAULIN: STABILITY OF CLOSE-CELL Al FOAMS DEPENDING ON THE USAGE OF DIFFERENT FOAMING AGENTS
988 Materiali in tehnologije / Materials and technology 49 (2015) 6, 983–988
W. RATTANASAKULTHONG et al.: EVOLUTION OF THE MICROSTRUCTURE AND MAGNETIC PROPERTIES ...
989–992
EVOLUTION OF THE MICROSTRUCTURE AND MAGNETIC
PROPERTIES OF A COBALT-SILICON-BASED ALLOY IN THE
EARLY STAGES OF MECHANICAL MILLING
RAZVOJ MIKROSTRUKTURE IN MAGNETNIH LASTNOSTI
ZLITINE Co-Si V ZA^ETNEM STADIJU MEHANSKEGA
LEGIRANJA
Watcharee Rattanasakulthong1, Chitnarong Sirisathitkul2, Peter Franz Rogl3
1Department of Physics, Faculty of Science, Kasetsart University, Bangkok, Thailand
2Magnet Laboratory, Molecular Technology Research Unit, School of Science, Walailak University, Nakhon Si Thammarat, Thailand
3Institute of Materials Chemistry and Research, University of Vienna, Vienna, Austria
chitnarong.siri@gmail.com
Prejem rokopisa – received: 2014-12-11; sprejem za objavo – accepted for publication: 2014-12-22
doi:10.17222/mit.2014.300
The early stages in the mechanical alloying of amount fractions x = 40 % cobalt (Co) and x = 60 % silicon (Si) powders were
investigated using X-ray diffractometry (XRD), scanning electron microscopy (SEM), differential thermal analysis (DTA) and
vibrating-sample magnetometry (VSM). After 2–8 h of ball-milling, the characteristic XRD peaks of the face-centered-cubic
(fcc) Si and hexagonal close-packed (hcp) Co phases remained sharp without a cobalt-silicide phase. As the milling progressed,
the particle size observed by SEM tended to reduce, being accompanied by smoother edges and a narrow size distribution. On
the DTA curves, between 200 °C and 1200 °C, exothermic peaks indicated a ferromagnetic-to-paramagnetic transition, whereas
endothermic peaks corresponded to the lattice recovery, the transition from a hcp to a fcc Co and melting. The longest milling of
up to 8 h significantly increased the magnetic squareness and the coercive field.
Keywords: Co-Si alloys, ball milling, XRD, VSM, DTA
Preiskovan je bil za~etni stadij mehanskega legiranja zlitine, sestavljene iz prahov z mno`inskim delom x = 40 % kobalta (Co) in
x = 60 % silicija (Si). Uporabljena je bila rentgenska difraktometrija (XRD), vrsti~na elektronska mikroskopija (SEM), diferen-
cialno termi~na analiza (DTA) in vibracijska magnetometrija vzorcev (VSM). Po mletju od 2 h do 8 h v krogli~nem mlinu so {e
ostali ostri zna~ilni XRD-vrhovi faz fcc Si in hcp Co brez Co-silicidne faze. S podalj{evanjem ~asa mletja se s SEM opazi
zmanj{anje velikosti delcev prahov, zo`anje njihove velikostne porazdelitve in pove~anje zaobljenja robov delcev. Na
DTA-krivuljah se je med 200 °C in 1200 °C pojavil eksotermni vrh, ki je posledica prehoda iz feromagnetne v paramagnetno
fazo, endotermni vrh pa je posledica poprave re{etke zaradi prehoda iz hcp v fcc Co in nato taljenja. Najdalj{i ~asi mletja (do 8
h) pomembno vplivajo na magnetno koercitivnost in pove~anje kvadrati~nosti histerezne zanke.
Klju~ne besede: zlitina na osnovi Co-Si, krogli~no mletje, XRD, VSM, DTA
1 INTRODUCTION
Mechanical milling has been successfully used in
producing a variety of magnetic amorphous alloys,
intermetallic compounds, nanocomposites and nano-
crystalline powders.1,2 Milling ferromagnetic powders
(i.e., Co, Fe, Ni) with non-magnetic metals (e.g., Au, Cu,
Ag) gives rise to mechanical alloys with a giant magne-
toresistance (GMR) effect.3 Since an addition of Si
reduces the magnetic anisotropy and the eddy-current
loss in commercial steels, Fe–Si mechanical alloys also
received much interest.4–7 By contrast, Co–Si alloys only
gained attention in the last decade after being recognized
as hydrogen-storage materials for nickel-metal hydride
batteries. The discharge capacity and cycling ability of
negative electrodes are reportedly improved when using 1
: 1 or 2 : 1 Co–Si milled for 10–80 h.8–10 In addition to
homogenizing their sizes and shapes, the ball milling
leads to several compounds as shown by the Co–Si phase
diagram.11 The Co2Si, CoSi and CoSi2 phases affect the
magnetic and hydrogen-storage properties of the
alloys.10–14
In this work, the evolution of the phases during the
initial period of the milling of Co with x = 60 % Si is
examined. The thermal and magnetic properties of these
Co–Si powders from the early stage of the ball-milling
for up to 8 h are also reported.
2 EXPERMENTAL WORK
Elemental crystalline cobalt powder (a 99.8 % purity
with the average particle size of less than 2 μm) and
silicon powder (a 99 % purity with the average particle
size of less than 44 μm) were mixed in an atomic ratio of
40 : 60 in a steel vial loaded with steel balls of 3 mm in
diameter. The ball-to-powder mass ratio was around
15 : 1. The vial was spun at 595 r/min on a vario-planetary
mill (Fritsch) for (2, 4, 6 and 8) h in a dry condition
under air. The structural and magnetic properties of the
milled powders were characterized with X-ray diffrac-
Materiali in tehnologije / Materials and technology 49 (2015) 6, 989–992 989
UDK 621.318.1:621.762:543.422.3 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 49(6)989(2015)
tion (XRD) using Cu-K radiation and vibrating sample
magnetometry (VSM), respectively. The coercive field
was determined from the x-intercept of the hysteresis
loop obtained with VSM, whereas the y-intercept corres-
ponded to the remanent magnetization. A ratio of the
remanent magnetization to the saturation magnetization
is referred to as the magnetic squareness in Table 1. The
thermal properties were studied using a differential
thermal analysis (DTA) with a heating rate of 5 °C min–1.
Table 1: Magnetic squareness and coercive field of Co – x(Si) = 60 %
powders
Tabela 1: Magnetna kvadratnost in koercitivno polje prahov iz Co –
x(Si) = 60 %
Milling time
(h)
Magnetic
squareness
Coercive field
(kA m–1)
2 0.28 15.15
4 0.29 15.32
6 0.28 15.48
8 0.37 19.41
3 RESULTS AND DISCUSSION
The XRD patterns of the Co–Si powders after the
milling shown in Figure 1 have characteristic peaks of
the fcc Si phase (2
 = 28.440 °, 47.300 ° and 56.120 °)
and the hcp Co phase (2
 = 47.220 °, 44.080 ° and
41.440 °). According to the literature15, allotropic Co
undergoes a transition from the hcp to the fcc structure at
a temperature around 450 °C. Previous works suggested
that the ball milling can also induce such an allotropic
transformation by virtue of a defect accumulation.16–19
The mixed phase may be converted into a highly
distorted hcp structure in the early stage of milling but
further milling leads to a fcc structure as a result of the
stacking faults from the plastic deformation.17 The CoSi
and CoSi2 phases are not clearly detected. The formation
of silicide compounds generally requires prolonged
milling and, as demonstrated with Pd–Si1, it is dependent
on the volume fractions of the two elemental powders. In
some cases, the milling is carried out at a high
temperature to stimulate the formation of compounds.1
In our case, all Co and Si peaks are rather sharp
indicating a high degree of crystallinity. However, the
peaks clearly broaden and their intensities are decreased
after the milling for 8 h. The small peaks of Si slightly
below 70 ° in the case of milling for 2 h and 4 h dis-
appear after the longer milling.
The morphologies of the Co–Si powders at various
milling times obtained with SEM are shown in Figure 2.
After the milling for 2 h, several particles are still bulky,
with straight and sharp edges. The edges are gradually
broken or rubbed off and the particles become smoother
as the milling progresses. Furthermore, the particle size
is clearly reduced after the longest milling time of 8 h. It
was also reported that a decrease in the particle size of
Fe–Si powders began to be noticeable only after 5 h of
ball milling.7 It is known that milling modifies the size of
ball-milled powders by virtue of fracturing and cold
welding. Each process dominates in a different stage of
milling and the welding of particles is dominant in the
initial milling time of up to 3 h.5,6 The narrower size
distribution seen in the case of 8 h milling is a result of
simultaneous fracturing of larger particles and cold
welding of smaller particles. The powder from the early
stage of milling can be modeled as brittle Si embedded
in more ductile Co particles.
It is seen from the DTA curves in Figure 3 that after
2 h to 8 h of milling, the samples exhibit a broad exo-
thermic peak centered around 600 °C. The area is in-
creased with the milling time because more energy is
released in the case of prolonged milling. Another exo-
thermic peak is observed that was shifted from above to
W. RATTANASAKULTHONG et al.: EVOLUTION OF THE MICROSTRUCTURE AND MAGNETIC PROPERTIES ...
990 Materiali in tehnologije / Materials and technology 49 (2015) 6, 989–992
Figure 2: SEM micrographs of Co – x(Si) = 60 % powders after
milling for: a) 2h, b) 4h, c) 6h and d) 8 h
Slika 2: SEM-posnetki prahov Co – x(Si) = 60 % po mletju: a) 2 h, b)
4 h, c) 6 h in d) 8 h
Figure 1: XRD patterns of Co – x(Si) = 60 % powders after milling
for 2–8 h
Slika 1: XRD-posnetki prahov Co – x(Si) = 60 % po mletju od 2 h do
8 h
below 900 °C due to the increase in the milling time.
This small exothermic peak may correspond to the ferro-
magnetic-to-paramagnetic transition. In addition to both
exothermic peaks, there are endothermic peaks at around
1100 °C corresponding to the melt appearing in the
system and at around 450 °C resembling the transition
from hcp to fcc in bulk Co.15 The latter is less notable in
the case of a longer milling time because the defect
accumulation due to milling already facilitates the
transformation to the fcc phase. The heat absorption also
reduces the defect and dislocation density in the
lattice-recovery process at around 200–300 °C which is
clearly detected only in the case of 2 h milling.
The hysteresis loops of the milled Co–Si powders are
shown in Figure 4 and their magnetic parameters are
summarized in Table 1. Like the other granular Co
alloys3, the magnetization is not saturated under the
magnetic field of about 200 kA m–1. Both the coercive
field and squareness (approximated from the ratio of the
remanence to the maximum magnetization in 200 kA m–1)
remain rather constant in the case of 2–6 h milling but
increase significantly after the milling for 8 h. Interest-
ingly, the size of Co particles is slightly modified during
the 2–6 h milling. This can then be related to the depen-
dence of the magnetic properties on the particle size of
magnetic mechanical alloys. The coercive field is also
related to the lattice imperfections as the milling is pro-
gressed because they impede the domain-wall move-
ment.4,5 Although Co has a strong crystalline aniso-
tropy20, it is noted that these coercive-field values are
comparable to those of the Fe alloys with a high fraction
of Si, which could be further reduced with heat treat-
ments.5,7
4 CONCLUSION
The ball-milling of Co – 60 % Si for up to 8 h does
not significantly induce silicide and amorphous phases.
However, there are some modifications in the morpho-
logy and particle size during this early stage of milling.
As a result, the thermal and magnetic properties of
Co–Si powders are considerably influenced by the
milling time. In addition to the endothermic DTA peaks
corresponding to the lattice recovery, the allotropic Co
transition, the melting and the ferromagnetic-to-para-
magnetic transition give rise to an exothermic peak,
which shifts to a lower temperature as the milling pro-
gresses. The considerable reduction in the particle size
after the milling for 8 h results in an enhanced coercive
field and magnetic squareness.
Acknowledgements
This work was funded by Faculty of Science, Kaset-
sart University (ScRF-E5-2553). The authors would like
to thank Scientific Equipments Center, Prince of Songkla
University for the characterization facilities and Juree-
porn Noodam for her assistance in the characterizations.
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992 Materiali in tehnologije / Materials and technology 49 (2015) 6, 989–992
IN MEMORIAM
prof. dr. Milanu Trbi`anu, zaslu`nemu profesorju Univerze v Ljubljani
Poslovil se je dr. Milan Trbi`an, zaslu`ni profesor
Univerze v Ljubljani. Bil je eden tistih, ki so v zadnjih
desetletjih bistveno prispevali k uspe{nemu razvoju
livarstva v Sloveniji na ve~ podro~jih.
S prvo generacijo kamni{kih maturantov je leta 1953
kon~al gimnazijo in se vpisal na {tudij metalurgije na
takratni Fakulteti za rudarstvo, metalurgijo in kemijsko
tehnologijo, kjer je diplomiral leta 1960. Prakti~ne
izku{nje si je kot mlad in`enir pridobival v livarni Titan,
leta 1964 pa je bil izvoljen za asistenta na Katedri za
`elezarstvo.
Njegovo raziskovalno delo so bile preiskave litih
kovinskih materialov, predvsem vpliv termo{oka na
toplotno utrujenost litega `eleza. Iz te tematike je leta
1973 tudi doktoriral na Montanisti~ni univerzi v Leobnu.
Razvil je lastno preiskovalno napravo za ponavljajo~e se
obremenitve s termo{okom. Leta 1976 je bil izvoljen za
docenta, leta 1986 za izrednega ter leta 1992 za rednega
profesorja in predstojnika Katedre za livarstvo; naziv
zaslu`nega profesorja mu je bil podeljen leta 2011.
Njegova raziskovalna dejavnost je bila usmerjena v
probleme v livarstvu. @e pri diplomi je preiskoval
materiale za izdelavo livarskih form in jeder.
Takoj po zaposlitvi na Univerzi je za~el sodelovati s
podjetjem Termit Morav~e pri razvoju opla{~enih peskov
za postopek izdelave ulitkov Croning.
Poleg rednega dela je pomembno oblikoval Dru{tvo
livarjev Slovenije. Leta 1965 je bil izvoljen za tajnika
Dru{tva livarjev in opravljal to delo vse do leta 1992, ko
je postal predsednik dru{tva. V tej vlogi je ostal do leta
2005.
S predavanji je sodeloval na mnogih nacionalnih,
mednarodnih in svetovnih kongresih livarjev. Bil je tudi
pobudnik letnega posvetovanja livarjev v Portoro`u, ki
vsako leto privabi ve~ kot 250 udele`encev iz {tevilnih
dr`av.
Prof. Trbi`an je s svojim pedago{ko-vzgojnim, razis-
kovalnim, inovativnim in dru{tvenim delovanjem prispe-
val tudi k uspehu na{ih livarn, ki z visokotehnolo{kimi
ulitki uspe{no konkurirajo na mednarodnem trgu in
zagotavljajo lepo {tevilo delovnih mest.
S svojim zna~ilnim temperamentom je odkrival nova
obzorja in nas razveseljeval v ~love{kem in strokovnem
smislu. Profesor Trbi`an je bil {e poln na~rtov, orga-
niziral je `e sre~anje ob svoji 80-letnici, a nepri~akovana
smrt mu je prepre~ila vse nadaljnje na~rte.
Zaslu`nega profesorja dr. Milana Trbi`ana bomo
ohranili v najbolj{i lu~i v svojih spominih.
Prof. dr. Primo` Mrvar
Predstojnik Oddelka za materiale
in metalurgijo NTF UL
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LETNO KAZALO – INDEX
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ISSN 1580-2949
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MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY
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LETNIK / VOLUME 49, 2015/1, 2, 3, 4, 5, 6
2015/1
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Z borom dopirani hidrogenirani amorfni polprevodnik MEMS
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Wear behaviour of B4C reinforced hybrid aluminum-matrix composites
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Corrosion of CrN-coated stainless steel in a NaCl solution (w = 3 %)
Korozija nerjavnega jekla s CrN-prevleko v raztopini NaCl (w = 3 %)
I. Kucuk, C. Sarioglu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19
Preparation and properties of master alloys Nb-Al and Ta-Al for melting and casting of -TiAl intermetallics
Priprava in lastnosti predzlitin Nb-Al in Ta-Al za taljenje in ulivanje intermetalnih zlitin -TiAl
J. Juøica, T. ^egan, K. Skotnicová, D. Petlák, B. Smetana, V. Matìjka . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27
Decreasing the carbonitride size and amount in austenitic steel with heat treatment and thermomechanical processing
Zmanj{anje velikosti karbonitridov v avstenitnem jeklu s toplotno obdelavo in termomehansko predelavo
P. Martínek, P. Podaný, J. Nacházel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 31
Deep cryogenic treatment of H11 hot-working tool steel
Globoka kriogenska obdelava orodnega jekla H11 za delo v vro~em
P. Suchmann, D. Jandova, J. Niznanska . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37
Numerical prediction of the compound layer growth during the gas nitriding of Fe-M binary alloys
Numeri~no napovedovanje rasti spojinske plasti med plinskim nitriranjem binarnih zlitin Fe-M
R. Kouba, M. Keddam. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43
Antibacterial composite based on nanostructured ZnO mesoscale particles and a poly(vinyl chloride) matrix
Protibakterijski kompozit na osnovi nanostrukturnih delcev ZnO in osnove iz polivinil klorida
J. Sedlák, P. Ba`ant, J. Klofá~, M. Pastorek, I. Kuøitka. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55
Water-soluble cores – verifying development trends
Jedra, topna v vodi – preverjanje smeri razvoja
E. Adámková, P. Jelínek, J. Beòo, F. Mik{ovský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 61
Mathematical modelling and physical simulation of the hot plastic deformation and recrystallization of steel with
micro-additives
Matemati~no modeliranje in fizikalna simulacija vro~e plasti~ne predelave in rekristalizacije jekla z mikrododatki
E. Kalinowska-Ozgowicz, W. Wajda, W. Ozgowicz . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 69
Synthesis of NiTi/Ni-TiO2 composite nanoparticles via ultrasonic spray pyrolysis
Sinteza kompozitnih nanodelcev NiTi/Ni-TiO2 z ultrazvo~no razpr{ilno pirolizo
P. Majeri~, R. Rudolf, I. An`el, J. Bogovi}, S. Stopi}, B. Friedrich . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75
Hydroxyapatite coatings on Cp-Titanium Grade-2 surfaces prepared with plasma spraying
Nanos hidroksiapatita na povr{ino Cp-Titana Grade-2 z nabrizgavanjem s plazmo
R. Rudolf, D. Stamenkovi}, Z. Aleksi}, M. Jenko, I. \ordevi}, A. Todorovi}, V. Jokanovi}, K. T. Rai} . . . . . . . . . . . . . . . . . . . . . . . . . 81
Heat-transfer characteristics of a non-Newtonian Au nanofluid in a cubical enclosure with differentially heated side walls
Zna~ilnosti prenosa toplote nenewtonske Au nanoteko~ine v kockastem ohi{ju z razli~no gretima stranskima stenama
P. Ternik, R. Rudolf, Z. @uni~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87
Amplitude–frequency response of an aluminium cantilever beam determined with piezoelectric transducers
Amplitudno-frekven~ni odziv konzolnega nosilca iz aluminija, ugotovljen s piezoelektri~nimi pretvorniki
Z. La{ová, R. Zem~ík . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95
Micromechanical model of the substituents of a unidirectional fiber-reinforced composite and its response to the tensile cyclic
loading
Mikromehanski model nadomestkov kompozitov, oja~anih z enosmernimi vlakni, in njihov odgovor na cikli~no natezno obremenjevanje
T. Kroupa, H. Srbová, R. Zem~ík . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 99
996 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
LETNO KAZALO – INDEX
Dry-cutting options with a chainsaw at the Hotavlje I natural-stone quarry
Mo`nosti suhega rezanja z veri`no `ago v kamnolomu naravnega kamna Hotavlje I.
J. Kortnik, B. Markoli . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103
Effect of sliding speed on the frictional behavior and wear performance of borided and plasma-nitrided W9Mo3Cr4V
high-speed steel
Vpliv hitrosti drsenja na vedenje in obrabo boriranega in v plazmi nitriranega hitroreznega jekla W9Mo3Cr4V
I. Gunes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111
Tribological behaviour of A356/10SiC/3Gr hybrid composite in dry-sliding conditions
Tribolo{ko vedenje hibridnega kompozita A356/10SiC/3Gr pri suhem drsenju
B. Stojanovi}, M. Babi}, N. Miloradovi}, S. Mitrovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 117
New solid-polymer-electrolyte material for dye-sensitized solar cells
Novi elektrolitni material na osnovi trdnega polimera za son~ne celice, ob~utljive za svetlobo
V. K. Singh, B. Bhattacharya, S. Shukla, P. K. Singh . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 123
Compacting the powder of Al-Cr-Mn alloy with SPS
Kompaktiranje prahu zlitine Al-Cr-Mn s SPS
T. F. Kubatík, Z. Pala, P. Novák. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 129
Effect of tempering on the microstructure and mechanical properties of resistance-spot-welded DP980 dual-phase steel
Vpliv popu{~anja na mikrostrukturo in mehanske lastnosti to~kasto varjenega dvofaznega jekla DP980
F. Nikoosohbat, S. Kheirandish, M. Goodarzi, M. Pouranvari . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 133
Optimization of the process parameters for surface roughness and tool life in face milling using the Taguchi analysis
Optimizacija procesnih parametrov glede na hrapavost povr{ine in trajnostno dobo orodja pri ~elnem rezkanju z uporabo Taguchijeve analize
M. Sarýkaya, H. Dilipak, A. Gezgin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 139
Thin tin monosulfide films deposited with the HVE method for photovoltaic applications
Tanka plast HVE kositrovega monosulfida za uporabo v fotovoltaiki
N. Revathi, S. Bereznev, J. Lehner, R. Traksmaa, M. Safonova, E. Mellikov, O. Volobujeva . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 149
Mechanical properties of the austenitic stainless steel X15CrNiSi20-12 after recycling
Mehanske lastnosti avstenitnega nerjavnega jekla X15CrNiSi20-12 po recikliranju
A. Deli}, O. Kablar, A. Zuli}, D. Kova~evi}, N. Hod`i}, E. Bar~i}. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 153
Comparison of refractory coatings based on talc, cordierite, zircon and mullite fillers for lost-foam casting
Primerjava ognjevzdr`nih premazov na osnovi smukca, kordierita, cirkona in mulitnih polnil za ulivanje v forme z izparljivim modelom
Z. A}imovi}-Pavlovi}, A. Terzi}, Lj. Andri}, M. Pavlovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 157
Selection of the most appropriate welding technology for hardfacing of bucket teeth
Izbira najbolj primerne tehnologije trdega navarjanja zoba zajemalke
V. Lazi}, A. Sedmak, R. R. Nikoli}, M. Mutavd`i}, S. Aleksandrovi}, B. Krsti}, D. Milosavljevi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 165
Characterization of TiO2 nanoparticles with high-resolution FEG scanning electron microscopy
Karakterizacija nanodelcev TiO2 z visokolo~ljivostno FEG vrsti~no elektronsko mikroskopijo
Z. Samard`ija, D. Lapornik, K. Gradi{ek, D. Verhov{ek, M. ^eh . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 173
2015/2
Editor’s Preface / Predgovor urednika
M. Torkar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 181
Pitting corrosion of TiN-coated stainless steel in 3 % NaCl solution
Jami~asta korozija nerjavnega jekla s prevleko TiN v 3-odstotni raztopini NaCl
I. Kucuk, C. Sarioglu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 183
Design of a wideband planar antenna on an epoxy-resin-reinforced woven-glass material
[irokopasovna ploskovna antena na epoksi smoli, oja~ani s steklenimi vlakni
R. Azim, M. T. Islam . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 193
Influence of the strain rate on the PLC effect and acoustic emission in single crystals of the CuZn30 alloy compressed at an
elevated temperature
Vpliv hitrosti deformacije na pojav PLC in akusti~no emisijo monokristalov zlitine CuZn30, stiskane pri povi{ani temperaturi
W. Ozgowicz, B. Grzegorczyk, A. Pawe³ek, A. Pi¹tkowski, Z. Ranachowski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 197
Determination of elastic-plastic properties of Alporas foam at the cell-wall level using microscale-cantilever bending tests
Dolo~anje elasti~nih in plasti~nih lastnosti pene Alporas na ravni celi~ne stene z upogibnimi preizkusi z mikroskopsko iglo
T. Doktor, D. Kytýø, P. Koudelka, P. Zlámal, T. Fíla, O. Jirou{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 203
Electrochemical behavior of biocompatible alloys
Elektrokemijsko vedenje biokompatibilnih zlitin
I. Petrá{ová, M. Losertová. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 207
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 997
LETNO KAZALO – INDEX
Preparation and thermal stability of ultra-fine and nano-grained commercially pure titanium wires using CONFORM
equipment
Priprava komercialne ultradrobne in nanozrnate Ti-`ice z opremo CONFORM in njena termi~na stabilnost
T. Kubina, J. Dlouhý, M. Köver, M. Dománková, J. Hodek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 213
Estimation of the thermal contact conductance from unsteady temperature measurements
Dolo~anje kontaktne toplotne prevodnosti iz neravnote`nega merjenja temperature
J. Kvapil, M. Pohanka, J. Horský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 219
Potentiodynamic and XPS studies of X10CrNi18-8 steel after ethylene oxide sterilization
Potenciodinami~ne in XPS analize jekla X10CrNi18-8 po sterilizaciji z etilen oksidom
W. Walke, J. Przondziono . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 223
TEM replica of a fluoride-miserite glass-ceramic glaze microstructure
TEM-replike mikrostrukture steklokerami~ne fluor-mizeritne glazure
J. Ma. Rincón, R. Casasola . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 229
Experimental analysis and modeling of the buckling of a loaded honeycomb sandwich composite
Eksperimentalna analiza in modeliranje upogibanja obremenjenega satastega sendvi~nega kompozita
A. Bentouhami, B. Keskes. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 235
Fatigue behaviour of X70 steel in crude oil
Vedenje jekla X70 pri utrujanju v surovi nafti
¼. Gajdo{, M. [perl, J. Bystrianský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 243
Use of Larson-Miller parameter for modeling the progress of isothermal solidification during transient-liquid-phase bonding
of IN718 superalloy
Uporaba Larson-Millerjevega parametra za modeliranje izotermnega strjevanja pri spajanju z vmesno teko~o fazo superzlitine IN718
M. Pouranvari, S. M. Mousavizadeh . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 247
Effect of severe air-blast shot peening on the wear characteristics of CP titanium
Vpliv intenzivnega povr{inskega kovanja s peskanjem z zrakom na obrabne lastnosti CP-titana
A. C. Karaoglanli . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 253
Finite-element minimization of the welding distortion of dissimilar joints of carbon steel and stainless steel
Uporaba kon~nih elementov za zmanj{anje popa~enja oblike pri varjenju ogljikovega in nerjavnega jekla
E. Ranjbarnodeh, M. Pouranvari, M. Farajpour . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 259
Magnesium-alloy die forgings for automotive applications
Izkovki iz magnezijevih zlitin za avtomobilsko industrijo
M. Madaj, M. Greger, V. Karas . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 267
Resistance to electrochemical corrosion of the extruded magnesium alloy AZ80 in NaCl solutions
Odpornost ekstrudirane magnezijeve zlitine AZ80 proti elektrokemijski koroziji v raztopini NaCl
J. Przondziono, E. Hadasik, W. Walke, J. Szala, J. Michalska, J. Wieczorek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 275
Microwave-assisted hydrothermal synthesis of Ag/ZnO sub-microparticles
Hidrotermi~na sinteza podmikrometrskih delcev Ag/ZnO z mikrovalovi
L. Münster, P. Ba`ant, M. Machovský, I. Kuøitka. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 281
Determination of the cause of the formation of transverse internal cracks on a continuously cast slab
Ugotavljanje vzrokov za nastanek notranjih pre~nih razpok v kontinuirno ulitem slabu
Z. Franìk, M. Masarik, J. [míd . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 285
Effective preparation of non-linear material models using a programmed optimization script for a nurimerical simulation of
sheet-metal processing
U~inkovita priprava nelinearnih modelov materiala s programiranim optimizacijskim zapisom za numeri~no simulacijo obdelave plo~evine
M. Urbánek, F. Tikal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 291
Neutralization of waste filter dust with CO2
Nevtralizacija odpadnega filtrskega prahu s CO2
A. Kra~un, I. An`el, L. Fras Zemlji~, A. Stergar{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 297
The influence of the morphology of iron powder particles on their compaction in an automatic die
Vpliv morfologije delcev `elezovega prahu na njegovo sposobnost za avtomatsko enoosno stiskanje
B. [u{tar{i~, M. Godec, ^. Donik, I. Paulin, S. Glode`, M. [ori, M. Ratej, N. Javornik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 303
2015/3
Wear mechanisms and surface engineering of forming tools
Obrabni mehanizmi in in`eniring povr{ine preoblikovalnih orodij
B. Podgornik, V. Leskov{ek. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 313
Predicting the physical properties of drawn Nylon-6 fibers using an artificial-neural-network model
Napovedovanje fizikalnih lastnosti vle~enih vlaken iz najlona 6 z uporabo modela umetne nevronske mre`e
R. Semnani Rahbar, M. Vadood. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 325
998 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
LETNO KAZALO – INDEX
Influence of the impact angle and pressure on the spray cooling of vertically moving hot steel surfaces
Vpliv vpadnega kota in tlaka na ohlajanje z brizganjem na vertikalno premikajo~e se vro~e povr{ine jekla
M. Hnízdil, M. Chabi~ovský, M. Raudenský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 333
Techniques of measuring spray-cooling homogeneity
Tehnike merjenja homogenosti hlajenja z brizganjem
M. Chabi~ovský, M. Raudenský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 337
Mineralogical and geochemical characterization of roman slag from the archaeological site near Mo{nje (Slovenia)
Mineralo{ka in geokemi~na karakterizacija rimske `lindre z arheolo{kega najdi{~a pri Mo{njah (Slovenija)
S. Kramar, J. Lux, H. Pristacz, B. Mirti~, N. Rogan - [muc . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 343
Reynolds differential equation singularity using processes of small straining with lubrication
Reynoldsova diferencialna ena~ba pri procesih majhne deformacije z mazanjem
D. ]ur~ija, F. Vodopivec, I. Mamuzi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 349
Prediction of the catastrophic tool failure in hard turning through acoustic emission
Napovedovanje katastrofi~ne po{kodbe kerami~nih vlo`kov pri stru`enju z akusti~no emisijo
M. ^illiková, B. Mi~ieta, M. Neslu{an, R. ^ep, I. Mrkvica, J. Petrù, T. Zlámal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 355
Erosive wear resistance of silicon carbide-cordierite ceramics: influence of the cordierite content
Odpornost keramike silicijev karbid-kordierit proti obrabi pri eroziji: vpliv vsebnosti kordierita
M. Po{arac-Markovi}, Dj. Veljovi}, A. Deve~erski, B. Matovi}, T. Volkov-Husovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 365
Influence of the substrate temperature on the structural, optical and thermoelectric properties of sprayed V2O5 thin films
Vpliv temperature podlage na strukturne, opti~ne in termoelektri~ne lastnosti napr{ene tanke plasti V2O5
Y. Vijayakumar, K. N. Reddy, A. V. Moholkar, M. V. R. Reddy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 371
Deep micro-hole drilling for hadfield steel by electro-discharge machining (EDM)
Vrtanje globokih mikrolukenj v jekla hadfield z metodo elektrorazreza (EDM)
V. Yilmaz, M. Sarýkaya, H. Dilipak . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 377
Surface analysis of electrochromic CuxO films in their colored and bleached states
Povr{inska analiza elektrokromiznih plasti CuxO v njihovih obarvanih in obeljenih stanjih
M. M. Ristova, M. Milun, B. Pejova . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 387
Thermodynamic analysis of the precipitation of carbonitrides in microalloyed steels
Termodinamska analiza izlo~anja karbonitridov v mikrolegiranih jeklih
M. Opiela . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 395
Experimental investigation of the crack-initiation moment of Charpy specimens under impact loading
Eksperimentalna preiskava trenutka iniciacije razpoke pri udarni obremenitvi Charpyjevih vzorcev
V. Kharchenko, E. Kondryakov, A. Panasenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 403
Structural, thermal and magnetic properties of Fe-Co-Ni-B-Si-Nb bulk amorphous alloy
Strukturne, termi~ne in magnetne lastnosti masivne amorfne zlitine Fe-Co-Ni-B-Si-Nb
S. Lesz, M. Nabia³ek, R. Nowosielski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 409
The nano-wetting aspect at the liquid-metal/SiC interface
Vidik nanoomakanja na stiku staljena kovina-SiC
M. Mihailovi}, K. Rai}, A. Patari}, T. Volkov - Husovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 413
The effect of a superplasticizer admixture on the mechanical fracture parameters of concrete
Vpliv dodatka superplastifikatorja na parametre mehanskega zloma betona
H. [imonová, I. Havlíková, P. Danìk, Z. Ker{ner, T. Vymazal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 417
Properties and structure of Cu-Ti-Zr-Ni amorphous powders prepared by mechanical alloying
Lastnosti in struktura amorfnih prahov Cu-Ti-Zr-Ni, pripravljenih z mehanskim legiranjem
A. Guwer, R. Nowosielski, A. Lebuda . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 423
Interaction of Cr2N and Cr2N/Ag thin films with CuZn-brass counterpart during ball-on-disc testing
Interakcija Cr2N in Cr2N/Ag tankih plasti v paru s CuZn-medenino med preizkusom krogla na disk
P. Bílek, P. Jur~i, P. Dulová, M. Hudáková, J. Pta~inová, M. Pa{ák. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 429
Using simulated spectra to test the efficiency of spectral processing software in reducing the noise in Auger electron spectra
Uporaba simuliranega spektra za preizkus u~inkovitosti programske opreme predelave spektra pri zmanj{anju {uma spektra Augerjevih
elektronov
B. Poniku, I. Beli~, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 435
Surface behavior of AISI 4140 modified with the pulsed-plasma technique
Lastnosti povr{ine AISI 4140, spremenjene s tehniko pulzirajo~e plazme
Y. Y. Özbek, M. Durman . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 441
Preparation and dielectric properties of thermo-vaporous BaTiO3 ceramics
Priprava in dielektri~ne lastnosti termo-parno porozne keramike BaTiO3
A. Kholodkova, M. Danchevskaya, N. Popova, L. Pavlyukova, A. Fionov . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 447
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 999
LETNO KAZALO – INDEX
Hybrid sol-gel coatings doped with cerium to protect magnesium alloys from corrosion
Hibridni sol-gel-nanosi, dopirani s cerijem, za korozijsko za{~ito magnezijevih zlitin
N. V. Murillo-Gutiérrez, F. Ansart, J.-P. Bonino, M.-J. Menu, M. Gressier . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 453
Influence of the tool geometry and process parameters on the static strength and hardness of friction-stir spot-welded
aluminium-alloy sheets
Vpliv geometrije orodja in parametrov procesa na stati~no trdnost in trdoto pri vrtilno-tornem to~kastem varjenju plo~evin iz Al-zlitine
H. Güler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 457
The stabilization of nano silver on polyester filament for a machine-made carpet
Stabilizacija nanodelcev srebra na poliestrskem vlaknu za strojno izdelavo preprog
K. M. Shojaei, A. Farrahi, H. Farrahi, A. Farrahi . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 461
Evaluation of the thermal resistance of selected bentonite binders
Ocena toplotne upornosti izbranih bentonitnih veziv
J. Beòo, J. Vontorová, V. Matìjka, K. Gál . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 465
Development of numerical models for the heat-treatment-process optimisation in a closed-die forging production
Razvoj numeri~nih modelov za optimizacijo postopka toplotne obdelave pri proizvodnji odkovkov v zaprtih utopnih orodjih
L. Male~ek, M. Fedorko, F. Van~ura, H. Jirková, B. Ma{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 471
Alojz Pre{ern, dipl. in`. metalurgije (1920–2015) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 477
2015/4
Impact-toughness investigations of duplex stainless steels
Preiskave udarne `ilavosti dupleksnega nerjavnega jekla
S. Topolska, J. £abanowski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 481
Effects of the artificial-aging temperature and time on the mechanical properties and springback behavior of AA6061
Vpliv temperature in ~asa umetnega staranja na mehanske lastnosti in vzmetnost AA6061
A. Polat, M. Avsar, F. Ozturk . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 487
Monitoring of polyurethane dispersions after the synthesis
Spremljanje poliuretanskih disperzij po sintezi
M. Ocepek, J. Zabret, J. Kecelj, P. Venturini, J. Golob . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 495
Spectroscopic and porosimetric analyses of Roman pottery from an archaeological site near Mo{nje, Slovenia
Spektroskopske in porozimetri~ne preiskave rimske lon~enine z arheolo{kega najdi{~a pri Mo{njah, Slovenija
S. Kramar, J. Lux . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 503
Non-linear finite-element simulations of the tensile tests of textile composites
Nelinearna simulacija nateznih preizkusov tekstilnih kompozitov s kon~nimi elementi
T. Kroupa, K. Kunc, R. Zem~ík, T. Mandys. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 509
Influence of the type and number of prepreg layers on the flexural strength and fatigue life of honeycomb sandwich structures
Vpliv vrste in {tevila plasti na utrditev upogibne trdnosti in zdr`ljivosti pri utrujanju satastih sendvi~nih konstrukcij
L. Fojtl, S. Rusnakova, M. Zaludek, V. Rusnák . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 515
Potential for obtaining an ultrafine microstructure of low-carbon steel using accumulative roll bonding
Mo`nosti doseganja ultradrobnozrnate mikrostrukture pri spajanju malooglji~nega jekla z akumulativnim valjanjem
T. Kubina, J. Gubi{ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 521
Optimization of the annealing of plaster moulds for the manufacture of metallic foams with an irregular cell structure
Optimiranje postopka `arjenja mav~nih form za izdelavo kovinskih pen z nepravilno strukturo celic
I. Kroupová, P. Lichý, F. Radkovský, J. Beòo, V. Bednáøová, I. Lána . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 527
The kinetics of small-impurity grain-boundary-segregation formation in cold-rolled deep-drawing 08C-Al and IF steels during
post-deformation annealing
Kinetika nastanka segregacije ne~isto~ po mejah zrn med `arjenjem po hladnem valjanju jekel 08C-Al in IF za globoki vlek
A. Rashkovskiy, A. Kovalev, D. Wainstein, I. Rodionova, Y. Bykova, D. Zakharova . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 531
Application of a control-measuring apparatus and peltier modules in the bulk-metallic-glass production using the
pressure-casting method
Uporaba kontrolno-merilne naprave in peltierjevih modulov pri izdelavi masivnih kovinskih stekel po postopku tla~nega litja
W. Pilarczyk, A. Pilarczyk. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 537
Evaluation of the structural changes in 9 % Cr creep-resistant steel using an electrochemical technique
Ocena sprememb strukture v jeklu z 9 % Cr, odpornem proti lezenju, z uporabo elektrokemijske tehnike
J. Rapouch, J. Bystriansky, V. Sefl, M. Svobodova. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 543
Effect of the by-pass cement-kiln dust and fluidized-bed-combustion fly ash on the properties of fine-grained alkali-activated
slag-based composites
Vpliv prahu iz pe~i za cement in lete~ega pepela iz vrtin~aste plasti na lastnosti drobnozrnatega, z alkalijami aktiviranega kompozita na
osnovi `lindre
V. Bílek Jr., L. Paøízek, L. Kalina . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 549
1000 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
LETNO KAZALO – INDEX
Microstructure, magnetic and mechanical properties of the bulk amorphous alloy Fe61Co10Ti4Y5B20
Mikrostruktura, magnetne in mehanske lastnosti masivne amorfne zlitine Fe61Co10Ti4Y5B20
K. Bloch, M. Nabia³ek, J. Gondro . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 553
Modified cement-based mortars: crack initiation and volume changes
Modificirane malte na osnovi cementa: iniciacija razpok in volumenske spremembe
I. Havlikova, V. Bilek Jr., L. Topolar, H. Simonova, P. Schmid, Z. Kersner. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 557
Fracture properties of plain and steel-polypropylene-fiber-reinforced high-performance concrete
Lastnosti loma navadnega in visokozmogljivega betona, oja~anega s polipropilenskimi vlakni
P. Smarzewski, D. Barnat-Hunek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 563
Preparation of porous ceramic materials based on CaZrO3
Priprava porozne keramike na osnovi CaZrO3
E. Œnie¿ek, J. Szczerba, I. Jastrzêbska, E. Kleczyk, Z. Pêdzich . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 573
DP780 dual-phase-steel spot welds: critical fusion-zone size ensuring the pull-out failure mode
To~kasti zvari jekla DP780 z dvofazno strukturo: kriti~na velikost staljene cone, ki zagotavlja poru{enje z izpuljenjem
M. Pouranvari, S. P. H. Marashi, H. L. Jaber . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 579
Application of computed tomography in comparison with the standardized methods for determining the permeability of
cement-composite structures
Uporaba ra~unalni{ke tomografije v primerjavi s standardiziranimi metodami dolo~anja prepustnosti cementnih kompozitnih struktur
T. Komárková, M. Králíková, P. Kovács, D. Kocáb, T. Stavaø . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 587
Properties of polymer-filled aluminium foams under moderate strain-rate loading conditions
Lastnosti aluminijevih pen, napolnjenih s polimernimi materiali, pri zmernih obremenitvah
T. Doktor, P. Zlámal, T. Fíla, P. Koudelka, D. Kytýø, O. Jirou{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 597
Quality of the structure of ash bodies based on different types of ash
Kvaliteta strukture telesa iz pepela na osnovi razli~nih vrst pepela
V. Cerny . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 601
Sintered board materials based on recycled glass
Sintrani plo{~ati materiali na osnovi recikliranega stekla
T. Melichar, J. Byd`ovský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 607
Mechanical and wetting properties of nanosilica/epoxy-coated stainless steel
Mehanske in povr{inske lastnosti premaza iz silicijevih nanodelcev in epoksidne smole na nerjavnem jeklu
M. Conradi, G. Intihar, M. Zorko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 613
Development of composite salt cores for foundry applications
Razvoj kompozitnih slanih jeder za uporabo v livarstvu
J. Beòo, E. Adámková, F. Mik{ovský, P. Jelínek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 619
Carbide morphology and ferrite grain size after accelerated carbide spheroidisation and refinement (ASR) of C45 steel
Morfologija karbidov in velikost feritnih zrn po pospe{eni sferoidizaciji in rafinaciji (ASR) jekla C45
J. Dlouhy, D. Hauserova, Z. Novy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 625
Use of micromachining to shape the structure and electrical properties of the front electrode of a silicon solar cell
Uporaba mikroobdelovanja za oblikovanje strukture in elektri~nih lastnosti prednje elektrode silicijeve son~ne celice
M. Musztyfaga-Staszuk . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 629
Fabrication of TiO2 nanotubes for bioapplications
Izdelava TiO2-nanocevk za biomedicinsko uporabo
M. Kulkarni, K. Mrak-Polj{ak, A. Fla{ker, A. Mazare, P. Schmuki, A. Kos, S. ^u~nik, S. Sodin-[emrl, A. Igli~. . . . . . . . . . . . . . . . . . . 635
Assessment of the impact-echo method for monitoring the long-standing frost resistance of ceramic tiles
Ocena metode impact-echo za kontrolo dolgotrajne odpornosti kerami~nih plo{~ic proti zmrzali
M. Matysik, I. Plskova, Z. Chobola . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 639
Investigation on new creep- and oxidation-resistant materials
Preiskava novega materiala, odpornega proti lezenju in oksidaciji
O. Khalaj, B. Masek, H. Jirkova, A. Ronesova, J. Svoboda . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 645
Effect of preheating on mechanical properties in induction sintering of metal-powder material Fe and w(Cu) = 3 %
Vpliv predgrevanja na mehanske lastnosti indukcijsko sintranega materiala, izdelanega iz kovinskega prahu Fe in w(Cu) = 3 %
G. Akpýnar, E. Atik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 653
2015/5
Kinetic study and characterization of borided AISI 4140 steel
[tudij kinetike in karakterizacija boriranega jekla AISI 4140
M. Keddam, M. Ortiz-Domínguez, O. A. Gómez-Vargas, A. Arenas-Flores, M. Á. Flores-Rentería, M. Elias-Espinosa,
A. García-Barrientos . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 665
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 1001
LETNO KAZALO – INDEX
Kinetics of the precipitation in austenite HSLA steels
Kinetika izlo~anja v avstenitnih HSLA-jeklih
E. Kalinowska-Ozgowicz, R. Kuziak, W. Ozgowicz, K. Lenik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 673
Combined influence of V and Cr on the AlSi10MgMn alloy with a high Fe level
Vzajemni vpliv V in Cr na zlitino AlSi10MgMn z visoko vsebnostjo Fe
D. Bolibruchová, M. @ihalová . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 681
A reliable approach to a rapid calculation of the grain size of polycrystalline thin films after excimer laser crystallization
Zanesljiv na~in hitrega izra~una velikosti zrn v polikristalni tanki plasti po UV-laserski kristalizaciji
C. C. Kuo . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 687
Effects of different stirrer-pin forms on the joining quality obtained with friction-stir welding
Vpliv razli~nih oblik vrtilnih konic na kvaliteto spoja pri tornem vrtilnem varjenju
H. Basak, K. Kaptan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 693
Monitoring early-age concrete with the acoustic-emission method and determining the change in the electrical properties
Pregled sve`ega betona z metodo akusti~ne emisije in dolo~anjem sprememb elektri~nih lastnosti
L. Pazdera, L. Topolar, M. Korenska, T. Vymazal, J. Smutny, V. Bilek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 703
Effect of the aggregate type on the properties of alkali-activated slag subjected to high temperatures
Vpliv vrste agregata na lastnosti z alkalijo aktivirane `lindre, izpostavljene visokim temperaturam
P. Rovnaník, Á. Dufka. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 709
Microstructure evolution of advanced high-strength TRIP-aided bainitic steel
Razvoj mikrostrukture naprednega visokotrdnostnega bainitnega jekla z uporabo TRIP
A. Grajcar. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 715
Effect of electric current on the production of NiTi intermetallics via electric-current-activated sintering
Vpliv elektri~nega toka pri izdelavi intermetalne zlitine NiTi s sintranjem, aktiviranim z elektri~nim tokom
T. Yener, S. Siddique, F. Walther, S. Zeytin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 721
Control of soft reduction of continuous slab casting with a thermal model
Kontrola mehke redukcije pri kontinuirnem litju slabov s termi~nim modelom
J. Stetina, P. Ramik, J. Katolicky . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 725
Optimization of operating conditions in a laboratory SOFC testing device
Optimizacija obratovalnih razmer laboratorijske gorivne celice SOFC
T. Skalar, M. Lubej, M. Marin{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 731
Production of shaped semi-products from AHS steels by internal pressure
Izdelava polproizvodov iz AHS-jekel, oblikovanih z notranjim tlakom
I. Vorel, H. Jirková, B. Ma{ek, P. Kurka . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 739
Composite-material printed antenna for a multi-standard wireless application
Tiskana antena iz kompozitnega materiala za ve~standardno brez`i~no uporabo
T. Alam, M. R. I. Faruque, M. T. Islam, N. Misran . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 745
Friction and wear behaviour of ulexite and cashew in automotive brake pads
Odpornost proti trenju in obrabi avtomobilskih zavornih oblog z uleksitom in prahom iz indijskega oreha
Ý. Sugözü, Ý. Mutlu, A. Keskin. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 751
Diffusion kinetics and characterization of borided AISI H10 steel
Kinetika difuzije in karakterizacija boriranega jekla AISI H10
I. Gunes, M. Ozcatal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 759
Application of the Taguchi method to optimize the cutting conditions in hard turning of a ring bore
Uporaba Taguchi-jeve metode za optimizacijo trdega stru`enja roba izvrtine
M. Boy, I. Ciftci, M. Gunay, F. Ozhan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 765
Thermal fatigue of single-crystal superalloys: experiments, crack-initiation and crack-propagation criteria
Toplotno utrujanje monokristalnih superzlitin: preizkusi, merila za nastanek in napredovanje razpoke
L. Getsov, A. Semenov, S. Semenov, A. Rybnikov, E. Tikhomirova . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 773
Investigating the effects of cutting parameters on the built-up-layer and built-up-edge formation during the machining of AISI
310 austenitic stainless steels
Preiskava vplivov parametrov rezanja na nastanek nakopi~ene plasti in nakopi~enega roba med stru`enjem avstenitnega nerjavnega
jekla AISI 310
M. B. Bilgin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 779
Mortar-type identification for the purpose of reconstructing fragmented Roman wall paintings (Celje, Slovenia)
Analiza ometov za rekonstrukcijo fragmentov rimskih stenskih poslikav (Celje, Slovenija)
M. Gutman, M. Lesar Kikelj, J. Kuret, S. Kramar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 785
1002 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
LETNO KAZALO – INDEX
Au-nanoparticle synthesis via ultrasonic spray pyrolysis with a separate evaporation zone
Sinteza nanodelcev zlata z ultrazvo~no razpr{ilno pirolizo z lo~eno cono izhlapevanja
P. Majeri~, B. Friedrich, R. Rudolf . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 791
Determination of the critical fraction of solid during the solidification of a PM-cast aluminium alloy
Dolo~anje kriti~nega dele`a strjene faze med strjevanjem v PM ulite aluminijeve zlitine
R. Kayikci, M. Colak, S. Sirin, E. Kocaman, N. Akar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 797
Microstructural evolution of Inconel 625 during hot rolling
Mikrostrukturni razvoj Inconela 625 med vro~im valjanjem
F. Tehovnik, J. Burja, B. Podgornik, M. Godec, F. Vode. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 801
Thermophysical properties and microstructure of magnesium alloys of the Mg-Al type
Termofizikalne lastnosti in mikrostruktura magnezijevih zlitin tipa Mg-Al
P. Lichý, J. Beòo, I. Kroupová, I. Vasková . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 807
Micro-encapsulated phase-change materials for latent-heat storage: thermal characteristics
Mikroenkapsulirani materiali s fazno premeno za shranjevanje latentne toplote: toplotne zna~ilnosti
M. Ostrý, D. Dostálová, T. Klubal, R. Pøikryl, P. Charvát. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 813
Ceramic masonry units intended for the masonry resistant to high humidity
Kerami~ni gradbeni elementi, namenjeni za zgradbe, odporne proti visoki vlagi
J. Zach, V. Novák, J. Hroudová, M. Sedlmajer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 817
Flame resistance and mechanical properties of composites based on new advanced resin system FR4/12
Negorljivost in mehanske lastnosti kompozitov na osnovi novih naprednih sistemov smol FR4/12
V. Rusnák, S. Rusnáková, L. Fojtl, M. @aludek, A. ^apka . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 821
Effect of process parameters on the microstructure and mechanical properties of friction-welded joints of AISI 1040/AISI
304L steels
Vpliv procesnih parametrov na mikrostrukturo in mehanske lastnosti torno varjenih spojev jekel AISI 1040/AISI 304L
Ý. Kirik, N. Özdemýr . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 825
Influence of inoculation methods and the amount of an added inoculant on the mechanical properties of ductile iron
Vpliv metod modifikacije in koli~ine dodanega modifikatorja na mehanske lastnosti duktilnega `eleza
H. Avdu{inovi}, A. Gigovi}-Geki}, D. ]ubela, R. Sunulahpa{i}, N. Mujezinovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 833
One-step green synthesis of graphene/ZnO nanocomposites for non-enzymatic hydrogen peroxide sensing
Enostopenjska zelena sinteza nanokompozita grafen-ZnO za neencimatsko detekcijo vodikovega peroksida
S. S. Low, M. T. T. Tan, P. S. Khiew, H. S. Loh, W. S. Chiu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 837
Material parameters for a numerical simulation of a compaction process for sintered double-height gears
Materialni parametri za numeri~no simulacijo postopka stiskanja sintranih dvovi{inskih zobnikov
T. Verlak, M. [ori, S. Glode` . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 841
2015/6
Problems and normative evaluation of bond-strength tests for coated reinforcement and concrete
Problemi in normativna ocena preizkusov trdnosti vezi med armaturo s prekritjem in betonom
P. Pokorný, M. Kouøil, J. Stoulil, P. Bou{ka, P. Simon, P. Juránek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 847
Modeling of occurrence of surface defects of C45 steel with genetic programming
Modeliranje pojava povr{inskih napak pri jeklu C45 z genetskim programiranjem
M. Kova~i~, R. Jager . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 857
Effects of various helically angled grinding wheels on the surface roughness and roundness in grinding cylindrical surfaces
Vpliv razli~nih kotov vija~nice pri brusilnih kolutih na hrapavost povr{ine in okroglost pri bru{enju valjastih povr{in
M. Gavas, M. Kýna, U. Köklü . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 865
Characterization of cast-iron gradient castings
Karakterizacija lito`eleznega gradientnega ulitka
D. Mitrovi}, P. Mrvar, M. Petri~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 871
Comparative mechanical and corrosion studies on magnesium, zinc and iron alloys as biodegradable metals
Primerjalna {tudija mehanskih in korozijskih lastnosti biorazgradljivih zlitin magnezija, cinka in `eleza
D. Vojtìch, J. Kubásek, J. ^apek, I. Pospí{ilová. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 877
Microstructural changes of fine-grained concrete exposed to a sulfate attack
Mikrostrukturne spremembe drobnozrnatega betona, izpostavljenega sulfatu
M. Vy{vaøil, P. Bayer, M. Rovnaníková . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 883
Effect of thermomechanical treatment on the corrosion behaviour of Si- and Al-containing high-Mn austenitic steel with Nb
and Ti micro-additions
Vpliv termomehanske obdelave na korozijsko vedenje manganskega avstenitnega jekla z vsebnostjo Si in Al, mikrolegiranega z Nb in Ti
A. Grajcar, A. Plachciñska, S. Topolska, M. Kciuk . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 889
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 1003
LETNO KAZALO – INDEX
Surface free energy of hydrophobic coatings of hybrid-fiber-reinforced high-performance concrete
Prosta energija povr{ine hidrofobnih premazov na visokozmogljivem betonu, oja~anem s hibridnimi vlakni
D. Barnat-Hunek, P. Smarzewski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 895
Developing continuous-casting-process control based on advanced mathematical modelling
Uporaba naprednega matemati~nega modeliranja za razvoj kontrole postopka kontinuirnega ulivanja
J. Falkus, K. Mi³kowska-Piszczek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 903
Elastic behaviour of magnesia-chrome refractories at elevated temperatures
Elasti~no vedenje ognjevzdr`nih gradiv magnezija-krom pri povi{anih temperaturah
I. Jastrzêbska, J. Szczerba, J. Szlêzak, E. Œnie¿ek, Z. Pêdzich . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 913
Study on the magnetization-reversal behavior of annealed Sm-Fe-Co-Si-Cu ribbons
[tudij vedenja pri obratu magnetizacije `arjenih trakov Sm-Fe-Co-Si-Cu
M. Doœpial, S. Garus, M. Nabialek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 919
Evaluation of the chitosan-coating effectiveness on a dental titanium alloy in terms of microbial and fibroblastic attachment
and the effect of aging
Ocena u~inkovitosti nanosa hitozana na oprijemanje mikrobov in fibroblastov na dentalni titanovi zlitini ter na pojav staranja
U. T. Kalyoncuoglu, B. Yilmaz, S. Gungor, Z. Evis, P. Uyar, G. Akca, G. Kansu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 925
Structure and properties of the carburised surface layer on 35CrSiMn5-5-4 steel after nanostructurization treatment
Struktura in lastnosti naoglji~ene povr{ine jekla 35CrSiMn5-5-4 po nanostrukturni obdelavi
E. Sko³ek, K. Wasiak, W. A. Œwi¹tnicki. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 933
Optimization of the surface roughness by applying the Taguchi technique for the turning of stainless steel under cooling
conditions
Uporaba Taguchi-jeve metode za optimiranje hrapavosti povr{ine pri stru`enju nerjavnega jekla z ohlajanjem
M. Sarýkaya . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 941
Effect of cryogenic treatment applied to M42 HSS drills on the machinability of Ti-6Al-4V alloy
Vpliv podhlajevanja svedrov M42 HSS na obdelovalnost zlitine Ti-6Al-4V
T. Kývak, U. ªeker . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 949
Load-capacity prediction for the carbon- or glass-fibre-reinforced plastic part of a wrapped pin joint
Napoved nosilnosti plasti~nih delov zati~nega spoja, oja~anega z ogljikovimi ali steklastimi vlakni
J. Krystek, R. Kottner . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 957
A meshless model of electromagnetic braking for the continuous casting of steel
Brezmre`ni model elektromagnetnega zaviranja pri kontinuiranem ulivanju jekla
K. Mramor, R. Vertnik, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 961
Non-singular method of fundamental solutions for three–dimensional isotropic elasticity problems with displacement boundary
conditions
Nesingularna metoda fundamentalnih re{itev za deformacijo tridimenzijskih elasti~nih problemov z deformacijskimi robnimi pogoji
Q. Liu, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 969
Solid-state sintering of (K0.5Na0.5)NbO3 synthesized from an alkali-carbonate-based low-temperature calcined powder
Sintranje v trdnem keramike (K0,5Na0,5)NbO3, sintetizirane iz nizkotemperaturno kalciniranega prahu, pripravljenega na osnovi alkalijskih
karbonatov
M. Feizpour, T. Ebadzadeh, D. Jenko. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 975
Stability of close-cell Al foams depending on the usage of different foaming agents
Stabilnost aluminijevih pen z zaprto poroznostjo glede na uporabo razli~nih penilnih sredstev
I. Paulin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 983
Evolution of the microstructure and magnetic properties of a cobalt-silicon-based alloy in the early stages of mechanical
milling
Razvoj mikrostrukture in magnetnih lastnosti zlitine Co-Si v za~etnem stadiju mehanskega legiranja
W. Rattanasakulthong, C. Sirisathitkul, P. F. Rogl . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 989
In memoriam prof. dr. Milan Trbi`an . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 993
Letnik 49 (2015), 1–6 – Volume 49 (2015), 1–6 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 995
1004 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
LETNO KAZALO – INDEX
MATERIALI IN TEHNOLOGIJE / MATERIALS AND TECHNOLOGY
AVTORSKO KAZALO / AUTHOR INDEX
LETNIK / VOLUME 49, 2015, 1–6, A–@
A
A}imovi}-Pavlovi} Z. 157
Adámková E. 61, 619
Akar N. 797
Akca G. 925
Akpýnar G. 653
Alam T. 745
Aleksandrovi} S. 165
Aleksi} Z. 81
Andri} Lj. 157
Ansart F. 453
An`el I. 75, 297
Arenas-Flores A. 665
Atik E. 653
Avdu{inovi} H. 833
Avsar M. 487
Azim R. 193
B
Babi} M. 117
Bar~i} E. 153
Barnat-Hunek D. 563, 895
Basak H. 693
Bayer P. 883
Ba`ant P. 55, 281
Bednáøová V. 527
Beli~ I. 435
Beòo J. 61, 465, 527,619, 807
Bentouhami A. 235
Bereznev S. 149
Bhattacharya B. 123
Bílek Jr. V. 549, 557
Bílek P. 429
Bilek V. 703
Bilgin M. B. 779
Bloch K. 553
Bogovi} J. 75
Bolibruchová D. 681
Bonino J.-P. 453
Bou{ka P. 847
Boy M. 765
Burja J. 801
Byd`ovský J. 607
Bykova Y. 531
Bystrianský J. 243, 543
C
Calleja W. 3
Casasola R. 229
Castaño V. M. 3
Cerny V. 601
Chabi~ovský M. 333, 337
Charvát P. 813
Chiu W. S. 837
Chobola Z. 639
Ciftci I. 765
Colak M. 797
Conradi M. 613
]
]ubela D. 833
]ur~ija D. 349
^
^apek J. 877
^apka A. 821
^egan T. 15, 27
^eh M. 173
^ep R. 355
^illiková M. 355
^u~nik S. 635
D
Danchevskaya M. 447
Danìk P. 417
Deli} A. 153
Deve~erski A. 365
Dharmalingam S. 9
Dilipak H. 139, 377
Dlouhý J. 213, 625
Doktor T. 203, 597
Dománková M. 213
Donik ^. 303
Doœpial M. 919
Dostálová D. 813
Dufka Á. 709
Dulová P. 429
Durman M. 441
\
\ordevi} I. 81
E
Ebadzadeh T. 975
Elias-Espinosa M. 665
Evis Z. 925
F
Falkus J. 903
Farajpour M. 259
Farrahi A. 461
Farrahi H. 461
Faruque M. R. I. 745
Fedorko M. 471
Feizpour M. 975
Fíla T. 203, 597
Fionov A. 447
Fla{ker A. 635
Flores-Rentería M. Á. 665
Fojtl L. 515, 821
Franìk Z. 285
Fras Zemlji~ L. 297
Friedrich B. 75, 791
G
Gajdo{ ¼. 243
Gál K. 465
Galindo M. 3
García-Barrientos A. 665
Garus S. 919
Gavas M. 865
Getsov L. 773
Gezgin A. 139
Gigovi}-Geki} A. 833
Glode` S. 303, 841
Godec M. 303, 801
Golob J. 495
Gómez-Vargas O. A. 665
Gondro J. 553
Goodarzi M. 133
Gowrisankar A. 9
Gradi{ek K. 173
Grajcar A. 715, 889
Greger M. 267
Gressier M. 453
Grzegorczyk B. 197
Gubi{ J. 521
Gunay M. 765
Gunes I. 111, 759
Gungor S. 925
Gutman M. 785
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 1005
LETNO KAZALO – INDEX
Guwer A. 423
Güler H. 457
H
Hadasik E. 275
Hauserova D. 625
Havlíková I. 417, 557
Hidalga de la J. 3
Hnízdil M. 333
Hodek J. 213
Hod`i} N. 153
Horský J. 219
Hroudová J. 817
Hudáková M. 429
I
Igli~ A. 635
Intihar G. 613
Islam M. T. 193, 745
J
Jaber H. L. 579
Jager R. 857
Jandova D. 37
Jastrzêbska I. 573, 913
Javornik N. 303
Jelínek P. 61, 619
Jenko D. 975
Jenko M. 81, 435
Jirková H. 471, 645, 739
Jirou{ek O. 203, 597
Jokanovi} V. 81
Jur~i P. 429
Juránek P. 847
Juøica J. 15, 27
K
Kablar O. 153
Kalina L. 549
Kalinowska-Ozgowicz E. 69, 673
Kalyoncuoglu U. T. 925
Kansu G. 925
Kaptan K. 693
Karaoglanli A. C. 253
Karas V. 267
Katolicky J. 725
Kayikci R. 797
Kciuk M. 889
Kecelj J. 495
Keddam M. 43, 665
Ker{ner Z. 417, 557
Keskes B. 235
Keskin A. 751
Khalaj O. 645
Kharchenko V. 403
Kheirandish S. 133
Khiew P. S. 837
Kholodkova A. 447
Kýna M. 865
Kirik Ý. 825
Kývak T. 949
Kleczyk E. 573
Klofá~ J. 55
Klubal T. 813
Kocáb D. 587
Kocaman E. 797
Komárková T. 587
Kondryakov E. 403
Korenska M. 703
Kortnik J. 103
Kos A. 635
Kottner R. 957
Kouba R. 43
Koudelka P. 203, 597
Kouøil M. 847
Kova~evi} D. 153
Kova~i~ M. 857
Kovács P. 587
Kovalev A. 531
Kra~un A. 297
Králíková M. 587
Kramar S. 343, 503, 785
Kroupa T. 509
Kroupa T. 99
Kroupová I. 527, 807
Krsti} B. 165
Krystek J. 957
Kubásek J. 877
Kubatík T. F. 129
Kubina T. 213, 521
Kucuk I. 19, 183
Kulkarni M. 635
Kunc K. 509
Kuo C. C. 687
Kuret J. 785
Kuøitka I. 55, 281
Kurka P. 739
Kursa M. 15
Kuziak R. 673
Kvapil J. 219
Kytýø D. 203, 597
Köklü U. 865
Köver M. 213
L
£abanowski J. 481
Lána I. 527
Lapornik D. 173
La{ová Z. 95
Lazi} V. 165
Lebuda A. 423
Lehner J. 149
Lenik K. 673
Lesar Kikelj M. 785
Leskov{ek V. 313
Lesz S. 409
Lichý P. 527, 807
Liu Q. 969
Loh H. S. 837
López F. 3
Losertová M. 207
Low S. S. 837
Lubej M. 731
Lux J. 343, 503
M
Machovský M. 281
Madaj M. 267
Majeri~ P. 75, 791
Malcharcziková J. 15
Male~ek L. 471
Mamuzi} I. 349
Mandys T. 509
Marashi S. P. H. 579
Marin{ek M. 731
Markoli B. 103
Martínek P. 31
Masarik M. 285
Ma{ek B. 471, 645, 739
Matìjka V. 27, 465
Matovi} B. 365
Matysik M. 639
Mazare A. 635
Melichar T. 607
Mellikov E. 149
Menu M.-J. 453
Mi~ieta B. 355
Michalska J. 275
Michenka V. 15
Mihailovi} M. 413
Mik{ovský F. 61, 619
Mi³kowska-Piszczek K. 903
Miloradovi} N. 117
Milosavljevi} D. 165
Milun M. 387
Mirti~ B. 343
Misran N. 745
Mitrovi} D. 871
Mitrovi} S. 117
Moholkar A. V. 371
Mousavizadeh S. M. 247
1006 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
LETNO KAZALO – INDEX
Mrak-Polj{ak K. 635
Mramor K. 961
Mrkvica I. 355
Mrvar P. 871
Mujezinovi} N. 833
Murillo-Gutiérrez N. V. 453
Musztyfaga-Staszuk M. 629
Mutavd`i} M. 165
Mutlu Ý. 751
Münster L. 281
N
Nabia³ek M. 409, 553, 919
Nacházel J. 31
Neslu{an M. 355
Nikoli} R. R. 165
Nikoosohbat F. 133
Niznanska J. 37
Novák P. 129
Novák V. 817
Nowosielski R. 409, 423
Novy Z. 625
O
Ocepek M. 495
Opiela M. 395
Ortiz-Domínguez M. 665
Ostrý M. 813
Ozcatal M. 759
Ozgowicz W. 69, 197, 673
Ozhan F. 765
Ozturk F. 487
Özbek Y. Y. 441
Özdemýr N. 825
P
Pala Z. 129
Palomino-Merino R. 3
Panasenko A. 403
Paøízek L. 549
Pastorek M. 55
Pa{ák M. 429
Patari} A. 413
Paulin I. 303, 983
Pawe³ek A. 197
Pavlovi} M. 157
Pavlyukova L. 447
Pazdera L. 703
Pêdzich Z. 573, 913
Pejova B. 387
Petlák D. 27
Petrá{ová I. 207
Petri~ M. 871
Petrù J. 355
Pi¹tkowski A. 197
Pilarczyk A. 537
Pilarczyk W. 537
Plachciñska A. 889
Plskova I. 639
Po{arac-Markovi} M. 365
Podaný P. 31
Podgornik B. 313, 801
Pohanka M. 219
Pohludka M. 15
Pokorný P. 847
Polat A. 487
Poniku B. 435
Popova N. 447
Pospí{ilová I. 877
Pouranvari M. 133, 247, 259, 579
Pristacz H. 343
Pøikryl R. 813
Przondziono J. 223, 275
Pta~inová J. 429
R
Radika N. 9
Radkovský F. 527
Rai} K. T. 81, 413
Ramik P. 725
Ranachowski Z. 197
Ranjbarnodeh E. 259
Rapouch J. 543
Rashkovskiy A. 531
Ratej M. 303
Rattanasakulthong W. 989
Raudenský M. 333
Raudenský M. 337
Reddy K. N. 371
Reddy M. V. R. 371
Revathi N. 149
Rincón J. Ma. 229
Ristova M. M. 387
Rodionova I. 531
Rogan - [muc N. 343
Rogl P. F. 989
Ronesova A. 645
Rovnaník P. 709
Rovnaníková M. 883
Rudolf R. 75, 81, 87, 791
Rusnák V. 515, 821
Rusnáková S. 515, 821
Rybnikov A. 773
S
Safonova M. 149
Samard`ija Z. 173
Sarýkaya M. 139, 377, 941
Sarioglu C. 19, 183
Schmid P. 557
Schmuki P. 635
Sedlák J. 55
Sedlmajer M. 817
Sedmak A. 165
Sefl V. 543
ªeker U. 949
Semenov A. 773
Semenov S. 773
Semnani Rahbar R. 325
Shojaei K. M. 461
Shukla S. 123
Siddique S. 721
Simon P. 847
Simonova H. 557
Singh P. K. 123
Singh V. K. 123
Sirin S. 797
Sirisathitkul C. 989
Skalar T. 731
Sko³ek E. 933
Skotnicová K. 27
Smarzewski P. 563, 895
Smetana B. 27
Smutny J. 703
Œnie¿ek E. 573, 913
Sodin-[emrl S. 635
Srbová H. 99
Stamenkovi} D. 81
Stavaø T. 587
Stergar{ek A. 297
Stetina J. 725
Stojanovi} B. 117
Stopi} S. 75
Stoulil J. 847
Subramanian R. 9
Suchmann P. 37
Sugözü Ý. 751
Sunulahpa{i} R. 833
Œwi¹tnicki W. A. 933
Svoboda J. 645
Svobodova M. 543
Szala J. 275
Szczerba J. 573, 913
Szlêzak J. 913
[
[arler B. 961, 969
[imonová H. 417
[míd J. 285
[ori M. 303, 841
[perl M. 243
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 1007
LETNO KAZALO – INDEX
[u{tar{i~ B. 303
T
Tan M. T. T. 837
Tehovnik F. 801
Ternik P. 87
Terzi} A. 157
Thiyagarajan T. 9
Tikal F. 291
Tikhomirova E. 773
Todorovi} A. 81
Topolar L. 557, 703
Topolska S. 481, 889
Torkar M. 181
Traksmaa R. 149
U
Urbánek M. 291
Uyar P. 925
V
Vadood M. 325
Van~ura F. 471
Vasková I. 807
Veljovi} Dj. 365
Venturini P. 495
Verhov{ek D. 173
Verlak T. 841
Vertnik R. 961
Vijayakumar Y. 371
Vode F. 801
Vodopivec F. 349
Vojtìch D. 877
Volkov - Husovi} T. 365, 413
Volobujeva O. 149
Vontorová J. 465
Vorel I. 739
Vy{vaøil M. 883
Vymazal T. 417, 703
W
Wainstein D. 531
Wajda W. 69
Walke W. 223, 275
Walther F. 721
Wasiak K. 933
Wieczorek J. 275
Y
Yener T. 721
Yilmaz B. 925
Yilmaz V. 377
Z
Zabret J. 495
Zach J. 817
Zakharova D. 531
Zaludek M. 515
Zem~ík R. 95, 99, 509
Zeytin S. 721
Zlámal P. 203, 597
Zlámal T. 355
Zorko M. 613
Zuli} A. 153
Zúñiga C. 3
@
@aludek M. 821
@ihalová M. 681
@uni~ Z. 87
1008 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
LETNO KAZALO – INDEX
MATERIALI IN TEHNOLOGIJE / MATERIALS AND
TECHNOLOGY
VSEBINSKO KAZALO / SUBJECT INDEX
LETNIK / VOLUME 49, 2015, 1–6
Kovinski materiali – Metallic materials
Boron-doped hydrogenated amorphous semiconductor MEMS
Z borom dopirani hidrogenirani amorfni polprevodnik MEMS
M. Galindo, C. Zúñiga, R. Palomino-Merino, F. López, W. Calleja, J. de la Hidalga, V. M. Castaño . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3
Wear behaviour of B4C reinforced hybrid aluminum-matrix composites
Vedenje hibridnega kompozita na osnovi aluminija, oja~anega z B4C, pri obrabi
T. Thiyagarajan, R. Subramanian, S. Dharmalingam, N. Radika, A. Gowrisankar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9
Influence of the HIP process on the properties of as-cast Ni-based alloys
Vpliv vro~ega izostatskega stiskanja na lastnosti Ni-zlitin z lito strukturo
J. Malcharcziková, M. Pohludka, V. Michenka, T. ^egan, J. Juøica, M. Kursa. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15
Corrosion of CrN-coated stainless steel in a NaCl solution (w = 3 %)
Korozija nerjavnega jekla s CrN-prevleko v raztopini NaCl (w = 3 %)
I. Kucuk, C. Sarioglu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19
Preparation and properties of master alloys Nb-Al and Ta-Al for melting and casting of -TiAl intermetallics
Priprava in lastnosti predzlitin Nb-Al in Ta-Al za taljenje in ulivanje intermetalnih zlitin -TiAl
J. Juøica, T. ^egan, K. Skotnicová, D. Petlák, B. Smetana, V. Matìjka . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27
Decreasing the carbonitride size and amount in austenitic steel with heat treatment and thermomechanical processing
Zmanj{anje velikosti karbonitridov v avstenitnem jeklu s toplotno obdelavo in termomehansko predelavo
P. Martínek, P. Podaný, J. Nacházel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 31
Deep cryogenic treatment of H11 hot-working tool steel
Globoka kriogenska obdelava orodnega jekla H11 za delo v vro~em
P. Suchmann, D. Jandova, J. Niznanska . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37
Numerical prediction of the compound layer growth during the gas nitriding of Fe-M binary alloys
Numeri~no napovedovanje rasti spojinske plasti med plinskim nitriranjem binarnih zlitin Fe-M
R. Kouba, M. Keddam. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43
Amplitude–frequency response of an aluminium cantilever beam determined with piezoelectric transducers
Amplitudno-frekven~ni odziv konzolnega nosilca iz aluminija, ugotovljen s piezoelektri~nimi pretvorniki
Z. La{ová, R. Zem~ík . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95
Effect of sliding speed on the frictional behavior and wear performance of borided and plasma-nitrided W9Mo3Cr4V
high-speed steel
Vpliv hitrosti drsenja na vedenje in obrabo boriranega in v plazmi nitriranega hitroreznega jekla W9Mo3Cr4V
I. Gunes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111
Tribological behaviour of A356/10SiC/3Gr hybrid composite in dry-sliding conditions
Tribolo{ko vedenje hibridnega kompozita A356/10SiC/3Gr pri suhem drsenju
B. Stojanovi}, M. Babi}, N. Miloradovi}, S. Mitrovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 117
Compacting the powder of Al-Cr-Mn alloy with SPS
Kompaktiranje prahu zlitine Al-Cr-Mn s SPS
T. F. Kubatík, Z. Pala, P. Novák. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 129
Effect of tempering on the microstructure and mechanical properties of resistance-spot-welded DP980 dual-phase steel
Vpliv popu{~anja na mikrostrukturo in mehanske lastnosti to~kasto varjenega dvofaznega jekla DP980
F. Nikoosohbat, S. Kheirandish, M. Goodarzi, M. Pouranvari . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 133
Optimization of the process parameters for surface roughness and tool life in face milling using the Taguchi analysis
Optimizacija procesnih parametrov glede na hrapavost povr{ine in trajnostno dobo orodja pri ~elnem rezkanju z uporabo Taguchijeve analize
M. Sarýkaya, H. Dilipak, A. Gezgin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 139
Mechanical properties of the austenitic stainless steel X15CrNiSi20-12 after recycling
Mehanske lastnosti avstenitnega nerjavnega jekla X15CrNiSi20-12 po recikliranju
A. Deli}, O. Kablar, A. Zuli}, D. Kova~evi}, N. Hod`i}, E. Bar~i}. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 153
LETNO KAZALO – INDEX
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 1009
Selection of the most appropriate welding technology for hardfacing of bucket teeth
Izbira najbolj primerne tehnologije trdega navarjanja zoba zajemalke
V. Lazi}, A. Sedmak, R. R. Nikoli}, M. Mutavd`i}, S. Aleksandrovi}, B. Krsti}, D. Milosavljevi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 165
Pitting corrosion of TiN-coated stainless steel in 3 % NaCl solution
Jami~asta korozija nerjavnega jekla s prevleko TiN v 3-odstotni raztopini NaCl
I. Kucuk, C. Sarioglu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 183
Influence of the strain rate on the PLC effect and acoustic emission in single crystals of the CuZn30 alloy compressed at an
elevated temperature
Vpliv hitrosti deformacije na pojav PLC in akusti~no emisijo monokristalov zlitine CuZn30, stiskane pri povi{ani temperaturi
W. Ozgowicz, B. Grzegorczyk, A. Pawe³ek, A. Pi¹tkowski, Z. Ranachowski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 197
Determination of elastic-plastic properties of Alporas foam at the cell-wall level using microscale-cantilever bending tests
Dolo~anje elasti~nih in plasti~nih lastnosti pene Alporas na ravni celi~ne stene z upogibnimi preizkusi z mikroskopsko iglo
T. Doktor, D. Kytýø, P. Koudelka, P. Zlámal, T. Fíla, O. Jirou{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 203
Electrochemical behavior of biocompatible alloys
Elektrokemijsko vedenje biokompatibilnih zlitin
I. Petrá{ová, M. Losertová. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 207
Preparation and thermal stability of ultra-fine and nano-grained commercially pure titanium wires using CONFORM
equipment
Priprava komercialne ultradrobne in nanozrnate Ti-`ice z opremo CONFORM in njena termi~na stabilnost
T. Kubina, J. Dlouhý, M. Köver, M. Dománková, J. Hodek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 213
Estimation of the thermal contact conductance from unsteady temperature measurements
Dolo~anje kontaktne toplotne prevodnosti iz neravnote`nega merjenja temperature
J. Kvapil, M. Pohanka, J. Horský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 219
Potentiodynamic and XPS studies of X10CrNi18-8 steel after ethylene oxide sterilization
Potenciodinami~ne in XPS analize jekla X10CrNi18-8 po sterilizaciji z etilen oksidom
W. Walke, J. Przondziono . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 223
Experimental analysis and modeling of the buckling of a loaded honeycomb sandwich composite
Eksperimentalna analiza in modeliranje upogibanja obremenjenega satastega sendvi~nega kompozita
A. Bentouhami, B. Keskes. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 235
Fatigue behaviour of X70 steel in crude oil
Vedenje jekla X70 pri utrujanju v surovi nafti
¼. Gajdo{, M. [perl, J. Bystrianský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 243
Use of Larson-Miller parameter for modeling the progress of isothermal solidification during transient-liquid-phase bonding
of IN718 superalloy
Uporaba Larson-Millerjevega parametra za modeliranje izotermnega strjevanja pri spajanju z vmesno teko~o fazo superzlitine IN718
M. Pouranvari, S. M. Mousavizadeh . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 247
Effect of severe air-blast shot peening on the wear characteristics of CP titanium
Vpliv intenzivnega povr{inskega kovanja s peskanjem z zrakom na obrabne lastnosti CP-titana
A. C. Karaoglanli . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 253
Finite-element minimization of the welding distortion of dissimilar joints of carbon steel and stainless steel
Uporaba kon~nih elementov za zmanj{anje popa~enja oblike pri varjenju ogljikovega in nerjavnega jekla
E. Ranjbarnodeh, M. Pouranvari, M. Farajpour . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 259
Magnesium-alloy die forgings for automotive applications
Izkovki iz magnezijevih zlitin za avtomobilsko industrijo
M. Madaj, M. Greger, V. Karas . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 267
Resistance to electrochemical corrosion of the extruded magnesium alloy AZ80 in NaCl solutions
Odpornost ekstrudirane magnezijeve zlitine AZ80 proti elektrokemijski koroziji v raztopini NaCl
J. Przondziono, E. Hadasik, W. Walke, J. Szala, J. Michalska, J. Wieczorek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 275
Determination of the cause of the formation of transverse internal cracks on a continuously cast slab
Ugotavljanje vzrokov za nastanek notranjih pre~nih razpok v kontinuirno ulitem slabu
Z. Franìk, M. Masarik, J. [míd . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 285
Neutralization of waste filter dust with CO2
Nevtralizacija odpadnega filtrskega prahu s CO2
A. Kra~un, I. An`el, L. Fras Zemlji~, A. Stergar{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 297
The influence of the morphology of iron powder particles on their compaction in an automatic die
Vpliv morfologije delcev `elezovega prahu na njegovo sposobnost za avtomatsko enoosno stiskanje
B. [u{tar{i~, M. Godec, ^. Donik, I. Paulin, S. Glode`, M. [ori, M. Ratej, N. Javornik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 303
LETNO KAZALO – INDEX
1010 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
Wear mechanisms and surface engineering of forming tools
Obrabni mehanizmi in in`eniring povr{ine preoblikovalnih orodij
B. Podgornik, V. Leskov{ek. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 313
Influence of the impact angle and pressure on the spray cooling of vertically moving hot steel surfaces
Vpliv vpadnega kota in tlaka na ohlajanje z brizganjem na vertikalno premikajo~e se vro~e povr{ine jekla
M. Hnízdil, M. Chabi~ovský, M. Raudenský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 333
Techniques of measuring spray-cooling homogeneity
Tehnike merjenja homogenosti hlajenja z brizganjem
M. Chabi~ovský, M. Raudenský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 337
Reynolds differential equation singularity using processes of small straining with lubrication
Reynoldsova diferencialna ena~ba pri procesih majhne deformacije z mazanjem
D. ]ur~ija, F. Vodopivec, I. Mamuzi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 349
Prediction of the catastrophic tool failure in hard turning through acoustic emission
Napovedovanje katastrofi~ne po{kodbe kerami~nih vlo`kov pri stru`enju z akusti~no emisijo
M. ^illiková, B. Mi~ieta, M. Neslu{an, R. ^ep, I. Mrkvica, J. Petrù, T. Zlámal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 355
Deep micro-hole drilling for hadfield steel by electro-discharge machining (EDM)
Vrtanje globokih mikrolukenj v jekla hadfield z metodo elektrorazreza (EDM)
V. Yilmaz, M. Sarýkaya, H. Dilipak . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 377
Surface analysis of electrochromic CuxO films in their colored and bleached states
Povr{inska analiza elektrokromiznih plasti CuxO v njihovih obarvanih in obeljenih stanjih
M. M. Ristova, M. Milun, B. Pejova . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 387
Thermodynamic analysis of the precipitation of carbonitrides in microalloyed steels
Termodinamska analiza izlo~anja karbonitridov v mikrolegiranih jeklih
M. Opiela . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 395
Experimental investigation of the crack-initiation moment of Charpy specimens under impact loading
Eksperimentalna preiskava trenutka iniciacije razpoke pri udarni obremenitvi Charpyjevih vzorcev
V. Kharchenko, E. Kondryakov, A. Panasenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 403
Structural, thermal and magnetic properties of Fe-Co-Ni-B-Si-Nb bulk amorphous alloy
Strukturne, termi~ne in magnetne lastnosti masivne amorfne zlitine Fe-Co-Ni-B-Si-Nb
S. Lesz, M. Nabia³ek, R. Nowosielski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 409
The nano-wetting aspect at the liquid-metal/SiC interface
Vidik nanoomakanja na stiku staljena kovina-SiC
M. Mihailovi}, K. Rai}, A. Patari}, T. Volkov - Husovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 413
Properties and structure of Cu-Ti-Zr-Ni amorphous powders prepared by mechanical alloying
Lastnosti in struktura amorfnih prahov Cu-Ti-Zr-Ni, pripravljenih z mehanskim legiranjem
A. Guwer, R. Nowosielski, A. Lebuda . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 423
Interaction of Cr2N and Cr2N/Ag thin films with CuZn-brass counterpart during ball-on-disc testing
Interakcija Cr2N in Cr2N/Ag tankih plasti v paru s CuZn-medenino med preizkusom krogla na disk
P. Bílek, P. Jur~i, P. Dulová, M. Hudáková, J. Pta~inová, M. Pa{ák. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 429
Surface behavior of AISI 4140 modified with the pulsed-plasma technique
Lastnosti povr{ine AISI 4140, spremenjene s tehniko pulzirajo~e plazme
Y. Y. Özbek, M. Durman . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 441
Hybrid sol-gel coatings doped with cerium to protect magnesium alloys from corrosion
Hibridni sol-gel-nanosi, dopirani s cerijem, za korozijsko za{~ito magnezijevih zlitin
N. V. Murillo-Gutiérrez, F. Ansart, J.-P. Bonino, M.-J. Menu, M. Gressier . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 453
Influence of the tool geometry and process parameters on the static strength and hardness of friction-stir spot-welded
aluminium-alloy sheets
Vpliv geometrije orodja in parametrov procesa na stati~no trdnost in trdoto pri vrtilno-tornem to~kastem varjenju plo~evin iz Al-zlitine
H. Güler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 457
Impact-toughness investigations of duplex stainless steels
Preiskave udarne `ilavosti dupleksnega nerjavnega jekla
S. Topolska, J. £abanowski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 481
Effects of the artificial-aging temperature and time on the mechanical properties and springback behavior of AA6061
Vpliv temperature in ~asa umetnega staranja na mehanske lastnosti in vzmetnost AA6061
A. Polat, M. Avsar, F. Ozturk . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 487
Influence of the type and number of prepreg layers on the flexural strength and fatigue life of honeycomb sandwich structures
Vpliv vrste in {tevila plasti na utrditev upogibne trdnosti in zdr`ljivosti pri utrujanju satastih sendvi~nih konstrukcij
L. Fojtl, S. Rusnakova, M. Zaludek, V. Rusnák . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 515
LETNO KAZALO – INDEX
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 1011
Potential for obtaining an ultrafine microstructure of low-carbon steel using accumulative roll bonding
Mo`nosti doseganja ultradrobnozrnate mikrostrukture pri spajanju malooglji~nega jekla z akumulativnim valjanjem
T. Kubina, J. Gubi{ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 521
Optimization of the annealing of plaster moulds for the manufacture of metallic foams with an irregular cell structure
Optimiranje postopka `arjenja mav~nih form za izdelavo kovinskih pen z nepravilno strukturo celic
I. Kroupová, P. Lichý, F. Radkovský, J. Beòo, V. Bednáøová, I. Lána . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 527
The kinetics of small-impurity grain-boundary-segregation formation in cold-rolled deep-drawing 08C-Al and IF steels during
post-deformation annealing
Kinetika nastanka segregacije ne~isto~ po mejah zrn med `arjenjem po hladnem valjanju jekel 08C-Al in IF za globoki vlek
A. Rashkovskiy, A. Kovalev, D. Wainstein, I. Rodionova, Y. Bykova, D. Zakharova . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 531
Application of a control-measuring apparatus and peltier modules in the bulk-metallic-glass production using the
pressure-casting method
Uporaba kontrolno-merilne naprave in peltierjevih modulov pri izdelavi masivnih kovinskih stekel po postopku tla~nega litja
W. Pilarczyk, A. Pilarczyk. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 537
Evaluation of the structural changes in 9 % Cr creep-resistant steel using an electrochemical technique
Ocena sprememb strukture v jeklu z 9 % Cr, odpornem proti lezenju, z uporabo elektrokemijske tehnike
J. Rapouch, J. Bystriansky, V. Sefl, M. Svobodova. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 543
Microstructure, magnetic and mechanical properties of the bulk amorphous alloy Fe61Co10Ti4Y5B20
Mikrostruktura, magnetne in mehanske lastnosti masivne amorfne zlitine Fe61Co10Ti4Y5B20
K. Bloch, M. Nabia³ek, J. Gondro . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 553
DP780 dual-phase-steel spot welds: critical fusion-zone size ensuring the pull-out failure mode
To~kasti zvari jekla DP780 z dvofazno strukturo: kriti~na velikost staljene cone, ki zagotavlja poru{enje z izpuljenjem
M. Pouranvari, S. P. H. Marashi, H. L. Jaber . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 579
Properties of polymer-filled aluminium foams under moderate strain-rate loading conditions
Lastnosti aluminijevih pen, napolnjenih s polimernimi materiali, pri zmernih obremenitvah
T. Doktor, P. Zlámal, T. Fíla, P. Koudelka, D. Kytýø, O. Jirou{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 597
Carbide morphology and ferrite grain size after accelerated carbide spheroidisation and refinement (ASR) of C45 steel
Morfologija karbidov in velikost feritnih zrn po pospe{eni sferoidizaciji in rafinaciji (ASR) jekla C45
J. Dlouhy, D. Hauserova, Z. Novy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 625
Investigation on new creep- and oxidation-resistant materials
Preiskava novega materiala, odpornega proti lezenju in oksidaciji
O. Khalaj, B. Masek, H. Jirkova, A. Ronesova, J. Svoboda . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 645
Effect of preheating on mechanical properties in induction sintering of metal-powder material Fe and w(Cu) = 3 %
Vpliv predgrevanja na mehanske lastnosti indukcijsko sintranega materiala, izdelanega iz kovinskega prahu Fe in w(Cu) = 3 %
G. Akpýnar, E. Atik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 653
Kinetic study and characterization of borided AISI 4140 steel
[tudij kinetike in karakterizacija boriranega jekla AISI 4140
M. Keddam, M. Ortiz-Domínguez, O. A. Gómez-Vargas, A. Arenas-Flores, M. Á. Flores-Rentería, M. Elias-Espinosa,
A. García-Barrientos . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 665
Kinetics of the precipitation in austenite HSLA steels
Kinetika izlo~anja v avstenitnih HSLA-jeklih
E. Kalinowska-Ozgowicz, R. Kuziak, W. Ozgowicz, K. Lenik . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 673
Combined influence of V and Cr on the AlSi10MgMn alloy with a high Fe level
Vzajemni vpliv V in Cr na zlitino AlSi10MgMn z visoko vsebnostjo Fe
D. Bolibruchová, M. @ihalová . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 681
Effects of different stirrer-pin forms on the joining quality obtained with friction-stir welding
Vpliv razli~nih oblik vrtilnih konic na kvaliteto spoja pri tornem vrtilnem varjenju
H. Basak, K. Kaptan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 693
Microstructure evolution of advanced high-strength TRIP-aided bainitic steel
Razvoj mikrostrukture naprednega visokotrdnostnega bainitnega jekla z uporabo TRIP
A. Grajcar. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 715
Effect of electric current on the production of NiTi intermetallics via electric-current-activated sintering
Vpliv elektri~nega toka pri izdelavi intermetalne zlitine NiTi s sintranjem, aktiviranim z elektri~nim tokom
T. Yener, S. Siddique, F. Walther, S. Zeytin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 721
Control of soft reduction of continuous slab casting with a thermal model
Kontrola mehke redukcije pri kontinuirnem litju slabov s termi~nim modelom
J. Stetina, P. Ramik, J. Katolicky . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 725
LETNO KAZALO – INDEX
1012 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
Production of shaped semi-products from AHS steels by internal pressure
Izdelava polproizvodov iz AHS-jekel, oblikovanih z notranjim tlakom
I. Vorel, H. Jirková, B. Ma{ek, P. Kurka . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 739
Diffusion kinetics and characterization of borided AISI H10 steel
Kinetika difuzije in karakterizacija boriranega jekla AISI H10
I. Gunes, M. Ozcatal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 759
Application of the Taguchi method to optimize the cutting conditions in hard turning of a ring bore
Uporaba Taguchi-jeve metode za optimizacijo trdega stru`enja roba izvrtine
M. Boy, I. Ciftci, M. Gunay, F. Ozhan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 765
Thermal fatigue of single-crystal superalloys: experiments, crack-initiation and crack-propagation criteria
Toplotno utrujanje monokristalnih superzlitin: preizkusi, merila za nastanek in napredovanje razpoke
L. Getsov, A. Semenov, S. Semenov, A. Rybnikov, E. Tikhomirova . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 773
Investigating the effects of cutting parameters on the built-up-layer and built-up-edge formation during the machining of AISI
310 austenitic stainless steels
Preiskava vplivov parametrov rezanja na nastanek nakopi~ene plasti in nakopi~enega roba med stru`enjem avstenitnega nerjavnega
jekla AISI 310
M. B. Bilgin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 779
Determination of the critical fraction of solid during the solidification of a PM-cast aluminium alloy
Dolo~anje kriti~nega dele`a strjene faze med strjevanjem v PM ulite aluminijeve zlitine
R. Kayikci, M. Colak, S. Sirin, E. Kocaman, N. Akar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 797
Microstructural evolution of Inconel 625 during hot rolling
Mikrostrukturni razvoj Inconela 625 med vro~im valjanjem
F. Tehovnik, J. Burja, B. Podgornik, M. Godec, F. Vode. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 801
Thermophysical properties and microstructure of magnesium alloys of the Mg-Al type
Termofizikalne lastnosti in mikrostruktura magnezijevih zlitin tipa Mg-Al
P. Lichý, J. Beòo, I. Kroupová, I. Vasková . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 807
Effect of process parameters on the microstructure and mechanical properties of friction-welded joints of AISI 1040/AISI
304L steels
Vpliv procesnih parametrov na mikrostrukturo in mehanske lastnosti torno varjenih spojev jekel AISI 1040/AISI 304L
Ý. Kirik, N. Özdemýr . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 825
Influence of inoculation methods and the amount of an added inoculant on the mechanical properties of ductile iron
Vpliv metod modifikacije in koli~ine dodanega modifikatorja na mehanske lastnosti duktilnega `eleza
H. Avdu{inovi}, A. Gigovi}-Geki}, D. ]ubela, R. Sunulahpa{i}, N. Mujezinovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 833
Modeling of occurrence of surface defects of C45 steel with genetic programming
Modeliranje pojava povr{inskih napak pri jeklu C45 z genetskim programiranjem
M. Kova~i~, R. Jager . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 857
Effects of various helically angled grinding wheels on the surface roughness and roundness in grinding cylindrical surfaces
Vpliv razli~nih kotov vija~nice pri brusilnih kolutih na hrapavost povr{ine in okroglost pri bru{enju valjastih povr{in
M. Gavas, M. Kýna, U. Köklü . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 865
Characterization of cast-iron gradient castings
Karakterizacija lito`eleznega gradientnega ulitka
D. Mitrovi}, P. Mrvar, M. Petri~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 871
Comparative mechanical and corrosion studies on magnesium, zinc and iron alloys as biodegradable metals
Primerjalna {tudija mehanskih in korozijskih lastnosti biorazgradljivih zlitin magnezija, cinka in `eleza
D. Vojtìch, J. Kubásek, J. ^apek, I. Pospí{ilová. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 877
Effect of thermomechanical treatment on the corrosion behaviour of Si- and Al-containing high-Mn austenitic steel with Nb
and Ti micro-additions
Vpliv termomehanske obdelave na korozijsko vedenje manganskega avstenitnega jekla z vsebnostjo Si in Al, mikrolegiranega z Nb in Ti
A. Grajcar, A. Plachciñska, S. Topolska, M. Kciuk . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 889
Study on the magnetization-reversal behavior of annealed Sm-Fe-Co-Si-Cu ribbons
[tudij vedenja pri obratu magnetizacije `arjenih trakov Sm-Fe-Co-Si-Cu
M. Doœpial, S. Garus, M. Nabialek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 919
Evaluation of the chitosan-coating effectiveness on a dental titanium alloy in terms of microbial and fibroblastic attachment
and the effect of aging
Ocena u~inkovitosti nanosa hitozana na oprijemanje mikrobov in fibroblastov na dentalni titanovi zlitini ter na pojav staranja
U. T. Kalyoncuoglu, B. Yilmaz, S. Gungor, Z. Evis, P. Uyar, G. Akca, G. Kansu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 925
Structure and properties of the carburised surface layer on 35CrSiMn5-5-4 steel after nanostructurization treatment
Struktura in lastnosti naoglji~ene povr{ine jekla 35CrSiMn5-5-4 po nanostrukturni obdelavi
E. Sko³ek, K. Wasiak, W. A. Œwi¹tnicki. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 933
LETNO KAZALO – INDEX
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 1013
Optimization of the surface roughness by applying the Taguchi technique for the turning of stainless steel under cooling
conditions
Uporaba Taguchi-jeve metode za optimiranje hrapavosti povr{ine pri stru`enju nerjavnega jekla z ohlajanjem
M. Sarýkaya . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 941
Effect of cryogenic treatment applied to M42 HSS drills on the machinability of Ti-6Al-4V alloy
Vpliv podhlajevanja svedrov M42 HSS na obdelovalnost zlitine Ti-6Al-4V
T. Kývak, U. ªeker . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 949
Stability of close-cell Al foams depending on the usage of different foaming agents
Stabilnost aluminijevih pen z zaprto poroznostjo glede na uporabo razli~nih penilnih sredstev
I. Paulin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 983
Evolution of the microstructure and magnetic properties of a cobalt-silicon-based alloy in the early stages of mechanical
milling
Razvoj mikrostrukture in magnetnih lastnosti zlitine Co-Si v za~etnem stadiju mehanskega legiranja
W. Rattanasakulthong, C. Sirisathitkul, P. F. Rogl . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 989
Anorganski materiali – Inorganic materials
Water-soluble cores – verifying development trends
Jedra, topna v vodi – preverjanje smeri razvoja
E. Adámková, P. Jelínek, J. Beòo, F. Mik{ovský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 61
Hydroxyapatite coatings on Cp-Titanium Grade-2 surfaces prepared with plasma spraying
Nanos hidroksiapatita na povr{ino Cp-Titana Grade-2 z nabrizgavanjem s plazmo
R. Rudolf, D. Stamenkovi}, Z. Aleksi}, M. Jenko, I. \ordevi}, A. Todorovi}, V. Jokanovi}, K. T. Rai} . . . . . . . . . . . . . . . . . . . . . . . . . 81
Dry-cutting options with a chainsaw at the Hotavlje I natural-stone quarry
Mo`nosti suhega rezanja z veri`no `ago v kamnolomu naravnega kamna Hotavlje I.
J. Kortnik, B. Markoli . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103
Thin tin monosulfide films deposited with the HVE method for photovoltaic applications
Tanka plast HVE kositrovega monosulfida za uporabo v fotovoltaiki
N. Revathi, S. Bereznev, J. Lehner, R. Traksmaa, M. Safonova, E. Mellikov, O. Volobujeva . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 149
Comparison of refractory coatings based on talc, cordierite, zircon and mullite fillers for lost-foam casting
Primerjava ognjevzdr`nih premazov na osnovi smukca, kordierita, cirkona in mulitnih polnil za ulivanje v forme z izparljivim modelom
Z. A}imovi}-Pavlovi}, A. Terzi}, Lj. Andri}, M. Pavlovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 157
TEM replica of a fluoride-miserite glass-ceramic glaze microstructure
TEM-replike mikrostrukture steklokerami~ne fluor-mizeritne glazure
J. Ma. Rincón, R. Casasola . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 229
Microwave-assisted hydrothermal synthesis of Ag/ZnO sub-microparticles
Hidrotermi~na sinteza podmikrometrskih delcev Ag/ZnO z mikrovalovi
L. Münster, P. Ba`ant, M. Machovský, I. Kuøitka. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 281
Mineralogical and geochemical characterization of roman slag from the archaeological site near Mo{nje (Slovenia)
Mineralo{ka in geokemi~na karakterizacija rimske `lindre z arheolo{kega najdi{~a pri Mo{njah (Slovenija)
S. Kramar, J. Lux, H. Pristacz, B. Mirti~, N. Rogan - [muc . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 343
Erosive wear resistance of silicon carbide-cordierite ceramics: influence of the cordierite content
Odpornost keramike silicijev karbid-kordierit proti obrabi pri eroziji: vpliv vsebnosti kordierita
M. Po{arac-Markovi}, Dj. Veljovi}, A. Deve~erski, B. Matovi}, T. Volkov-Husovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 365
Influence of the substrate temperature on the structural, optical and thermoelectric properties of sprayed V2O5 thin films
Vpliv temperature podlage na strukturne, opti~ne in termoelektri~ne lastnosti napr{ene tanke plasti V2O5
Y. Vijayakumar, K. N. Reddy, A. V. Moholkar, M. V. R. Reddy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 371
Preparation and dielectric properties of thermo-vaporous BaTiO3 ceramics
Priprava in dielektri~ne lastnosti termo-parno porozne keramike BaTiO3
A. Kholodkova, M. Danchevskaya, N. Popova, L. Pavlyukova, A. Fionov . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 447
Spectroscopic and porosimetric analyses of Roman pottery from an archaeological site near Mo{nje, Slovenia
Spektroskopske in porozimetri~ne preiskave rimske lon~enine z arheolo{kega najdi{~a pri Mo{njah, Slovenija
S. Kramar, J. Lux . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 503
Effect of the by-pass cement-kiln dust and fluidized-bed-combustion fly ash on the properties of fine-grained alkali-activated
slag-based composites
Vpliv prahu iz pe~i za cement in lete~ega pepela iz vrtin~aste plasti na lastnosti drobnozrnatega, z alkalijami aktiviranega kompozita na
osnovi `lindre
V. Bílek Jr., L. Paøízek, L. Kalina . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 549
LETNO KAZALO – INDEX
1014 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
Preparation of porous ceramic materials based on CaZrO3
Priprava porozne keramike na osnovi CaZrO3
E. Œnie¿ek, J. Szczerba, I. Jastrzêbska, E. Kleczyk, Z. Pêdzich . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 573
Quality of the structure of ash bodies based on different types of ash
Kvaliteta strukture telesa iz pepela na osnovi razli~nih vrst pepela
V. Cerny . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 601
Sintered board materials based on recycled glass
Sintrani plo{~ati materiali na osnovi recikliranega stekla
T. Melichar, J. Byd`ovský . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 607
Development of composite salt cores for foundry applications
Razvoj kompozitnih slanih jeder za uporabo v livarstvu
J. Beòo, E. Adámková, F. Mik{ovský, P. Jelínek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 619
Use of micromachining to shape the structure and electrical properties of the front electrode of a silicon solar cell
Uporaba mikroobdelovanja za oblikovanje strukture in elektri~nih lastnosti prednje elektrode silicijeve son~ne celice
M. Musztyfaga-Staszuk . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 629
Assessment of the impact-echo method for monitoring the long-standing frost resistance of ceramic tiles
Ocena metode impact-echo za kontrolo dolgotrajne odpornosti kerami~nih plo{~ic proti zmrzali
M. Matysik, I. Plskova, Z. Chobola . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 639
A reliable approach to a rapid calculation of the grain size of polycrystalline thin films after excimer laser crystallization
Zanesljiv na~in hitrega izra~una velikosti zrn v polikristalni tanki plasti po UV-laserski kristalizaciji
C. C. Kuo . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 687
Effect of the aggregate type on the properties of alkali-activated slag subjected to high temperatures
Vpliv vrste agregata na lastnosti z alkalijo aktivirane `lindre, izpostavljene visokim temperaturam
P. Rovnaník, Á. Dufka. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 709
Optimization of operating conditions in a laboratory SOFC testing device
Optimizacija obratovalnih razmer laboratorijske gorivne celice SOFC
T. Skalar, M. Lubej, M. Marin{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 731
Friction and wear behaviour of ulexite and cashew in automotive brake pads
Odpornost proti trenju in obrabi avtomobilskih zavornih oblog z uleksitom in prahom iz indijskega oreha
Ý. Sugözü, Ý. Mutlu, A. Keskin. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 751
Elastic behaviour of magnesia-chrome refractories at elevated temperatures
Elasti~no vedenje ognjevzdr`nih gradiv magnezija-krom pri povi{anih temperaturah
I. Jastrzêbska, J. Szczerba, J. Szlêzak, E. Œnie¿ek, Z. Pêdzich . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 913
Solid-state sintering of (K0.5Na0.5)NbO3 synthesized from an alkali-carbonate-based low-temperature calcined powder
Sintranje v trdnem keramike (K0,5Na0,5)NbO3, sintetizirane iz nizkotemperaturno kalciniranega prahu, pripravljenega na osnovi alkalijskih
karbonatov
M. Feizpour, T. Ebadzadeh, D. Jenko. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 975
Polimeri – Polymers
Antibacterial composite based on nanostructured ZnO mesoscale particles and a poly(vinyl chloride) matrix
Protibakterijski kompozit na osnovi nanostrukturnih delcev ZnO in osnove iz polivinil klorida
J. Sedlák, P. Ba`ant, J. Klofá~, M. Pastorek, I. Kuøitka. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55
Micromechanical model of the substituents of a unidirectional fiber-reinforced composite and its response to the tensile cyclic
loading
Mikromehanski model nadomestkov kompozitov, oja~anih z enosmernimi vlakni, in njihov odgovor na cikli~no natezno obremenjevanje
T. Kroupa, H. Srbová, R. Zem~ík . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 99
New solid-polymer-electrolyte material for dye-sensitized solar cells
Novi elektrolitni material na osnovi trdnega polimera za son~ne celice, ob~utljive za svetlobo
V. K. Singh, B. Bhattacharya, S. Shukla, P. K. Singh . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 123
Design of a wideband planar antenna on an epoxy-resin-reinforced woven-glass material
[irokopasovna ploskovna antena na epoksi smoli, oja~ani s steklenimi vlakni
R. Azim, M. T. Islam . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 193
Predicting the physical properties of drawn Nylon-6 fibers using an artificial-neural-network model
Napovedovanje fizikalnih lastnosti vle~enih vlaken iz najlona 6 z uporabo modela umetne nevronske mre`e
R. Semnani Rahbar, M. Vadood. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 325
Monitoring of polyurethane dispersions after the synthesis
Spremljanje poliuretanskih disperzij po sintezi
M. Ocepek, J. Zabret, J. Kecelj, P. Venturini, J. Golob . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 495
LETNO KAZALO – INDEX
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 1015
Composite-material printed antenna for a multi-standard wireless application
Tiskana antena iz kompozitnega materiala za ve~standardno brez`i~no uporabo
T. Alam, M. R. I. Faruque, M. T. Islam, N. Misran . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 745
Flame resistance and mechanical properties of composites based on new advanced resin system FR4/12
Negorljivost in mehanske lastnosti kompozitov na osnovi novih naprednih sistemov smol FR4/12
V. Rusnák, S. Rusnáková, L. Fojtl, M. @aludek, A. ^apka . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 821
Load-capacity prediction for the carbon- or glass-fibre-reinforced plastic part of a wrapped pin joint
Napoved nosilnosti plasti~nih delov zati~nega spoja, oja~anega z ogljikovimi ali steklastimi vlakni
J. Krystek, R. Kottner . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 957
Nanomateriali in nanotehnologije – Nanomaterials and nanotechnology
Synthesis of NiTi/Ni-TiO2 composite nanoparticles via ultrasonic spray pyrolysis
Sinteza kompozitnih nanodelcev NiTi/Ni-TiO2 z ultrazvo~no razpr{ilno pirolizo
P. Majeri~, R. Rudolf, I. An`el, J. Bogovi}, S. Stopi}, B. Friedrich . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75
Heat-transfer characteristics of a non-Newtonian Au nanofluid in a cubical enclosure with differentially heated side walls
Zna~ilnosti prenosa toplote nenewtonske Au nanoteko~ine v kockastem ohi{ju z razli~no gretima stranskima stenama
P. Ternik, R. Rudolf, Z. @uni~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87
Characterization of TiO2 nanoparticles with high-resolution FEG scanning electron microscopy
Karakterizacija nanodelcev TiO2 z visokolo~ljivostno FEG vrsti~no elektronsko mikroskopijo
Z. Samard`ija, D. Lapornik, K. Gradi{ek, D. Verhov{ek, M. ^eh . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 173
The stabilization of nano silver on polyester filament for a machine-made carpet
Stabilizacija nanodelcev srebra na poliestrskem vlaknu za strojno izdelavo preprog
K. M. Shojaei, A. Farrahi, H. Farrahi, A. Farrahi . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 461
Mechanical and wetting properties of nanosilica/epoxy-coated stainless steel
Mehanske in povr{inske lastnosti premaza iz silicijevih nanodelcev in epoksidne smole na nerjavnem jeklu
M. Conradi, G. Intihar, M. Zorko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 613
Fabrication of TiO2 nanotubes for bioapplications
Izdelava TiO2-nanocevk za biomedicinsko uporabo
M. Kulkarni, K. Mrak-Polj{ak, A. Fla{ker, A. Mazare, P. Schmuki, A. Kos, S. ^u~nik, S. Sodin-[emrl, A. Igli~. . . . . . . . . . . . . . . . . . . 635
Au-nanoparticle synthesis via ultrasonic spray pyrolysis with a separate evaporation zone
Sinteza nanodelcev zlata z ultrazvo~no razpr{ilno pirolizo z lo~eno cono izhlapevanja
P. Majeri~, B. Friedrich, R. Rudolf . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 791
One-step green synthesis of graphene/ZnO nanocomposites for non-enzymatic hydrogen peroxide sensing
Enostopenjska zelena sinteza nanokompozita grafen-ZnO za neencimatsko detekcijo vodikovega peroksida
S. S. Low, M. T. T. Tan, P. S. Khiew, H. S. Loh, W. S. Chiu . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 837
Gradbeni materiali – Materials in civil engineering
The effect of a superplasticizer admixture on the mechanical fracture parameters of concrete
Vpliv dodatka superplastifikatorja na parametre mehanskega zloma betona
H. [imonová, I. Havlíková, P. Danìk, Z. Ker{ner, T. Vymazal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 417
Evaluation of the thermal resistance of selected bentonite binders
Ocena toplotne upornosti izbranih bentonitnih veziv
J. Beòo, J. Vontorová, V. Matìjka, K. Gál . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 465
Modified cement-based mortars: crack initiation and volume changes
Modificirane malte na osnovi cementa: iniciacija razpok in volumenske spremembe
I. Havlikova, V. Bilek Jr., L. Topolar, H. Simonova, P. Schmid, Z. Kersner. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 557
Fracture properties of plain and steel-polypropylene-fiber-reinforced high-performance concrete
Lastnosti loma navadnega in visokozmogljivega betona, oja~anega s polipropilenskimi vlakni
P. Smarzewski, D. Barnat-Hunek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 563
Application of computed tomography in comparison with the standardized methods for determining the permeability of
cement-composite structures
Uporaba ra~unalni{ke tomografije v primerjavi s standardiziranimi metodami dolo~anja prepustnosti cementnih kompozitnih struktur
T. Komárková, M. Králíková, P. Kovács, D. Kocáb, T. Stavaø . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 587
Monitoring early-age concrete with the acoustic-emission method and determining the change in the electrical properties
Pregled sve`ega betona z metodo akusti~ne emisije in dolo~anjem sprememb elektri~nih lastnosti
L. Pazdera, L. Topolar, M. Korenska, T. Vymazal, J. Smutny, V. Bilek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 703
LETNO KAZALO – INDEX
1016 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018
Mortar-type identification for the purpose of reconstructing fragmented Roman wall paintings (Celje, Slovenia)
Analiza ometov za rekonstrukcijo fragmentov rimskih stenskih poslikav (Celje, Slovenija)
M. Gutman, M. Lesar Kikelj, J. Kuret, S. Kramar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 785
Micro-encapsulated phase-change materials for latent-heat storage: thermal characteristics
Mikroenkapsulirani materiali s fazno premeno za shranjevanje latentne toplote: toplotne zna~ilnosti
M. Ostrý, D. Dostálová, T. Klubal, R. Pøikryl, P. Charvát. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 813
Ceramic masonry units intended for the masonry resistant to high humidity
Kerami~ni gradbeni elementi, namenjeni za zgradbe, odporne proti visoki vlagi
J. Zach, V. Novák, J. Hroudová, M. Sedlmajer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 817
Problems and normative evaluation of bond-strength tests for coated reinforcement and concrete
Problemi in normativna ocena preizkusov trdnosti vezi med armaturo s prekritjem in betonom
P. Pokorný, M. Kouøil, J. Stoulil, P. Bou{ka, P. Simon, P. Juránek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 847
Microstructural changes of fine-grained concrete exposed to a sulfate attack
Mikrostrukturne spremembe drobnozrnatega betona, izpostavljenega sulfatu
M. Vy{vaøil, P. Bayer, M. Rovnaníková . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 883
Surface free energy of hydrophobic coatings of hybrid-fiber-reinforced high-performance concrete
Prosta energija povr{ine hidrofobnih premazov na visokozmogljivem betonu, oja~anem s hibridnimi vlakni
D. Barnat-Hunek, P. Smarzewski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 895
Numeri~ne metode – Numerical methods
Mathematical modelling and physical simulation of the hot plastic deformation and recrystallization of steel with
micro-additives
Matemati~no modeliranje in fizikalna simulacija vro~e plasti~ne predelave in rekristalizacije jekla z mikrododatki
E. Kalinowska-Ozgowicz, W. Wajda, W. Ozgowicz . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 69
Effective preparation of non-linear material models using a programmed optimization script for a nurimerical simulation of
sheet-metal processing
U~inkovita priprava nelinearnih modelov materiala s programiranim optimizacijskim zapisom za numeri~no simulacijo obdelave plo~evine
M. Urbánek, F. Tikal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 291
Using simulated spectra to test the efficiency of spectral processing software in reducing the noise in Auger electron spectra
Uporaba simuliranega spektra za preizkus u~inkovitosti programske opreme predelave spektra pri zmanj{anju {uma spektra Augerjevih
elektronov
B. Poniku, I. Beli~, M. Jenko . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 435
Development of numerical models for the heat-treatment-process optimisation in a closed-die forging production
Razvoj numeri~nih modelov za optimizacijo postopka toplotne obdelave pri proizvodnji odkovkov v zaprtih utopnih orodjih
L. Male~ek, M. Fedorko, F. Van~ura, H. Jirková, B. Ma{ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 471
Non-linear finite-element simulations of the tensile tests of textile composites
Nelinearna simulacija nateznih preizkusov tekstilnih kompozitov s kon~nimi elementi
T. Kroupa, K. Kunc, R. Zem~ík, T. Mandys. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 509
Material parameters for a numerical simulation of a compaction process for sintered double-height gears
Materialni parametri za numeri~no simulacijo postopka stiskanja sintranih dvovi{inskih zobnikov
T. Verlak, M. [ori, S. Glode` . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 841
Developing continuous-casting-process control based on advanced mathematical modelling
Uporaba naprednega matemati~nega modeliranja za razvoj kontrole postopka kontinuirnega ulivanja
J. Falkus, K. Mi³kowska-Piszczek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 903
A meshless model of electromagnetic braking for the continuous casting of steel
Brezmre`ni model elektromagnetnega zaviranja pri kontinuiranem ulivanju jekla
K. Mramor, R. Vertnik, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 961
Non-singular method of fundamental solutions for three–dimensional isotropic elasticity problems with displacement boundary
conditions
Nesingularna metoda fundamentalnih re{itev za deformacijo tridimenzijskih elasti~nih problemov z deformacijskimi robnimi pogoji
Q. Liu, B. [arler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 969
Editor's Preface – Predgovor urednika
Editor’s Preface / Predgovor urednika
M. Torkar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 181
In Memoriam
Alojz Pre{ern, dipl. in`. metalurgije (1920–2015) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 477
Prof. dr. Milan Trbi`an . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 993
LETNO KAZALO – INDEX
Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018 1017
Letno kazalo – Index
Letnik 49 (2015), 1–6 – Volume 49 (2015), 1–6 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 995
LETNO KAZALO – INDEX
1018 Materiali in tehnologije / Materials and technology 49 (2015) 6, 995–1018