VSEBINA – CONTENTS
PREGLEDNI ^LANKI – REVIEW ARTICLES
Effect of heat treatment on the microstructure and mechanical properties of martensitic stainless-steel joints welded with
austenitic stainless-steel fillers
Vpliv toplotne obdelave na mikrostrukturo in mehanske lastnosti martenzitnih nerjavnih spojev, varjenih z avstenitnimi nerjavnimi
elektrodami
A. Calik, M. S. Karakaº . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 403
IZVIRNI ZNANSTVENI ^LANKI – ORIGINAL SCIENTIFIC ARTICLES
New combined bio-scouring and bio-bleaching process of cotton fabrics
Nov, zdru`eni postopek bioizkuhavanja in biobeljenja bomba`nih tkanin
N. [pi~ka, P. Forte Tav~er . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 409
Effect of mechanical activation on mullite formation in an alumina-quartz ceramics system
Vpliv mehanske aktivacije na nastanek mullita v kerami~nem sistemu glinica-kremen
E. Elmas, K. Yildiz, N. Toplan, H. Özkan Toplan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 413
Determination of the mechanical parameters of a bonded joint between a metal and a composite by comparing experiments
with a finite-element model
Dolo~anje mehanskih parametrov spoja med kovino in kompozitom s primerjavo preizkusov in modelom kon~nih elementov
P. Bernardin, J. Vacík, T. Kroupa, R. Kottner . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 417
Determination of the metal concentrations in an anode material for solid-oxide fuel cells
Dolo~itev vsebnosti kovin v anodnem materialu za gorivne celice s trdnim elektrolitom
T. Skalar, A. Golobi~, M. Marin{ek, J. Ma~ek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 423
Producing antibacterial silver-doped hydroxyapatite powders with chemical precipitation and reshaping in a spray dryer
Izdelava s srebrom dopiranega protibakterijskega prahu hidroksiapatita s kemijskim izlo~anjem in preoblikovanjem v razpr{ilnem
su{ilniku
F. E. Baþtan, Y. Y. Özbek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 431
Tribocorrosion degradation of protective coatings on stainless steel
Tribokorozijska degradacija za{~itnih prevlek na nerjavnem jeklu
D. Kek Merl, P. Panjan, M. ^ekada, P. Gselman, S. Paskvale . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 435
Behaviour of short cracks emanating from tiny drilled holes
Vedenje kratkih razpok, ki nastanejo iz majhnih izvrtin
V. Gliha, P. Maruschak, T. Vuherer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 441
Internal-oxidation kinetics of Ag-Cd alloys
Kinetika notranje oksidacije zlitin Ag-Cd
L. Balanovi}, V. ]osovi}, N. Talijan, D. @ivkovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 447
Effects of the ageing treatment on the superelastic behavior of a nitinol stent for an application in the esophageal duct:
a finite-element analysis
U~inek staranja na superelasti~no vedenje nitinolne opore pri uporabi v cevi po`iralnika: analiza kon~nih elementov
F. Nematzadeh, S. K. Sadrnezhaad . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 453
Investigation of the effect of skin-pass rolling on the formability of low-carbon steel sheets
Preiskava u~inka dresirnega valjanja na preoblikovalnost malooglji~ne jeklene plo~evine
D. Asefi, H. Monajatizadeh, A. Ansaripour, A. Salimi . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 461
Synthesis, characterization and sensing application of a solid alum/fly ash composite electrolyte
Sinteza, karakterizacija in mo`nost uporabe trdnega kompozitnega elektrolita galun-lete~i pepel
A. Sachdeva, R. Singh, P. K. Singh, B. Bhattacharya . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 467
ISSN 1580-2949
UDK 669+666+678+53 MTAEC9, 47(4)401–540(2013)
MATER.
TEHNOL.
LETNIK
VOLUME 47
[TEV.
NO. 4
STR.
P. 401–540
LJUBLJANA
SLOVENIJA
JULY–AUG.
2013
Different source atelocollagen thin films: preparation, process optimisation and its influence on the interaction with eukaryotic
cells
Razli~en izvir atelokolagenskih tankih plasti: priprava, optimizacija procesa in vpliv na interakcije z evkarionti~nimi celicami
J. López García, P. Humpolí~ek, M. Lehocký, I. Junkar, M. Mozeti~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 473
Evaluation of equilibrium isotherm models for the adsorption of Cu and Ni from wastewater on bentonite clay
Ocena modelov ravnote`nih izoterm za adsorpcijo Cu in Ni iz odpadnih vod na bentonitno glino
S. A. Al-Jlil, M. S. Latif . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 481
Modelling and optimization of a laser droplet-formation process
Modeliranje in optimizacija laserskega tvorjenja kapljice
T. Kokalj, I. Grabec, E. Govekar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 487
STROKOVNI ^LANKI – PROFESSIONAL ARTICLES
Comparison of the temperature fields of continuously cast steel slabs with different chemical compositions
Primerjava temperaturnega polja kontinuirno ulitih jeklenih slabov razli~nih kemijskih sestav
F. Kavicka, K. Stransky, B. Sekanina, J. Stetina, M. Masarik, T. Mauder, V. Gontarev . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 497
Investigation of the thermomechanical properties and microstructure of special magnesium alloys
Preiskava termomehanskih lastnosti in mikrostrukture posebnih magnezijevih zlitin
P. Lichý, M. Cagala, J. Beòo . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 503
Impact of casting speed on the temperature field of continuously cast steel billets
Vpliv livne hitrosti na temperaturno polje kontinuirno ulitih gredic iz jekla
L. Klimes, J. Stetina, P. Bucek . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 507
Development of the production of ultrafine-grained titanium with the conform equipment
Razvoj proizvodnje ultradrobnozrnatega titana z opremo conform
M. Duchek, T. Kubina, J. Hodek, J. Dlouhy. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 515
Conversion of the Ms70 alpha-brass torsion-test data into hot-forming processing maps
Pretvorba podatkov torzijskega preizkusa alfa medi Ms70 v procesne zemljevide vro~e predelave
T. Kubina, J. Boøuta, R. Pernis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 519
Recycled polymer/clay composites for heavy-metals adsorption
Recikliran kompozit polimer-glina za adsorpcijo te`kih kovin
F. D. Alsewailem, S. A. Aljlil . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 525
Influential factors in the surface-hardness testing of a nitrided layer
Vplivni faktorji pri preizku{anju trdote povr{ine nitrirane plasti
F. Cajner, D. Landek, I. Kumi}, A. Vugrin~i} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 531
Mechanical and microstructural properties of the fly-ash-based geopolymer paste and mortar
Mehanske in mikrostrukturne lastnosti geopolimernih veziv in malt iz lete~ega pepela
R. Zejak, I. Nikoli}, D. Ble~i}, V. Radmilovi}, V. Radmilovi} . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 535
A. CALIK, M. S. KARAKAª: EFFECT OF HEAT TREATMENT ON THE MICROSTRUCTURE ...
EFFECT OF HEAT TREATMENT ON THE
MICROSTRUCTURE AND MECHANICAL PROPERTIES
OF MARTENSITIC STAINLESS-STEEL JOINTS WELDED
WITH AUSTENITIC STAINLESS-STEEL FILLERS
VPLIV TOPLOTNE OBDELAVE NA MIKROSTRUKTURO IN
MEHANSKE LASTNOSTI MARTENZITNIH NERJAVNIH SPOJEV,
VARJENIH Z AVSTENITNIMI NERJAVNIMI ELEKTRODAMI
Adnan Calik, Mustafa Serdar Karakaº
Department of Manufacturing Engineering, Faculty of Technology, Suleyman Demirel University, 32260 Cunur, Isparta, Turkey
adnancalik@sdu.edu.tr
Prejem rokopisa – received: 2012-11-08; sprejem za objavo – accepted for publication: 2013-01-04
AISI 422 martensitic stainless steels were welded according to the ISO 15792-1 standard with austenitic stainless-steel filler.
The effects of tempering and preheating heat treatments were also evaluated. The microstructures of the welds were
characterized with optical microscopy. Mechanical properties were determined via microhardness, tensile and fatigue tests, and
compared to those of the unwelded AISI 422 steel. X-ray diffraction was used to characterize the phases present in the weld.
Chromium carbides were observed at the grain boundaries of the welds. The precipitation was caused mainly by the diffusion of
carbon into the austenitic stainless steel during the welding. The microhardness decreased from the base metal to the weld. The
tensile-strength and percent-elongation values were improved with preheating and tempering heat treatments. The highest
fatigue limit was obtained with preheating.
Keywords: welding, heat treatment, stainless steel, fatigue, hardness
Martenzitno nerjavno jeklo AISI 422 je bilo varjeno skladno s standardom ISO 15792-1 z elektrodami iz avstenitnega
nerjavnega jekla. Ocenjen je bil vpliv kaljenja in predogrevanja. Mikrostruktura zvarov je bila opredeljena s svetlobno
mikroskopijo. Mehanske lastnosti so bile dolo~ene z mikrotrdoto, nateznimi in utrujenostnimi preizkusi in primerjane z
nevarjenim jeklom AISI 422. Rentgenska difrakcija je bila uporabljena za karakterizacijo faz v zvaru. Po mejah zrn v zvaru je
bilo opaziti kromove karbide. Ve~ino izlo~anja je povzro~ila difuzija ogljika v avstenitno jeklo med varjenjem. Mikrotrdota se je
zmanj{evala od osnovnega materiala do zvara. Vrednosti natezne trdnosti in dele` raztezka so se pove~ali s predgrevanjem in s
toplotno obdelavo. Najvi{ja odpornost proti utrujanju je bila dose`ena s predgrevanjem.
Klju~ne besede: varjenje, toplotna obdelava, nerjavno jeklo, utrujenost, trdota
1 INTRODUCTION
For metals, fatigue failures occur more often in
welded joints than in unwelded structures. Weld defects
and imperfections affect the fatigue and other mechani-
cal properties of structures.1,2 In general, fatigue strength
decreases significantly with the introduction of any stress
raisers in metals. Machine parts contain stress raisers in
industrial applications. Fatigue cracks in structural parts
almost always initiate at geometrical irregularities.3
Weld defects and imperfections can be accepted as
critical stress raisers in welded structures. Fatigue pro-
perties of welded joints are more complicated because of
the nature of the weld. Slag inclusions, pores, undercuts,
incomplete fusion and residual stresses all affect the
crack initiation and crack-propagation stages of the
fatigue life of a welded joint.4–6 In addition, surface
roughness and section changes may also affect the
existing or newly developing fatigue cracks. The fatigue
behavior of a welded joint is influenced by the thickness
of a plate as well as by the other geometrical para-
meters.7 However, there is conflicting evidence with
regard to the effect of a welding procedure on the fatigue
strength. For the fillet-welded and butt-welded joints the
fatigue strength is strongly dependent on the welding
procedure.8,9
In spite of a poor weldability, martensitic stainless
steels may be welded in the annealed, stress-relieved and
tempered conditions. As an alloying element, Cr has
very significant effects on the metallurgical aspect of the
martensitic stainless steels. Cr and C are specifically
added to steel to ensure the formation of martensite after
the hardening. Ni can be added, and it enhances both the
yield strength and ductility of these steels. The hardness
in this type of material is dependent on the carbon
content.10 The carbon content can also affect the hard-
ness of the heat-affected zone (HAZ). Thus, the influ-
ence of hardness of HAZ may be controlled with
welding characteristics. In general, the welding of the
martensitic stainless steels with a carbon content higher
than 0.25 % is not recommended. In the 0.05–0.15 % C
range, the hardness increases with the increasing carbon
concentration. Austenitic or ferritic-austenitic electrodes
are used for martensitic stainless steels. The hardness
and ductility of a weld and HAZ vary in a welded struc-
ture. To avoid this harmful effect on the welded AISI 420
Materiali in tehnologije / Materials and technology 47 (2013) 4, 403–407 403
UDK 621.791:621.791.053:669.14.018.8 ISSN 1580-2949
Review article/Pregledni ~lanek MTAEC9, 47(4)403(2013)
steel, preheating and a slow cooling rate after the weld-
ing must be considered.11,12
In this study, an attempt has been made to clarify the
effects of pre- and post-weld treatments on the welding
of the martensitic stainless steel with an austenitic stain-
less-steel filler. The present study highlights the effects
of preheating and tempering on the microstructural,
hardness and tensile properties, and on the fatigue
behavior of the welded joints.
2 EXPERIMENTAL PROCEDURE
2.1 Materials
AISI 422 steel in its annealed state was used as the
base metal for the welding experiments. The chemical
composition of the steel is given in Table 1. This steel is
widely used in the transmission systems of diesel loco-
motives. Cylindrical as-cast billets of 80 mm diameters
and 1 m lengths were hot-forged at 800 °C into the plates
of 25 mm thickness.
The forged plates were then machined to the weld
specimen dimensions specified in the ISO 15792-1
standard. The geometry of the weld specimens is shown
in Figure 1a. Type 1.3 specimens were used and their
dimensions are given in Table 2. An OFHC (Oxygen
Free High Conductivity) copper backing strip was used
during the welding. The sequence of the welding passes
is shown in Figure 1b.
Table 2: Dimensions (mm) of the type 1.3 weld test specimen
depicted in Figure 1
Tabela 2: Dimenzije (mm) preizku{anca za varjenje vrste 1.3, prika-
zanega na sliki 1
t a b u l
20 ≥ 150 16 6 10 ≥ 150
The specimens were thoroughly cleaned with acetone
before the welding. The austenitic stainless-steel
electrodes (E310L) of a 3.25 mm diameter were used for
the welding tests. The chemical composition of the
welding electrodes is shown in Table 3. In order to
prevent hydrogen-induced cracking, the electrodes were
dried in a muffle furnace at 573 K for 2 h before the
welding. The welding procedure was then carried out
manually in a single V-butt-joint configuration using a
UMS 250 DC type electric-arc welding machine with a
total of 13 passes. The current was held constant at 170
A. The total of 13 runs were needed to fill the groove
between the butt joints.
Table 3: Chemical composition of the E310L filler electrode in mass
fractions, w/%
Tabela 3: Kemijska sestava varilne elektrode E310L v masnih dele`ih,
w/%
C Mn Si Ni Cr Fe
0.10 2.0 0.4 20 25 Bal.
The effects of preheating and tempering heat treat-
ments were investigated. The welding parameters and
notations used for the as-welded, preheated and tem-
pered specimens are summarized in Table 4. From here
onwards, the codes AW, PH and TP shall be used to
express the as-welded, preheated and tempered speci-
mens, respectively.
Table 4: Welding parameters used for the test specimens
Tabela 4: Parametri varjenja, uporabljeni pri preizkusnih vzorcih
Specimen Filler metal Preheating PWHT
AW E310L N/A N/A
PH E310L 623 K, 2 h N/A
TP E310L N/A 1023 K, 3 h
2.2 Metallography
The microstructures of the welded joints were eva-
luated using optical microscopy. The specimens exa-
mined were prepared using standard metallographic
techniques. The original microstructure of the original
AISI 422 steel (before the welding) is shown in Figure
2. This reveals a fine, fully martensitic microstructure.
A. CALIK, M. S. KARAKAª: EFFECT OF HEAT TREATMENT ON THE MICROSTRUCTURE ...
404 Materiali in tehnologije / Materials and technology 47 (2013) 4, 403–407
Figure 1: a) Shape and dimensions of a weld specimen as indicated in
the ISO 15792-1 standard; b) sequence of the welding passes
Slika 1: a) Oblika in dimenzije varjenih vzorcev, kot je prikazano v
ISO 15792-1 standardu, b) potek zaporedja prehodov v zvaru
Table 1: Chemical composition of the AISI 422 steel in mass fractions, w/%
Tabela 1: Kemijska sestava jekla AISI 422 v masnih dele`ih, w/%
Element C Si S P Mn Ni Cr Mo V Fe
Composition 0.220 0.220 0.023 0.02 0.403 1.268 11.15 0.086 0.017 Bal.
2.3 Mechanical testing
Tensile, fatigue, and microhardness test specimens
were extracted from the welded samples. Tensile tests
were carried out at room temperature in accord with the
ISO 6892-1 standard. Yield strength (YS), ultimate
tensile strength (UTS) and fraction of elongation (% EL)
values were determined. Fatigue tests were done using a
rotating bar bending fatigue testing machine in accord
with the ISO 1143-1 standard. Microhardness measure-
ments were made utilizing a Vickers indenter with a load
of 100 g and dwell time of 15 s. For each of the me-
chanical tests, a minimum of three measurements/tests
were made for each experimental condition.
3 RESULTS AND DISCUSSION
After the welding, the specimens generally consisted
of three distinctive regions: (i) the weld metal; (ii) the
heat-affected zone (HAZ) and (iii) the base metal that
was not affected by the welding process. The micro-
structures of the AW specimen are shown in Figure 3.
The weld metal shows a microstructure consisting main-
ly of austenite (Figure 3a), with delta ferrite present at
the grain boundaries.
The HAZ of the AW specimen revealed both mar-
tensite and austenite phases. The coexistence of the two
phases in the AW specimen is shown in Figure 3b. The
microstructure also shows the presence of precipitates
(possibly inclusions) located at the centers of the
austenite phases. The XRD pattern of the weld in Figure
4 indicates the presence of chromium carbides (Cr23C6,
Cr7C3). High temperatures, combined with the diffusion
of C from the martensitic stainless steel to the austenitic
stainless steel may have caused the formation of these
carbides during the welding.
The microstructures of the PH specimen are shown in
Figure 5. The microstructure of the weld metal reveals
an austenitic structure with possible delta ferrite
segregated at the grain boundaries (Figure 5a). The heat
affected zone (HAZ) between the weld and the base
metal reveals the presence of martensite and austenite
phases (Figure 5b).
The microstructures of the TP specimen are shown in
Figure 6. The microstructure of the weld metal is auste-
nitic, similar to the weld metal observed in the AW and
PH specimens. However, the dendritic morphology of
the austenitic structure is much more visible (Figure 6a).
A. CALIK, M. S. KARAKAª: EFFECT OF HEAT TREATMENT ON THE MICROSTRUCTURE ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 403–407 405
Figure 4: XRD pattern of the AW specimen indicating the presence of
chromium carbides
Slika 4: Rentgenski (XRD) posnetek, ki ka`e prisotnost kroma v
vzorcu AW
Figure 3: Microstructure of the AW specimen: a) weld metal; b) HAZ
Slika 3: Mikrostruktura AW-vzorcev: a) zvar, b) toplotno vplivana
cona (HAZ)
Figure 2: Microstructure of the AISI 422 base metal
Slika 2: Mikrostruktura osnovnega materiala AISI 422
The heat affected zone (HAZ) between the weld and the
base metal again reveals a microstructure consisting of
martensite and austenite phases (Figure 6b). Compared
to the AW and PH specimens, the martensite phase in the
TP specimen has a coarser appearance due to the effects
of tempering.
The hardness test results are summarized in Figure 7.
Microhardness measurements were made both across
and along the weld centerline. Figure 7a indicates a
decrease in the microhardness towards the weld metal for
all the welded joints. This is expected, since the base
metal consists of martensite and the weld metal consists
of either austenite or a mixture of austenite and some
delta ferrite. Figure 7b shows the microhardness distri-
bution for the weld from top to bottom. While the
microhardness distribution for the AW specimen is more
or less constant, a steady decrease is observed for the PH
and TP specimens. As welding was carried out in 13
passes, each pass resulted in a tempering treatment on
the previous pass, which led to the softening of the
underlying layers.
The tensile-test results for the welded specimens are
listed in Table 5. Failures occurred within the weld
metal. The YS showed no change with preheating and
remained constant at 490 MPa. The tempering treatment
A. CALIK, M. S. KARAKAª: EFFECT OF HEAT TREATMENT ON THE MICROSTRUCTURE ...
406 Materiali in tehnologije / Materials and technology 47 (2013) 4, 403–407
Figure 7: Microhardness distribution for the welded joints: a) across
the weld centerline and b) on the weld centerline from top to bottom
Slika 7: Razporeditev mikrotrdote v varjenih spojih: a) preko sredine
zvara in b) na sredini zvara od vrha do dna
Figure 6: Microstructure of the TP specimen: a) weld metal; b) HAZ
Slika 6: Mikrostruktura TP-vzorca: a) zvar, b) toplotno vplivana cona
(HAZ)
Figure 5: Microstructure of the PH specimen: a) weld metal; b) HAZ
Slika 5: Mikrostruktura PH-vzorca: a) zvar, b) toplotno vplivana cona
(HAZ)
resulted in a decrease in the YS to 467 MPa. The UTS
increased from 630 MPa to 660 MPa with preheating.
The tempering further increased the UTS to 683 MPa.
The elongation values increased with preheating and
tempering treatments. With preheating, the fraction of
EL values increased from 12.5 % to 18.8 %. With
tempering, the fraction of EL values increased to 25 %.
Table 5: Tensile-test results for the welded AISI 422 alloy for three
different heat treatments
Tabela 5: Rezultati nateznih preizkusov varjene AISI 422 zlitine pri
treh razli~nih toplotnih obdelavah
Specimen Yield strengthYS
Ultimate
tensile strength
UTS
Percent
elongation
%
AW 490 630 12.5
PH 490 660 18.8
TP 467 683 25.0
The fatigue-test results are given in Figure 8. Despite
its lower tensile strength (Figure 5), the unwelded speci-
men shows a superior fatigue strength compared to its
welded counterparts. The fatigue limit of the unwelded
specimen lies at 458 MPa, where the PH, TP and AW
specimens display the fatigue limits at (345, 312 and
305) MPa, respectively. These results indicate that the
fatigue strength is more dependent on the hardness of a
weld. The unwelded specimen is completely martensitic
and therefore displays the highest fatigue limit. The AW
specimen has the highest hardness among the welded
specimens, but it displays the lowest fatigue limit which
is probably due to residual stresses. The TP specimen
has the lowest hardness distribution among the welded
specimens displaying a fatigue limit which is only
slightly higher than that of the AW specimen. The PH
specimen, on the other hand, displays the highest fatigue
limit among the welded specimens, which is due to its
moderate hardness and also due to the stress relief
provided by the preheating.
4 CONCLUSIONS
The effect of preheating and post-weld tempering on
the microstructure and mechanical properties of welded
martensitic stainless steels was studied. The conclusions
drawn from this study are as follows:
1) The grain boundary precipitation of chromium
carbides was observed in the welded joints. The
precipitation was mainly caused by the diffusion of
carbon into the austenitic stainless steel during the
welding.
2) A decrease in the microhardness was observed
moving from the base metal to the weld. The decrease in
and absence of the martensite phase in the HAZ and the
weld, respectively, were the main causes for a decreased
hardness.
3) An increase in the tensile strength and percent
elongation was observed after the preheating and
tempering heat treatments.
4) Preheating provided the best fatigue results for the
welded specimens. The fatigue limits observed for the
tempered specimens were slightly lower than those
found in the preheated specimens.
5 REFERENCES
1 Welding Handbook, 9th ed., Vol. 1, Welding Science and Techno-
logy, American Welding Society, Miami 2007
2 S. J. Maddox, Recent advances in the fatigue assessment of weld
imperfections, Weld. J., 7 (1992), 42–50
3 G. S. Booth, J. G. Wylde, Procedural considerations relating to the
fatigue testing of steel weldments, in Fatigue and Fracture Testing of
Weldments, ASTM STP 1058, Eds., H. I. McHenry, J. M. Potter,
American Society for Testing and Materials, Philadelphia 1990, 3–15
4 T. N. Nguyen, M. A. Wahab, The effect of weld geometry and resi-
dual stresses on the fatigue of welded joints under combined loading,
J. Mater. Proc. Technol., 77 (1998), 201–208
5 G. Linnert, Common welding problems: Fatigue, Weld. J., 75 (1996),
59–60
6 A. Ohta, Y. Maeda, N. Suzuki, Fatigue strength of butt-welded joints
under constant maximum stress and random minimum stress condi-
tions, Fatigue Fract. Eng. M., 19 (1996) 19, 265–275
7 J. A. Ferreira, C. M. Branco, Fatigue analysis and prediction in fillet
welded joints in the low thickness range, Fatigue Fract. Eng. M., 13
(1990), 201–212
8 G. Magudeeswaran, V. Balasubramanian, G. Madhusudhan Reddy, T.
S. Balasubramanian, Effect of welding processes and consumables
on tensile and impact properties of high strength quenched and tem-
pered steel joints, J. Iron Steel Res. Int., 15 (2008), 87–94
9 L. R. Link, Fatigue crack growth of weldments, in Fatigue and Frac-
ture Testing of Weldments, ASTM STP 1058, Eds., H. I. McHenry,
J. M. Potter, American Society for Testing and Materials, Phi-
ladelphia 1990, 16–33
10 M. Ghosh, K. Kumar, R. S. Mishra, Friction stir lap welded advanced
high strength steels: Microstructure and mechanical properties,
Mater. Sci. Eng., A 528 (2011), 8111–8119
11 ASM Handbook, Vol. 6, Welding, Brazing and Soldering, 2nd ed.,
Metals Park, ASM International, Ohio 1993
12 A. J. Williams, P. J. Rieppel, C. B. Voldrich, Literature survey on
weld-metal cracking, WADC Technical Report, 52–143, Ohio 1952
A. CALIK, M. S. KARAKAª: EFFECT OF HEAT TREATMENT ON THE MICROSTRUCTURE ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 403–407 407
Figure 8: Wöhler (S – N) diagram showing fatigue test results
Slika 8: Wöhlerjeva krivulja (S – N) prikazuje rezultate preizkusov
utrujenosti

N. [PI^KA, P. FORTE TAV^ER: NEW COMBINED BIO-SCOURING AND BIO-BLEACHING PROCESS ...
NEW COMBINED BIO-SCOURING AND
BIO-BLEACHING PROCESS OF COTTON FABRICS
NOV, ZDRU@ENI POSTOPEK BIOIZKUHAVANJA IN BIOBELJENJA
BOMBA@NIH TKANIN
Nina [pi~ka, Petra Forte Tav~er
University of Ljubljana, Faculty of Natural Sciences and Engineering, Department of Textiles, Sne`ni{ka 5, 1101 Ljubljana, Slovenia
nina.spicka@ntf.uni-lj.si
Prejem rokopisa – received: 2012-08-27; sprejem za objavo – accepted for publication: 2013-01-03
A new commercial bio-bleaching process for cellulose fibres was investigated in this study. The process runs enzymatically with
arylesterase enzymes (EC 3.1.1.2) and hydrogen peroxide. The enzyme system catalyses the perhydrolysis of propylene glycol
diacetate. During the reaction propylene glycol and peracetic acid as a bleaching agent are formed in situ. The main advantage
of the bleaching with peracetic acid is that a satisfactory degree of whiteness of a cotton fabric can be obtained at 65 °C at a
neutral pH. The bleaching performance of the new bio-bleaching process on a traditionally alkaline-scoured and bio-scoured
100 % cotton fabric and the feasibility of a one-bath bio-scouring/bio-bleaching pre-treatment were investigated. The whiteness
degrees, tenacities at maximum loads and water absorbencies of the treated cotton fabrics were compared. The hydrogen
peroxide and peracetic acid concentrations, pH and TOC values of the remaining treatment solutions were measured. The
research showed that the new bio-bleaching system has a powerful bleaching ability under mild process conditions and that
bio-scouring and bio-bleaching can be efficiently combined in a one-bath process.
Keywords: cotton, bio-scouring, bio-bleaching, peracetic acid
V {tudiji je bil raziskan nov komercialni postopek biobeljenja celuloznih vlaken. Proces poteka encimsko z encimi
arilesterazami (EC 3.1.1.2) in vodikovim peroksidom. Med reakcijo se v kopeli tvorita propilen glikol in perocetna kislina kot
belilno sredstvo. Poglavitna prednost beljenja s perocetno kislino je zadostna stopnja beline bomba`ne tkanine, dose`ena pri 65 °C
v nevtralnem pH. Raziskana je bila belilna sposobnost novega biobelilnega postopka na tradicionalno alkalno izkuhani in
bioizkuhani 100 % bomba`ni tkanini in mo`nost enokopelnega procesa bioizkuhavanja in biobeljenja. Pri obdelanih bomba`nih
tkaninah je bila izmerjena stopnja beline, natezna trdnost ob maksimalni obremenitvi in vodovpojnost. Pri odpadnih
obdelovalnih kopelih je bila izmerjena koncentracija vodikovega peroksida in perocetne kisline, pH in TOC-vrednost. Raziskava
je pokazala, da ima nov biobelilni postopek mo~no belilno sposobnost v blagih procesnih razmerah in da sta lahko postopka
bioizkuhavanja in biobeljenja uspe{no zdru`ena v enokopelni proces.
Klju~ne besede: bomba`, bioizkuhavanje, biobeljenje, perocetna kislina
1 INTRODUCTION
The most important natural cellulose fibre is cotton,
whose use is constantly increasing. Natural cotton is
highly hydrophobic and slightly coloured. To prepare the
fibres for further treatment and use, pretreatment pro-
cesses are needed. With scouring, non-cellulose substan-
ces (wax, pectin, proteins, hemicelluloses, etc.) that
surround the fibre cellulose core are removed and, as a
result, fibres become hydrophilic. With bleaching the
natural pigments of cotton fibres are removed and fibres
become white. Traditionally, scouring and bleaching
processes are conducted at the temperatures up to 120 °C
in a very alkaline medium at a pH of 10 to 12. In these
treatments large amounts of auxiliary agents are added.
Due to the high working temperatures, a lot of energy is
consumed. Large amounts of water are used to rinse and
neutralise the alkaline-scoured and bleached fabrics.
Consequently, the textile industry is considered to be one
of the biggest water, energy, and chemical consumers. To
comply with the increasingly more rigorous environmen-
tal regulations and to save water and energy, biotechno-
logy and several types of enzymes have entered the
textile sector. Bio-scouring with enzyme pectinases is an
alternative to sodium hydroxide scouring in the removal
of non-cellulose substances from the cotton-fibre sur-
face. The process occurs at moderate temperatures in a
slightly acidic or alkaline medium which is dependent on
the type of pectinases.1,2
Some of the alternatives to bleaching with hydrogen
peroxide (HP) have been explored in the textile, and pulp
and paper industries, i.e., the enzymatic bleaching using
peroxidases,3 laccase/mediator system and glucose
oxidases4–6 and bleaching with the peracids either pro-
duced industrially7 or generated in situ from different
bleach activators.8
A new commercial bio-bleaching product, Gentle
Power Bleach from Huntsman, is available on the
market. It functions enzymatically with arylesterase
enzymes (EC 3.1.1.2) and HP. The enzyme system
catalyses the perhydrolysis of propylene glycol diacetate.
During the reaction propylene glycol and peracetic acid
(PAA) as the bleaching agents are formed in situ. If
bio-scouring and bio-bleaching could be combined into
one process, large amounts of water, energy, time and
auxiliary agents would be saved.
The objective of our work was to investigate the
bleaching performance of the new bio-bleaching process
Materiali in tehnologije / Materials and technology 47 (2013) 4, 409–412 409
UDK 677:677.21:677.027.262 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)409(2013)
on a traditionally alkaline-scoured and bio-scoured
100 % cotton fabric. Secondly, the feasibility of a one-
bath process combining bio-scouring and bio-bleaching
pre-treatments was investigated. The whiteness degree,
tenacity at maximum load and water absorbency of the
treated cotton fabrics were evaluated. After the treat-
ments, the pH and TOC (total organic carbon) values of
the remaining baths were measured. The quantities of
HP and PAA in the remaining bleaching baths were also
measured.
2 EXPERIMENTAL WORK
2.1 Materials
Desized cotton fabric, 100 g/m2, was obtained from
Tekstina, Slovenia. Beisol PRO (a commercial pectinase
solution), Felosan RG-N (a non-ionic wetting agent),
Cotoblanc HTD-N (an anionic wetting and dispersing
agent) and Lawotan RWS (a non-ionic wetting agent)
were supplied from CHT, Germany. The Gentle Power
Bleach chemicals (CLARITE® LTC, INVATEX® LTA,
INVAZYME® LTE) were supplied from Huntsman,
Switzerland. H2O2 35 % (HP) was obtained from
Belinka, Slovenia and Stabilizer SIFA from Clariant.
NaOH and Na2CO3 were purchased from Sigma Aldrich.
2.2 Treatment methods
The desized cotton fabric was treated according to
the procedures presented in Table 1. Two parallel sets of
experiments were undertaken. The treatments were
performed in a laboratory dyeing machine DL-6000 Plus
from Starlet in the 500 ml bakers at a liquor ratio of 1 :
20. After each treatment the samples were washed in hot
water, rinsed twice in cold water and air dried.
Table 1: Abbreviations of the treatments and treated samples with
treatment descriptions
Tabela 1: Okraj{ava obdelave in obdelanega vzorca z opisom obde-
lave
Treatment/sample
abbreviation Treatment description
AS alkaline scouring
TB (AS) two-bath alkaline scouring andtraditional bleaching with HP
BB (AS) two-bath alkaline scouring andbio-bleaching
BS bio-scouring
BB (BS) two-bath bio-scouring and bio-bleaching
BS/BB one-bath bio-scouring and bio-bleaching
BS/BB+ one-bath bio-scouring and bio-bleachingwith an additional final temperature rise
Alkaline scouring (AS) was carried out in a bath
containing 2 g/l of Cotoblanc HTD-N and 3 g/l of NaOH
at 95 °C for 40 min. Traditional bleaching (TB) was
carried out in a bath containing 10 ml/l of H2O2 (35 %),
0.1 g/l of Lawotan RWS, 0.5 g/l of Stabilizer SIFA and
4 g/l of NaOH at 90 °C for 60 min. The ingredients and
data conditions for bio-treatments are presented in
Tables 2 and 3.
Table 2: Ingredients of bio-treatments
Tabela 2: Sestavine bioobdelav
Bio-treatment
Ingredient BB(AS) BS
BB
(BS)
BS/
BB
BS/
BB+
1 g/l Felosan RG-N – + + – –
2 g/l Beisol PRO – + + + +
1.5 g/l CLARITE® LTC + – + + +
3 g/l INVATEX® LTA + – + + +
6 ml/l H2O2 35 % + – + + +
1 g/l INVAZYME® LTE + – + + +
2 g/l Na2CO3 + – + + +
+ the ingredient is included in the treatment bath
– the ingredient is not included in the treatment bath
Table 3: Data conditions for bio-treatments
Tabela 3: Podatki pogojev bioobdelave
Bio-treatment conditions
BB (AS) BS BB (BS) BS/BB BS/BB+
Bio-
scouring –
55 °C,
15 min
85 °C,
15 min
55 °C,
15 min
85 °C,
15 min
– –
Bio-
bleaching
65 °C,
60 min –
65 °C,
60 min
65 °C,
60 min
65 °C,
60 min
Final
treatment – – – –
85 °C,
15 min
– the treatment step was not included
2.3 Analytical methods
Prior to the measurements, samples were conditioned
for 24 h at 20 °C and 65 % relative humidity. The pH
was measured using a pH meter MA5740 (Iskra,
Slovenia). The rates of PAA and HP were measured with
iodometric titrations.5 The degree of whiteness was
measured with a spectrophotometer Spectraflash SF600
Plus (Datacolor, Switzerland) using the CIE method,
according to EN ISO 105-J02:1997(E). The water
absorbency was measured according to DIN 53 924 (the
velocity of the soaking water for textile fabrics, the
method for determining the wicking height). The
measurements of the tenacity at maximum load were
performed on a Tensile Tester Model 5567 (Instron,
USA). The total organic carbon (TOC) was measured
with a TOC-5000A (Shimadzu, Japan) according to ISO
8245.
3 RESULTS AND DISCUSSION
3.1 Concentration of oxidants during the bio-bleach-
ing process
The measurements confirmed that during the
bio-bleaching process HP is converted into PAA (Figure
1). Almost 0.02 mol/l of PAA was produced and con-
N. [PI^KA, P. FORTE TAV^ER: NEW COMBINED BIO-SCOURING AND BIO-BLEACHING PROCESS ...
410 Materiali in tehnologije / Materials and technology 47 (2013) 4, 409–412
sumed during the bleaching process. During the bleach-
ing with PAA, acetic acid is formed and the pH de-
creases. For this reason the starting pH was set to be
highly alkaline (pH 10.3). During the process the pH
decreased to 7.6.
3.2 Fabric analysis
3.2.1 Whiteness
The achieved degrees of whiteness (W) are presented
in Table 4. The alkaline-scoured sample (AS) had a
whiteness of 20.7 and the bio-scoured sample (BS) had a
whiteness of 10. After the alkaline scouring, the fibres
swelled, became smoother and clean of non-cellulose
impurities and the degree of whiteness increased.
Alkaline scouring is more intensive and removes some of
the coloured substances from the fibre that the bio-
scouring does not. The whiteness degrees of both
samples increased significantly after the bleaching
process. The highest whiteness (84.5) was obtained on
the HP-bleached and alkaline-scoured sample (TB (AS)).
Not much less (81.7) was obtained on the alkaline-
scoured and bio-bleached sample (BB (AS)), while the
whiteness degrees of bio-scoured and bio-bleached
samples were lower: 73.5 on the separately bio-scoured
and bio-bleached sample (BB (BS)), 67.8 on the
one-bath bio-scoured and bio-bleached sample (BS/BB)
and 72.8 on the one-bath bio-scoured and bio-bleached
sample with the final temperature rise. As it can be seen
in Figure 1, a high quantity of HP remains in the
bio-bleaching bath. At a pH bellow 10 °C and at 65 °C
HP is very stable and does not contribute to the bleach-
ing of the fibres. The rising of the temperature at the end
of the bio-bleaching process activated the remaining HP,
which further increased the degree of whiteness.
3.2.2 Tenacity at maximum load
There were no significant differences in the tenacity
values among all the treated samples in the weft direc-
tions whereas the tenacity values in the warp directions
were very different. The highest value was measured on
the alkaline-scoured sample (AS) and on most of the
bio-bleached samples. The high tenacity of the AS
sample is a result of the contraction of the fabric exposed
to a high process temperature. After the HP bleaching of
the alkaline-scoured sample (TB (AS)) the tenacity at
maximum load decreased significantly. On the other
hand, the tenacity after the bio-bleaching of the alka-
line-scoured sample remained similar (BB (AS)). These
results confirm that during a bio-bleaching process the
fabric does not get damaged.
3.2.3 Water absorbency
The remaining substances influence the water
absorbency. All the treated samples revealed very good
absorption properties. The highest rising height was
measured on the traditionally alkaline-scoured and
HP-bleached sample (TB (AS)). The rising height of the
bio-scoured samples was a bit lower but the differences
among all the treated samples were not significant. All
the samples could be considered to be absorbent.
Table 4: Whiteness degree (W), tenacity at maximum load (s) and
wicking height after 300 s of various sample treatments
Tabela 4: Stopnja beline (W), trdnost ob maksimalni obremenitvi (s)
in vi{ina kapilarnega dviga v ~asu 300 s razli~no obdelanih vzorcev
W s/(cN/10–6 kg m–1)
Wicking height
(cm)
Sample warp weft warp weft
AS 20.68 26.41 16.09 – –
TB (AS) 84.51 18.36 13.01 6.27 5.67
BB (AS) 81.68 25.50 13.57 – –
BS 10.11 22.08 13.20 5.62 4.75
BB (BS) 73.47 19.86 12.87 5.12 4.62
BS/BB 67.84 25.99 14.01 5.00 4.15
BS/BB+ 72.77 25.52 15.93 5.10 4.27
3.3 Ecological parameters of the remaining treatment
baths
The ecological parameters – the remaining concen-
trations of HP and PAA, the final pH and TOC values –
are presented in Table 5. High quantities of the
unconsumed HP remained in all the bleaching baths,
especially in the traditional one. In all the bio-bleaching
baths PAA was detected, but the concentrations were
low, meaning the PAA was consumed for the bleaching
of the cotton fabric.
Traditional alkaline scouring and bleaching are con-
ducted in an alkaline environment. These baths should be
neutralised prior to drainage into the sewage system.
During the neutralisation, salts that additionally load
wastewaters are produced. The final pH value of the
bio-scouring bath was 6 and it was around 7.5 for the
N. [PI^KA, P. FORTE TAV^ER: NEW COMBINED BIO-SCOURING AND BIO-BLEACHING PROCESS ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 409–412 411
Figure 1: Changes in the concentrations of hydrogen peroxide (HP)
and peracetic acid (PAA) and change in the pH during bio-bleaching
of an alkaline-scoured sample (during the first 15 min the bath was
heated to 65 °C)
Slika 1: Sprememba koncentracije vodikovega peroksida (HP) in
perocetne kisline (PAA) ter sprememba pH med biobeljenjem alkalno
izkuhanega vzorca (za~etnih 15 min je bila kopel ogrevana na 65 °C)
bio-bleaching baths. Since none of these treatment
processes requires a neutralisation of the fabric, the treat-
ment process can be shorter and less expensive.
Bio-scouring, traditional bleaching and alkaline
scouring baths had low TOC values and all the bio-
bleaching baths had significantly higher values. A bio-
bleaching bath includes more additives than other
processes; especially enzymes contribute a lot of organic
carbon. On the other hand, enzymatic processes require
much less water, energy and time, which compensate for
high TOC values from the ecological point of view.
Nevertheless, all the treatments exceeded the limit TOC
values (60 mg C/l) for direct drainage into the sewage
system.
Table 5: Ecological parameters of the remaining treatment baths
(concentrations of hydrogen peroxide – HP and peracetic acid – PAA,
final pH and TOC values)
Tabela 5: Ekolo{ki parametri odpadnih obdelovalnih kopeli (koncen-
tracija vodikovega peroksida – HP in perocetne kisline – PAA, kon~ni
pH in vrednosti TOC)
c (HP)/
(mol/l)
c (PAA)/
(mol/l) pH
TOC
/(mg/l)Treatment
AS – – 11.0 1096
TB (AS) 0.1084 – 11.7 558
BB (AS) 0.0542 0.0032 7.6 2372
BS – – 5.7 510
BB (BS) 0.0531 0.0034 7.4 2510
BS/BB 0.0588 0.0042 7.6 2673
BS/BB+ 0.0608 0.0012 7.4 2842
4 CONCLUSIONS
The research showed that the new bio-bleaching
system with enzymatically gained peracetic acid has a
powerful bleaching ability under mild process con-
ditions, i.e., a neutral pH and a temperature of 65 °C.
With the new bio-bleaching process the fabrics with a
high whiteness degree, good water absorbency and high
tenacity at maximum load were obtained having proper-
ties comparable to the results of the traditional process.
During the bio-bleaching process the cotton fabric did
not get damaged as during the traditional pre-treatment.
A slightly higher degree of whiteness was obtained on
the alkaline-scoured cotton fabric compared to the bio-
scoured one. Bio-scouring and bio-bleaching can be
efficiently combined in a one-bath process. The final
temperature rise of the one-bath process contributes to a
higher whiteness degree. This kind of a one-bath process
is shorter, consuming less energy and hence being less
expensive.
Acknowledgements
This work was supported by the Slovenian Research
Agency (P2-0213). N. [. thanks the Ministry of Higher
Education, Science and Technology for a Ph.D. grant
(1000-10-310153).
5 REFERENCES
1 P. Pre{a, P. Forte Tav~er, Low water and energy saving process for
cotton pretreatment, Textile Research Journal, 79 (2009), 76–88
2 P. Pre{a, P. Forte Tav~er, Bioscouring and bleaching of cotton with
pectinase enzyme and peracetic acid in one bath, Color Technol., 124
(2008), 36–42
3 K. Opwis, D. Knittel, E. Schollmeyer, P. Hoferichter, A. Cordes,
Simultaneous Application of Glucose Oxidases and Peroxidases in
Bleaching Process, Eng. Life. Sci., 8 (2008), 175–178
4 T. Tzanov, S. Costa, G. Gübitz, A. Cavaco-Paulo, Hydrogen peroxide
generation with immobilized glucose oxidase for textile bleaching,
Journal of Biotechnology, 93 (2002), 87–94
5 G. Buschle-Diller, X. D. Yang, Enzymatic Bleaching of Cotton
Fabric with Glucose Oxidase, Textile Research Journal, 71 (2001),
388–394
6 P. Forte Tav~er, Low-temperature bleaching of cotton induced by
glucose oxidase enzymes and hydrogen peroxide activators, Biocatal.
biotransform., 30 (2012), 20–26
7 P. Kri`man, F. Kova~, P. Forte Tav~er, Bleaching of cotton fabric
with peracetic acid in the presence of different activators, Coloration
Technology, 121 (2005), 304–309
8 S. H. Lim, N. C. Gürsoy, P. Hauser, D. Hinks, Performance of a new
cationic bleach activator on a hydrogen peroxide bleaching system,
Coloration Technology, 120 (2004), 114–118
N. [PI^KA, P. FORTE TAV^ER: NEW COMBINED BIO-SCOURING AND BIO-BLEACHING PROCESS ...
412 Materiali in tehnologije / Materials and technology 47 (2013) 4, 409–412
E. ELMAS et al.: EFFECT OF MECHANICAL ACTIVATION ON MULLITE FORMATION IN AN ALUMINA-QUARTZ ...
EFFECT OF MECHANICAL ACTIVATION ON MULLITE
FORMATION IN AN ALUMINA-QUARTZ CERAMICS
SYSTEM
VPLIV MEHANSKE AKTIVACIJE NA NASTANEK MULLITA V
KERAMI^NEM SISTEMU GLINICA-KREMEN
Eda Elmas1, Kenan Yildiz2, Nil Toplan2, H. Özkan Toplan2
1Kaleseramik, Çanakkale Kalebodur Ceramic Ind. Inc., 17430 Çan, Çanakkale, Turkey
2Sakarya University, Metallurgy and Materials Engineering Dept., 54187 Sakarya, Turkey
toplano@sakarya.edu.tr
Prejem rokopisa – received: 2012-08-31; sprejem za objavo – accepted for publication: 2013-01-04
Powder mixtures of alumina and quartz were being mechanically activated in a planetary mill for 2 h. Both non-activated and
activated samples were sintered at different temperatures (1250, 1300, 1325, 1350 and 1375) °C for (1, 2, 3 and 5) h and the
formation of a mullite phase was examined with an X-ray diffraction analysis. It was determined that the mechanical activation
increased the quantity of the mullite phase.
Keywords: mullite, mechanical activation, amorphization, X-ray diffraction, alumina, quartz
Me{anice prahov glinice in kremena so bile 2 h mehansko aktivirane v planetarnem mlinu. Neaktivirani in aktivirani vzorci so
bili sintrani (1, 2, 3 in 5) h pri razli~nih temperaturah (1250, 1300, 1325, 1350 in 1375) °C), z rentgensko difrakcijsko analizo pa
je bil preiskovan nastanek mullitne faze. Ugotovljeno je bilo, da mehanska aktivacija pove~a koli~ino mullitne faze.
Klju~ne besede: mullit, mehanska aktivacija, amorfizacija, rentgenska difrakcija, glinica, kremen
1 INTRODUCTION
Mullite (3Al2O3 · 2SiO2) is an important ceramic pha-
se in conventional ceramics (such as tableware, construc-
tion ceramics and refractories), advanced high-tempera-
ture structural materials, heat exchangers, catalysator
convertors, filters, optical devices and electronic pack-
aging materials. However, mullite suffers from its relati-
vely low fracture toughness which limits its application
in industrial use. The conventional route for the pre-
paration of mullite is the solid-state reaction between
alumina and silica, which is controlled with diffusion.
The mullite formation with this method takes place at a
relatively high temperature (>1500 °C). The mullitiza-
tion temperature and the morphology of mullite particles
depend on the particle size of the initial raw materials
and the preparation of the precursors before sintering.
Mullite has been synthesized in many ways like simple
sintering of alumina and silica powders, sol-gel method,
co-precipitation, hydrothermal and chemical-vapor-depo-
sition processes. The mullitization temperature is as high
as 1600 °C for the conventional fabrication method, i.e.,
the solid-state reaction of high-purity alumina and
quartz.1–3
Mechanical activation of starting materials is a pro-
mising method for a precursor preparation. The particle-
size reduction, which increases the contact surfaces
between the particles, is a direct consequence of milling.
Also, the energy of the system increases resulting in a
decrease in the reaction temperature.4 Different pro-
cesses can remarkably influence the reactivity of the
solids. Particularly, the mechanical treatments are
important as long as they can help to produce the
changes in the texture and structure of the solids. In
many cases, these alterations in the structure cause cer-
tain modifications in the phases formed with a thermal
treatment of the solids that were mechanochemically
treated.5,6
In this work, the effects of mechanical activation on
structural disordering (amorphization) in an alumina and
quartz ceramics system and a formation of mullite were
analyzed using X-ray diffraction (XRD) and scanning
electron microscopy (SEM).
2 EXPERIMENTAL METHODS
Alumina and quartz were supplied from the Çelvit
Ceramic Company, Turkey. The chemical composition of
alumina is 99.425 % Al2O3, 0.52 % SiO2 and 0.055 %
Na2O. The composition of quartz is 99.1 % SiO2, 0.28 %
Al2O3, 0.16 % CaO + MgO, 0.17 % K2O + Na2O, 0.05 %
Fe2O3, 0.05 % TiO2 and 0.19 % LOI. Alumina and quartz
were mixed to obtain the appropriate stoichiometric ratio
according to the chemical formula of mullite (3Al2O3 ·
2SiO2) in ashless rubber-lined ceramic jars for 2 h using
zirconia balls and distilled water as the milling media.
After drying, the mixture was made in a high-energy
planetary ball mill (Fristch) with a rotation speed of 600
r/min. The ball-to-powder weight ratio was adjusted to
20. The precursor milling was carried out for 2 h. An
Materiali in tehnologije / Materials and technology 47 (2013) 4, 413–416 413
UDK 666.3/.7 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)413(2013)
X-ray diffraction analysis was performed using a Rigaku
Ultima X-ray diffractometer and CuK radiation. A Joel
6060 LV scanning electron microscope was used for the
morphological analysis of non-activated and activated
mixed powders. The degree of amorphization (A/%) of
the mechanically activated powder was calculated from
equation (1):7,8
%A
B I
B I
x
x
= −
⎡
⎣⎢
⎤
⎦⎥
×1 1000
0
(1)
where Io is the integral intensity of the diffraction peak
for the non-activated mixture, Bo is the background of
the diffraction peak for the non-activated mixture and Ix
and Bx are the equivalent values for the mechanically
activated mixture. After the activation, both non-activa-
ted and activated powders were sintered in an electrical
furnace with a heating rate of 10 °C/min at (1250, 1300,
1325, 1350 and 1375) °C for (1, 2, 3 and 5) h and the
formulation of the mullite phase was examined with
XRD. In addition, the degree of mullite crystallization
(C/%) of non-activated and activated powders at
different firing temperatures and times was calculated
from equation (2):
%C
B I
B I
x
x
=
⎡
⎣⎢
⎤
⎦⎥
×0
0
100 (2)
where Io is the integral intensity of the diffraction peak
for the reference mullite powders, Bo is the background
of the diffraction peak for the non-activated mixture and
Ix and Bx are the equivalent values for the mullite
crystallization (the reference mullite is Nabaltec mullite
and the JCPDS card number is 01–079–1454). The
integrated intensity of the (121) XRD reflection (2
from 40.5° to 41.5°) was measured to determine to
content of the mullite phase.
3 RESULTS
3.1 Structural changes in the activated alumina-quartz
mixture powders
The X-ray diffraction analysis of non-activated and
activated mixture powders is given in Figure 1. A com-
parison of the peaks from the two diffraction patterns
shows that all the diffraction peaks get shorter after a
mechanically activation. This reflects the partial
amorphization and structural disordering in alumina and
quartz. Mechanical activation has already been reported
to amorphize materials.8
The scanning electron micrographs (SEM) of non-
activated and activated alumina and quartz mixture
powders can be observed in Figure 2a, b. The particle
size in the non-activated mixture is over 5 μm (Figure
2a). After the mechanical activation, the mixture of the
powders is agglomerated. The degrees of amorphization
of alumina and quartz were founded to be approximately
70 % and 85 %, respectively.
3.2 Mullite formation
Figure 3a, b shows the XRD patterns of non-
activated and activated quartz-alumina mixture powders
fired at 1250 °C and 1375 °C for 60 min. It is found that
the intensity of the mullite phase in the activated sample
increases. Figure 4a, b shows the mullite content of the
non-activated and activated quartz-alumina mixture
powders sintered at 1250–1375 °C for different times. In
the samples sintered at 1250 °C for 60 min, the mullite
content gradually increased from mass fraction 8.13 % to
45 % during the activation. On the other hand, when they
E. ELMAS et al.: EFFECT OF MECHANICAL ACTIVATION ON MULLITE FORMATION IN AN ALUMINA-QUARTZ ...
414 Materiali in tehnologije / Materials and technology 47 (2013) 4, 413–416
Figure 2: SEM micrographs of: a) non-activated and b) activated
alumina-quartz powders
Slika 2: SEM-posnetek: a) neaktiviranega in b) aktiviranega prahu
glinica-kremen
Figure 1: XRD patterns of non-activated and activated alumina-quartz
powders (A: alumina, Q: quartz)
Slika 1: XRD-posnetek neaktiviranega in aktiviranega prahu glinica-
kremen (A: glinica, Q: kremen)
were sintered at 1375 °C for 60 min, the mullite content
increased from 14.2 % to 92.3 % during the activation.
Mechanical treatment in a high-energy mill generates
a stress field within the solids. Stress relaxation can
occur via several mechanisms: (1) heat release, (2)
development of a surface area as a result of the brittle
fracture of the particles, (3) generation of various sorts of
structural defects and (4) stimulation of chemical
reactions within the solids. All of the relaxation channels
cause changes in the reactivity of the solid substance that
is under treatment, which is why the resulting action is
called mechanical activation.9 The concentration of the
mechanically induced defects and their spatial
distribution depend upon the condition of the energy
transfer in the mill. The creation of defects enhances the
stored energy in the solids and consequently causes a
decrease in the activation barrier to further processing of
the solids.10
4 CONCLUSION
The effect of mechanical activation on mullite for-
mation in a quartz-alumina ceramics system was studied
by using XRD and SEM. The mechanical activation
caused amorphization and structural disordering of the
quartz-alumina mixture. The degree of amorphization of
quartz and alumina were found to be approximately
70 % and 85 %, respectively. The application of high-
energy milling allowed a dramatic change in the struc-
ture and surface performance of alumina and quartz, so
the content of mullite was increased with mechanical
activation.
5 REFERENCES
1 P. M. Souto, R. R. Menezes, R. H. G. A. Kiminami, Sintering of
commercial mullite powder: Effect of MgO dopant, J. Mater. Process
Tech., 209 (2009), 548–53
2 Y. Dong, X. Feng, X. Feng, Y. Ding, X. Liu, G. Meng, Preparation of
Low-Cost Mullite Ceramics From natural Bauxite and Industrial
Waste Fly Ash, Journal of Alloys and Compounds, 460 (2008),
599–606
3 V. Viswabaskaran, F. D. Gnanam, M. Balasubramanian, Mullitisation
Behaviour of South Indian Clays, Ceramics International, 28 (2002),
557–564
E. ELMAS et al.: EFFECT OF MECHANICAL ACTIVATION ON MULLITE FORMATION IN AN ALUMINA-QUARTZ ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 413–416 415
Figure 4: Mass fractions of the mullite content in: a) non-activated
and b) activated sintered samples
Slika 4: Masni dele`i mullita v: a) neaktiviranih in b) aktiviranih sin-
tranih vzorcih
Figure 3: Comparison of the XRD patterns for the reference mullite
with the XRD patterns for non-activated and activated quartz-alumina
mixtures fired at: a) 1250 °C for 60 min and b) 1375 °C for 60 min
Slika 3: Primerjava XRD-posnetkov referen~nega mullita z XRD-po-
snetki neaktivirane in aktivirane me{anice kremena in glinice: a) 60
min segrevanega na 1250 °C in b) 60 min segrevanega na 1375 °C
4 N. Behmanesh, S. Heshmati-Manesh, A. Ataie, Role of mechanical
activation of precursors in solid state processing of nano-structured
mullite phase, J. Alloy. Compd., 450 (2008), 421–5
5 S. Tamborenea, A. D. Mazzoni, E. F. Aglietti, Mechanochemical
activation of minerals on the cordierite synthesis, Thermochimica
Acta, 411 (2004), 219–24
6 A. D. Mazzoni, E. F. Aglietti, E. Pereira, Preparation of Spinel
Powders at low temperature by mechanical activation, Latin. Am.
Res., 21 (1991), 63–8
7 S. M. Ohlberg, D. W. Strickler, Determination of percent crystallinity
of partly devitrified glass by X-ray diffraction, J. Am. Ceram. Soc.,
45 (1962), 170–171
8 P. Balaz, Mechanochemistry in nanoscience and minerals engineer-
ing, Springer-Verlag, Berlin 2008
9 V. V. Boldyrev, T. Tkacova, Mechanochemistry of solids: past, pre-
sent, and prospects, J. Mater. Synt. Proc., 8 (2000) 3/4, 121–32
10 U. Steinike, K. Tkacova, Mechanochemistry of solids-real structure
and reactivity, J. Mater. Synt. Proc., 8 (2000) 3/4, 197–203
E. ELMAS et al.: EFFECT OF MECHANICAL ACTIVATION ON MULLITE FORMATION IN AN ALUMINA-QUARTZ ...
416 Materiali in tehnologije / Materials and technology 47 (2013) 4, 413–416
P. BERNARDIN et al.: DETERMINATION OF THE MECHANICAL PARAMETERS OF A BONDED JOINT ...
DETERMINATION OF THE MECHANICAL
PARAMETERS OF A BONDED JOINT BETWEEN A
METAL AND A COMPOSITE BY COMPARING
EXPERIMENTS WITH A FINITE-ELEMENT MODEL
DOLO^ANJE MEHANSKIH PARAMETROV SPOJA MED KOVINO
IN KOMPOZITOM S PRIMERJAVO PREIZKUSOV IN MODELOM
KON^NIH ELEMENTOV
Petr Bernardin1, Josef Vacík1, Tomá{ Kroupa2, Radek Kottner3
1University of West Bohemia in Pilsen, Department of Machine Design, Univerzitní 22, 306 14 Plzeò, Czech Republic
2University of West Bohemia in Pilsen, Department of Mechanics, Univerzitní 22, 306 14, Plzeò, Czech Republic
3University of West Bohemia in Pilsen, European Centre of Excellence, NTIS – New Technologies for Information Society, Faculty of Applied
Sciences, Univerzitní 22, 306 14, Plzeò, Czech Republic
berny@kks.zcu.cz
Prejem rokopisa – received: 2012-09-01; sprejem za objavo – accepted for publication: 2012-11-16
The main goal of this work was to evaluate the application of a cohesive surface-based contact for bonded joints between
rotational parts. The numerical simulations were realized using the commercial finite-element software Abaqus. The mechanical
properties of the joint were identified using the gradient-optimization method implemented in the OptiSLang software. The
identification was performed by minimizing the difference between the force-displacement diagram obtained from the
numerical analyses and from the experiments. The shapes of the specimen and the bonded joint were designed in accordance
with joints commonly used in the machine industry.
Keywords: joint, bond, composite, metal, finite element, optimization, damage modelling, identification
Glavni cilj tega dela je oceniti uporabo spoja na osnovi kohezivnega stika povr{in med rotacijskimi deli. Numeri~ne simulacije
so bile izvr{ene z uporabo komercialne programske opreme kon~nih elementov Abaqus. Mehanske lastnosti spoja pa so bile
ugotovljene z uporabo gradientne optimizacijske metode, uporabljene s programsko opremo OptiSLang. Ugotavljanje je bilo
izvr{eno z zmanj{anjem razlike med vrednostmi na diagramu sila-raztezek, dobljenem z numeri~no analizo, in preizkusi. Oblike
vzorcev in spojev so bile izvedene skladno s spoji, ki se pogosto uporabljajo v strojni{tvu.
Klju~ne besede: stik, vez, kompozit, kovina, kon~ni element, optimizacija, modeliranje po{kodb, identifikacija
1 INTRODUCTION
Composite materials are being used in different
industrial fields. The number of applications in the
machinery industry is also increasing very quickly. Most
constructions usually consist of more than one material,
which places severe demands on the inter-part connec-
tions, for example, a metal-to-composite joint. Adhesive
bonding is a suitable option for such joints. For the
purpose of numerical simulations, the proper finite-ele-
ment model should be selected. There are many
possibilities for bonded joint modelling. The components
of the joint can be represented by 3D elements, whereas
the adhesive layer substitutes the system of spring
elements with a specific stiffness for each direction.1 The
most innovative and recently the most used approach
utilizes cohesive elements, i.e., the damage and cohesive
surface behaviour2 (stiffness and maximum nominal
stress), to define the material properties.3 These para-
meters are not usually provided by the manufacturers of
adhesives, since the parameters strongly depend on the
type of the bonded joint. The main goal of this work is
the determination of the specific values of these para-
meters on the basis of tensile tests. A special optimi-
zation cycle compares the results of the experiments and
the FE analyses and minimizes the difference.4
2 EXPERIMENT
The experiments were performed on specimens
consisting of two parts made of steel (Figure 1 – part #1
and #3) and composite material (Figure 1 – part #2). The
shape of the specimens and the bonded joint was
designed by respecting the normal joints used in the
machinery industry. The outer diameter of the composite
Materiali in tehnologije / Materials and technology 47 (2013) 4, 417–421 417
UDK 519.61/.64:669.018.95 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)417(2013)
Figure 1: Sketch of specimen
Slika 1: Skica vzorca
pipe was D = 30.32 mm and the thickness was t = 6.328
mm. The properties of the composite pipe are shown in
Table 1.
The assembly arrangement is shown in Figure 1.
The description of the layers, the fibre orientation
and the thickness of particular layers in the composite
pipe are shown in Table 2.
Table 2: Description of layers
Tabela 2: Opis plasti
Layer number Thickness (mm) Fibre Orientation (°)
1 0.677 18.79
2 0.638 –20.09
3 1.320 0.00
4 1.320 0.00
5 0.577 0.00
6 0.228 87.17
7 0.228 87.22
8 0.453 87.26
9 0.453 87.35
10 0.454 87.43
Four specimens were analyzed during the determi-
nation of the cohesive parameters. An epoxy adhesive
Spabond 345 LV5 was used to join the two cylindrical
surfaces of the specimens. The specimens were then
tensile loaded until failure using a Zwick Roell Z050
universal testing machine. The method of loading and
the location of the extensometer are shown in Figure 2.
2.1 Finite-element analysis
The finite-element models were created in the
commercial software Abaqus/CAE 6.11-1. All the parts
were modelled in accordance with the sketch in Figure
1. They were uniformly meshed except for the contact
surfaces. These surfaces were covered with a finer mesh.
Parts #1 and #3 were made of steel. Part #2 was made of
composite material. The idealization of the adhesive
layer was realized using surface-to-surface contact with
defined cohesive and damage behaviour. To simulate the
loading, the coupling function on both frontal surfaces
and the reference points was applied. The boundary
conditions were set on the reference points. The loading
was defined by the velocity vy = 0.033 mm/s. The 3-D
finite-element model is shown in Figure 3. Four parti-
cular specimens were modelled because of the different
size of the active surface of the bonded joint.
The surface-to-surface-contact with the defined
cohesive and damage properties was used for the modell-
ing. The cohesive properties of the bonded joint are
characterized by the cohesive stiffness in three directions
(knn, kss, ktt). A damage-modelling option was used for the
simulation of the degradation and the eventual failure of
the joint. The mechanism of failure consists of a
damage-initiation criterion and the damage-evolution
law.
The damage initiation refers to the beginning of the
degradation when the maximum contact stress defined
by user (tn, ts, tt) and the nominal traction stress are
equal, and the maximum contact stress ratio reaches the
value of one. This state describes the equation:2
max
t
t
t
t
t
t
n
n
s
s
t
t
0 0 0 1
⎧
⎨
⎩
⎫
⎬
⎭
= (1)
P. BERNARDIN et al.: DETERMINATION OF THE MECHANICAL PARAMETERS OF A BONDED JOINT ...
418 Materiali in tehnologije / Materials and technology 47 (2013) 4, 417–421
Figure 3: 3-D meshed model
Slika 3: 3-D model mre`e
Figure 2: Specimen testing using the Zwick Roell Z050 (Load in F
direction)
Slika 2: Preizkus vzorca z Zwick Roell Z050 (obremenitev v smeri F)
Table 1: Material properties of the composite pipe
Tabela 1: Lastnosti materiala kompozitne cevi
Matrix Vm
Epoxy resin 0.29
Fibre Vf
T700 0.71
E1/MPa E2/MPa E3/MPa
167662 6627 6627
v12 v23 v31
0.329 0.326 0.013
G12/MPa G23/MPa G31/MPa
5116 2498 5116
The damage evolution can be defined either by an
effective separation during complete failure or by the
energy that dissipates during the failure. The separation
() is described in Figure 4, while the energy (G0) is
represented by the area under the curve in Figure 4.
The relationship between the above-mentioned
parameters is described by equation:3
t
t
t
t
K K K
K K K
K K K
n
s
t
nn ns nt
ns ss st
nt st tt
=
⎧
⎨
⎪
⎩⎪
⎫
⎬
⎪
⎭⎪
=
⎡
⎣
⎢
⎢
⎤
⎦
⎥
⎥
⎧
⎨
⎪
⎩⎪
⎫
⎬
⎪
⎭⎪
=




n
s
t
K (2)
2.2 Identification
The cohesive and damage properties of the bonded
joint were identified using the optimization method
implemented in OptiSLang software. The cohesive and
damage parameters (, knn, kss, ktt, tn, ts, tt) were set as the
input parameters. The identification was performed by
minimizing the difference between the force-displace-
ment diagram obtained from the numerical analysis and
the experiment:
r
F F
Fg
FEA
i i
i
i
i
n
=
−
=
∑
( )
max( )
exp
exp
2
0
(3)
where FFEA
i is the force obtained by the experiment and
F iexp is the force obtained by the analysis. The difference
marked as rg is the objective function and was tracked at
particular points of two curves. This is the model
parameter to be minimalized.
3 DISCUSSION
The curves shown in Figure 5 describe the relation-
ship between the force and the displacement of all four
specimens. These values were obtained during the
testing. The reason for the disagreement between the
curves was the different area of the active bonding
surface. The real bonding surface area was influenced by
the shrinkage of the adhesive, which occurs within the
curing process. Another reason could be the errors
caused by the imperfect manual preparation of the test
specimens. Thanks to the particular ruptured specimens
it was possible to find the different widths of the bonded
joint surface, which was implemented in the 3D-model.
Using the optimization process, the necessary para-
meters for all the bonded joints were found for the tested
specimens (Table 3).
Table 3: Cohesive and damage parameters
Tabela 3: Parametri kohezije in po{kodb
Parameters 1 2 3 4
/mm 0.75 0.89 0.95 0.97
Knn/N/mm 1254.0 1158.0 1193.0 1194.0
Kss /N/mm 1212.0 1179.0 1112.0 1112.6
Ktt /N/mm 1000.0 6648.0 6007.0 850.9
tn /MPa 30.0 30.0 30.0 30.0
ts /MPa 30.0 30.0 30.0 30.0
tt /MPa 30.1 27.8 30.6 31.2
rg 41.19 11.7 4.1 16.8
Table 3 describes the cohesive and damage para-
meters obtained during the optimization process for four
specimens. In the ideal state, the found parameters had to
be the same, but some differences emerged, especially in
the value of Ktt. The specimens #1 and #4 show similar
values, as well as #2 and #3. This might be caused by a
different mechanism of failure during the test. With the
1st and 4th specimens, the pull-up of the fibre, not the
P. BERNARDIN et al.: DETERMINATION OF THE MECHANICAL PARAMETERS OF A BONDED JOINT ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 417–421 419
Figure 5: Comparison between four tested specimens
Slika 5: Primerjava med {tirimi preizku{enimi vzorci
Figure 4: Linear damage parameters
Slika 4: Linearni parametri po{kodbe
Figure 6: Specimens 1 and 4
Slika 6: Vzorca 1 in 4
adhesive failure, occurred. The failures of the specimens
#2 and #3 were located in the layer of the adhesive. The
properties gained from these specimens (bold text in
Table 3) are really describing the cohesive and damage
behaviour of the adhesive. These parameters do not
differ by more than 10 % and the rg ratio is also low. The
mechanism of the failure is shown in Figures 6 and 7.
The fitting of the interpolation curves of the experi-
ment and the FE analyses are shown from Figures 8 to
11.
The behaviour of the curves of the experiment and
the FE analyses differs after the 0.8 mm value of the
displacement and the curves were fitted up to this value.
After this value, the force obtained by the experiment
had a constant value, which was caused by the friction,
although the adhesive was already ruptured. The friction
was not taken into account in the FE models; therefore,
the force decreased to zero in the numerical simulations.
As is implied by the optimization process, the para-
meters of , ktt, tt have the largest influence on the shape
of the curve, while the other parameters have smaller
influences (knn, kss, tn, ts). The datasheet shear stress value
of the Spabond 345 LV5 is in the range between 29 MPa
and 37 MPa, which corresponds well with the observed
values.
4 CONCLUSION
A series of tensile tests was performed on specimens
consisting of steel and composite material. Particular
parts of the specimens were joined by the Spabond 345
LV5 adhesive and loaded using the Zwick Roell Z050
P. BERNARDIN et al.: DETERMINATION OF THE MECHANICAL PARAMETERS OF A BONDED JOINT ...
420 Materiali in tehnologije / Materials and technology 47 (2013) 4, 417–421
Figure 10: Specimen 3
Slika 10: Vzorec 3
Figure 7: Specimens 2 and 3
Slika 7: Vzorca 2 in 3
Figure 8: Specimen 1
Slika 8: Vzorec 1
Figure 11: Specimen 4
Slika 11: Vzorec 4
Figure 9: Specimen 2
Slika 9: Vzorec 2
universal testing machine. The objective function was
minimized with the help of the OptiSLang software. The
damage and cohesive parameters of the cohesive
surface-based contact were determined. The analyses
successfully confirmed the suitability of the experi-
mental data fitting for the identification of the cohesive
parameters. Nevertheless, further investigations should
follow in order to optimise the parameters, e.g., testing
according to ASTM standards.
Acknowledgement
The work has been supported by projects ME 10074
and GA P101/11/0288.
5 REFERENCES
1 F. P. Tahmasebi, Finite Element Modeling of an Adhesive in a
Bonded Joint, NASA Goggard Space Flight Center, 1999
2 M. L. Benzeggagh, M. Kenane, Measurement of Mixed-Mode
Delamination Fracture Toughness of Unidirectional Glass/Epoxy
Composites with Mixed-Mode Bending Apparatus, Composites
Science and Technology, 56 (1996) 4, 439–449
3 P. P. Camanho, C. G. Davila, Mixed-Mode Decohesion Finite Ele-
ments for the Simulation of Delamination in Composite Materials,
NASA/TM-2002-211737, 2002
4 P. Janda, M. Kosnar, T. Kroupa, V. La{ová, J. Vacík, Vyu`ití
programù ANSYS a OPTISLANG v konstrukci výrobních strojù,
Frymburk: ANSYS konference, 2010
5 Spabond, Epoxy Adhesive System, Newport Isle of Weight, 2004,
Available: http://www.gurit.com/files/documents/Spabond_345_v7.
pdf. [cited 12 June 2012]
P. BERNARDIN et al.: DETERMINATION OF THE MECHANICAL PARAMETERS OF A BONDED JOINT ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 417–421 421

T. SKALAR et al.: DETERMINATION OF THE METAL CONCENTRATIONS IN AN ANODE MATERIAL ...
DETERMINATION OF THE METAL CONCENTRATIONS
IN AN ANODE MATERIAL FOR SOLID-OXIDE FUEL
CELLS
DOLO^ITEV VSEBNOSTI KOVIN V ANODNEM MATERIALU ZA
GORIVNE CELICE S TRDNIM ELEKTROLITOM
Tina Skalar1,2, Amalija Golobi~1, Marjan Marin{ek1,2, Jadran Ma~ek1,2
1University of Ljubljana, Faculty of Chemistry and Chemical Technology, A{ker~eva 5, 1000 Ljubljana, Slovenia
2Center of Excellence for Low-Carbon Technologies (CO-NOT), Hajdrihova 19, 1000 Ljubljana, Slovenia
tina.skalar@fkkt.uni-lj.si
Prejem rokopisa – received: 2012-09-25; sprejem za objavo – accepted for publication: 2013-01-03
Ni-SDC is a very promising new anode material for SOFC systems. It exhibits a superb ionic and electronic conductivity at
intermediate temperatures (400–700 °C) in comparison to Ni-YSZ. Ni-SDC is a composite material requiring careful
microstructural and compositional tailoring during the preparation of the material. Ni-SDC was synthesized using a simplified
Pechini method with a reaction between metal acetates and ethylene glycol, the latter serving as a chelating agent, and a reaction
medium. The molar ratio between cerium and samarium in the ceramic part was 83 : 17 and the overall nickel content in the
final product was set to 38 %. During the preparation of the material, several different intermediates were synthesized (i.e., a
powdered product after the synthesis, an oxide mixture after the calcination and a mechanically ground powder after the
milling), which may differ according to their chemical and morphological properties. In this respect, the material chemical
composition expressed as concentrations of nickel, cerium and samarium was followed through the preparation sequence using
various analytical techniques, i.e., the volumetric and gravimetric methods, ICP-OES, SEM-EDS and XRD. It appears that the
obtained results diverge consistently with the analytical techniques used. Volumetry and gravimetry were used only for the
nickel-content determination. Additionally, all the metals were simultaneously determined with ICP-OES and XRD Rietveld
refinement in a bulk sample and with SEM-EDS for a point analysis. There is no unique answer as to which analytical method
should be used at various preparation steps. Instead, suitable analytical methods or their combinations were chosen with respect
to the analyzed material’s appearance and its morphological and microstructural characteristics.
Keywords: NiO-SDC powders, gravimetric method, volumetric method, SEM-EDS, XRD, ICP analysis
Kermet Ni-SDC se uporablja kot anodni material v SOFC gorivnih celicah (gorivne celice s trdnim elektrolitom). V primerjavi z
navadno uporabljenim anodnim materialom Ni-YSZ ima Ni-SDC vi{jo ionsko in elektronsko prevodnost pri temperaturah od
400 °C do 700 °C. Kermet Ni-SDC zahteva pazljivo prilagajanje mikrostrukture in sestave med svojo pripravo. Ni-SDC je bil
sintetiziran s poenostavljeno Pechinijevo metodo, kjer poteka reakcija med kovinskimi acetati in etilen glikolom. V kon~nem
produktu je bilo molsko razmerje med cerijem in samarijem 83 : 17, volumenski dele` niklja pa 38 %. Za ustrezen anodni
material je potrebnih ve~ stopenj obdelave, sinteza prekurzorja Ni-SDC, kalcinacija in mletje v atritorju. V vsaki stopnji se pri
materialu lahko spreminjata ali sestava ali morfologija. Kemijska sestava vseh treh materialov je bila izra`ena s koncentracijami
niklja, cerija in samarija. Za dolo~itev teh koncentracij so bile uporabljene razli~ne analizne metode: volumetrija, gravimetrija,
ICP-OES, SEM-EDS in XRD. Med dobljenimi rezultati se glede na uporabljeno metodo poka`ejo razlike. Volumetrijo in
gravimetrijo smo uporabili za dolo~anje koncentracije niklja v vzorcih. Vsebnost vseh treh kovin isto~asno pa smo dolo~ili z
ICP-OES, XRD in SEM-EDS, kjer se s slednjo to~kovno dolo~i koncentracije, s prvima dvema metodama pa v ve~jem
volumnu. Unikatnega odgovora, katera analitska metoda je najbolj primerna, ni. Najprimernej{o metodo oziroma njihovo
kombinacijo izberemo glede na morfolo{ke in mikrostrukturne lastnosti analiziranega materiala.
Klju~ne besede: NiO-SDC-prahovi, gravimetri~na metoda, volumetri~na metoda, SEM-EDS, XRD, ICP-analiza
1 INTRODUCTION
The solid-oxide fuel cell (SOFC) is one source of
alternative energy. These cells have a more efficient and
pollution-free transformation of fuel into electrical
power than the traditional combustion engines, which is
the reason for the growing research in this field. A SOFC
is typically composed of an anode, a cathode and elec-
trolytes, which should each have their specific proper-
ties. The anode should have a high ionic and electronic
conductivity for the transportation of oxide ions to the
reaction site and electrons from it. Its microstructure
should provide a good contact between the reactive gas,
electronic and ionic conductors (the triple-phase boun-
dary; hereafter: TPB). Finally, the desired molar ratio of
the metals in the cermet1,2 is also very important. The
volume content of nickel in the cermet influences the
open porosity of the material, which enables a contact
between the gasses and the anode.3 The electronic con-
ductivity increases with the increasing nickel content;
meanwhile, the ionic conductivities of the cermet de-
crease due to the weak connections among the SDC
grains.4 Furthermore, the ceramic phase of the cermet
(samaria-doped ceria; hereafter: SDC) should have a
proper composition of samarium and cerium. It is experi-
mentally determined that the best ionic conductivity is
achieved when the molar ratio of Ce : Sm is 80 : 20, and
the samarium is distributed homogeneously in the ce-
rium5. It is necessary to fulfill all these conditions to
provide a high-efficiency material for SOFCs . Previous
contributions6–8 mostly stress the importance of the metal
content in such materials, but this was rarely determined,
Materiali in tehnologije / Materials and technology 47 (2013) 4, 423–429 423
UDK 621.35/.36:621.352.6:666.798.2 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)423(2013)
even partially. The value of the metal content has the
highest importance when developing the materials for
SOFCs.
This report describes the material’s composition
expressed as concentrations of nickel, cerium and sama-
rium, which are followed through the preparation steps,
using several analytical techniques, such as the volume-
tric and gravimetric methods, ICP-OES, SEM-EDS and
XRD.
2 EXPERIMENTAL WORK
2.1 Sample preparation
A mixture of 17.68 g of Ce(C2H3O2)3 . xH2O (Sigma
Aldrich, 99.9 % pure, metals basis) and 3.65 g of
Sm(C2H3O2)3 . xH2O (Sigma Aldrich 99.9 % pure, metals
basis) was dissolved in 150 mL of deionized water. After
the acetates dissolved 100 mL of ethylene glycol
(C2H6O2, Sigma Aldrich, puriss. p. a. 99.5 %) was added.
The mixture was heated to 50 °C for 30 min, then the
temperature was raised to 80 °C and 32.93 g of nickel
acetate Ni (CH3COO)2.4H2O (Riedel-de Haën, min 98 %
purity) was added. The metal molar ratio in the reaction
mixture was 5.0 : 4.6 : 1. This mixture was heated and
vacuum dried for two and half hours in a vacuum dryer
(Büchi Heating bath B-490 Rotavapor R-200), equipped
with a vacuum pump (PC 2003 VARIO). Afterwards, the
product was dried at 80 °C for three days until the final
crystals were formed, which were then used as an
intermediate (Sample 1) to produce the anode material.
This intermediate was calcined at 900 °C in air for 1 h to
form a mixture of metal oxides (Sample 2). Sample 2
was homogenized in an atritor mill by wet milling in
isopropanol for one hour which resulted in Sample 3.
Nickel concentrations in the NiO-SDC material were
determined with the gravimetric method, using dimethyl-
glyoxime, and the volumetric method with the EDTA
titration. To determine cerium and samarium, and the
nickel weight content, SEM-EDS, XRD method and
ICP-OES analysis were used.
The samples that were used for gravimetry, volu-
metry and the ICP analysis were dissolved in different
solvents. The intermediate after the synthesis (Sample
1), which was a water-soluble product, was dissolved
with distilled water. Sample 2 was a cermet composed of
nickel oxide, samarium oxide and cerium oxide. It was
dissolved in concentrated (96–98 %) sulfuric (VI) acid
and evaporated to form the sulfates of the metals present.
The residue of metallic sulfates was then dissolved in 37
% hydrochloric acid and diluted in distilled water.
Sample 3 was dissolved using the same acid procedure
as for the Sample 2.
2.2 Analytical procedures
2.2.1 Thermal decomposition of Sample 1
Thermal decomposition of Sample 1 was analyzed in
the range from room temperature to 1000 °C in air with a
heat rate of 10 K min–1 with the TG/DTA analysis using
a Netzsch STA 449 F3 Jupiter apparatus.
2.2.2 Gravimetric determination with dimethylglyoxime
The nickel content was determined in 10 mL aliquots
of dissolved samples diluted with distilled water in an
Erlenmeyer flask. The pH level was adjusted to between
4 to 5 with an ammonia or hydrochloric solution,
depending on the pH value of the previous solution.
Then, 0.5 mL of NaCOOCH3 and 20 mL of CH3COOH
were added to form an acetate buffer. The solution was
then boiled with an electric heater. After the mixture
cooled down slightly, 10 mL of dimethylglyoxime (1 %
ethanol solution) was added and the red nickel preci-
pitate was filtered through the ceramic filter and washed
with hot water. The precipitate containing nickel was
dried at 120 °C for 1 h, cooled down in a desiccator and
weighed. This was repeated at least three times. The
weight percentage was calculated from the weight of the
complex.9 It is recommended that the analysis of the
standard sample be carried out in the same method of
analysis that was used for our samples. Therefore, the
gravimetric method was also used for the nickel (II)
oxide (NiO), the Aldrich nanopowder that was taken as
the standard material.
2.2.3 Volumetric method using EDTA as a titration
reagent
For a determination of nickel, 10 mL aliquots of
dissolved samples were diluted with distilled water in an
Erlenmeyer flask to 100 mL. The pH value was adjusted
to 10 with an ammonia solution. Then a solid murexide
indicator was added that formed a bright yellow color.
The sample was titrated with the standard 0.01 M EDTA
until the yellow color of the nickel-murexide complex
started to change to the purple color of a free indicator.
The volumetric-determination procedure was repeated
ten times for each sample. The weight percentage was
calculated from the EDTA volume used for the titration
of the nickel ions.9–11 Using this method, the nickel
content was also determined for the nickel (II) oxide
(NiO), the Aldrich nanopowder that was taken as the
standard material.
2.2.4 SEM-EDS
All the samples were characterized on an FE-SEM
Zeiss Ultra Plus microscope equipped with EDS (an
Oxford X-Max SDD 50 mm2 detector and INCA 4.14
X-ray microanalysis software). The sample preparation
included a fixation onto a conductive C tape and a sub-
sequent sputtering with platinum, without any polishing.
The detector was calibrated just before the analysis with
the Co-standard under operating conditions. The EDS
spectra were recorded on the flat regions of the samples
using a process time of 5, a lifetime of 120 s and an
accelerating voltage of 14 kV, which is an acceptable
compromise between the analyzing volume and the
overvoltage needed to excitate the Ce, Sm and Ni X-rays,
T. SKALAR et al.: DETERMINATION OF THE METAL CONCENTRATIONS IN AN ANODE MATERIAL ...
424 Materiali in tehnologije / Materials and technology 47 (2013) 4, 423–429
whose emission lines are found in the interval between
4.84 (L1) keV, 5.64 (L1) keV and 7.48 (K1) keV,
respectively. Using the Anderson-Halser estimation,12 the
X-ray production depth was approximately 0.6 μm. The
quantification of the X-ray spectra was performed with
respect to the standard procedure provided by the soft-
ware manufacturer (ZAF-based method13). For the stati-
stically reliable data in each case, five to seven different
fields of view in various regions of interest were ana-
lyzed.
2.2.5 XRD analysis
The X-ray powder-diffraction data for Samples 1, 2
and 3 were collected using a PANalytical X’Pert PRO
MPD diffractometer with the –2 reflection geometry, a
primary-side Johansson-type monochromator and the
CuK1 ( = 0.154 059 nm) radiation. The room-tempe-
rature-reflection data were acquired from the 2 angles
of 5° to 90° in the steps of 0.034°. A quantitative XRD
analysis with the Rietveld method using the TOPAS2.1
program suite14 was performed for Samples 2 and 3. The
background was modeled with a third-order polynomial.
We also refined zero error, scale factor, lattice parameter
a and one profile parameter (the crystallite size) for NiO
and SDC. The atom parameters were fixed, taken from
the published structures of SDC and NiO.15,16 The final
match between the observed and calculated profiles is
shown in Figures 1 and 2. The agreement factor Rwp was
0.094 5 and 0.078 5 for Samples 2 and 3, respectively.
2.2.6 ICP-OES
An analysis of nickel, samarium and cerium was
carried out with an inductively coupled plasma-optical
emission spectrometer (Varian 715-ES ICP-OES
Spectrometer). The ICP-OES spectrometer is used to
define concentrations of several elements in the solution.
The ICP-OES method was already used to detect the
metals (Ni, Ce, Gd, Ag) in the materials used for
SOFCs.1,17–20 Details about the instruments’ operating
conditions are depicted in Table 1.
3 RESULTS AND DISCUSSION
The results for the metal contents in Samples 1, 2 and
3 are shown in Tables 2, 3 and 4, respectively. The
calculated concentrations of cations for Samples 2 and 3
were determined according to the initial amounts of
added cations. It was assumed that during the process
there were no losses of the metals and that the entire che-
mical reaction took place in line with the expectations,
meaning that Samples 2 and 3 consist of oxides only.
The calculated metal mass fractions for Sample 1 were
calculated by considering its thermal decomposition into
oxides (Figure 3). The concentration range of the ele-
ments that are in the samples is in the detection range of
the classical analytical methods (gravimetry, volumetry).
The systematic error in the gravimetric determination of
nickel had to do with the specific characteristics of the
precipitated nickel dimethylglyoxime (the voluminosity
of the formed complex).21 The values obtained with the
complexometric titration of nickel deviate more from the
expected calculated values than in the case of gravimetry.
It was difficult to determine the color leap of the organic
indicator that provided the endpoint of titration. The pH
value of the prepared solution had a significant influence
on the results, as did the absence of oxidants in the
analyte.21 The rare elements showed a great similarity in
the chemical properties; therefore, the concentration of
these elements (cerium and samarium) was determined
only with instrumental methods.
Determination of the chemical compositions of the
samples was quite a challenge with regard to the quanti-
fication of the collected EDS spectra. The main difficulty
was found in the fact that the Ce and Sm characteristic
X-ray peaks in the EDS spectra mostly overlap. This
problem was particularly pronounced in the case of Sm,
since the superimposed Ce peaks on all the main Sm
peaks made a consistent quantitative compositional cal-
culation rather difficult. From this point of view, the
related WDS analysis seems to be more suitable for a
compositional investigation of such samples. However,
when compared to WDS, EDS has its advantages: it is
more available, quicker, simpler and easier to perform. In
fact, in material science, the EDS analysis is one of the
basic tools for compositional investigations and, as such,
it is very popular.
EDS and XRD are nondestructive approaches to a
material investigation, but both are less accurate than the
other three used analytical methods. Additionally, EDS
and XRD are both based on a similar physical pheno-
menon; however, there is one distinguishable difference
between them. The XRD quantitative analysis always
refers to the bulk-sample composition, while EDS
denotes the chemical composition in a much smaller
sample region. For this reason, EDS results always
depend upon the specimen surface preparation and are
also influenced by the sample topography.
From this point of view, the three investigated sam-
ples were somewhat dissimilar. After the synthesis, the
sample was composed of tiny crystals (Figure 4). To
determine the homogeneity of this sample, an X-ray sur-
T. SKALAR et al.: DETERMINATION OF THE METAL CONCENTRATIONS IN AN ANODE MATERIAL ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 423–429 425
Figure 1: Rietveld refinement for Sample 2: experimental and calcu-
lated (upper curves) and difference (lower curve) profiles. Vertical
bars denote the position of reflections for SDC (upper) and NiO
(lower).
Slika 1: Rietveldovo prilagajanje za vzorec 2: izmerjena in izra~unana
krivulja (zgoraj) ter diferen~na (spodaj). Pokon~ne ~rtice ozna~ujejo
lego uklonov za SDC (zgornje) in NiO (spodnje).
face mapping was performed (Figure 5). As expected,
Ce, Sm, Ni, O and C were the only elements detected in
the sample. From the element maps it could be seen that
the specimen was rather inhomogeneous. Some vari-
ations in the grayscale maps may be associated with the
specimen topography; however, it was evident that the
sample contains two separate phases. The elongated
crystals, typically 20–40 μm long and 5 μm wide, were
Ce- and Sm-rich, while smaller particles (3–10 μm in
diameter) of an undefined shape were Ni-rich. This was
also supported by the XRD diffraction pattern of Sample
1, presented in Figure 6. The qualitative XRD analysis
revealed that the sample consisted of an SDC metal-
organic precursor with a known structure and formula22
and of at least one additional crystalline phase with an
unknown composition and structure. The positions of the
first reflections of the unknown phase were at the angles
lower than 10° 2. They also had significantly lower
intensities in comparison with the SDC precursor phase.
Both facts indicated that the unknown phase could be a
metal-organic nickel complex, i.e., a precursor of NiO.
The XRD quantitative analysis using Rietveld refinement
was not possible for this sample, since it requires the
knowledge of the composition and the structure of all the
crystalline phases. Due to the observed inhomogeneity,
the only practical way of performing a quantitative EDS
analysis was by collecting the X-ray signals from the
regions of interest (ROI) of approximately 0.25 mm2 in
size (6–7 ROIs on each field of view). According to the
results summarized in Table 2, the determined element-
content ratio (ECR) among Ni, Ce and Sm (w(Ni):w(Ce):
w(Sm) = 5.0 : 5.1 : 1.0) was not in good accordance with
T. SKALAR et al.: DETERMINATION OF THE METAL CONCENTRATIONS IN AN ANODE MATERIAL ...
426 Materiali in tehnologije / Materials and technology 47 (2013) 4, 423–429
Figure 3: Thermal decomposition of Sample 1 to Sample 2 in air
Slika 3: Termi~ni razpad vzorca 1 do vzorca 2 na zraku
Figure 4: Tiny crystals in Sample 1 after the synthesis
Slika 4: Kristal~ki vzorca 1 po sintezi
Table 1: Operating conditions for the ICP-OES analysis
Tabela 1: Pogoji merjenja za ICP-OES analize
Parameter Value
RF generator power / kW 1.2
Frequency of RF generator / MHz 40
Plasma-gas flow rate / L min–1 15.0
Auxiliary-gas flow rate / L min–1 1.50
Nebulization-gas flow pressure / kPa 200
Gas Argon
Sample-uptake rate / mL min–1 1.9
Type of detector CCD
Type of spray chamber Sturman-Mastersdouble pass
Type of nebulizer V-groove
Element /nm–1
Ce: 418.659
446.021
Sm: 359.259
360.949
Ni: 216.555
231.604
Table 2: Results of metal mass fractions for Sample 1 obtained with
different analytical methods
Tabela 2: Masni dele`i treh kovin v vzorcu 1, dolo~eni z razli~nimi
analitskimi metodami
Used method w(Ni)/% w(Ce)/% w(Sm)/%
w(Ni) : w(Ce)
: w(Sm)
Calculated value 16.1 14.6 3.2 5.0 : 4.6 : 1.0
Gravimetric
method
16.5 ±
0.3 / / /
Volumetric
method
15.47 ±
0.02 / / /
SEM-EDS 12.3 ±0.5
12.6 ±
1.5 2.5 ± 0.5 5.0 : 5.1 : 1.0
ICP-OES 14.9 13.5 2.8 5.4 : 4.9 : 1.0
w/% – mass fraction, w/% – masni dele`
Figure 2: Rietveld refinement for Sample 3: experimental and calcu-
lated (upper curves) and difference (lower curve) profiles. Vertical
bars denote the position of reflections for SDC (upper) and NiO
(lower).
Slika 2: Rietveldovo prilagajanje za vzorec 3: izmerjena in izra~unana
krivulja (zgoraj) ter diferen~na (spodaj). Pokon~ne ~rtice ozna~ujejo
lego uklonov za SDC (zgornje) in NiO (spodnje).
the calculated ratio w(Ni) : w(Ce) : w(Sm) = 5.0 : 4.6 :
1.0. The highest deviations were observed for the
Ce-content determination, which was estimated to be too
high and the Sm-content determination, which was esti-
mated to be too low. The reason for the inaccurate Ce-
and Sm-content determinations was probably a poor
deconvolution of the overlapping peaks in the EDS
spectra.
The EDS spectra of Sample 1 also exhibited a cha-
racteristic carbon peak, which made the absolute quanti-
tative analysis unrealistic. The C peak was attributed to
two origins; i) carbon was contained in the structure of
the metal-organic intermediate in Sample 1, and ii) the
characteristic C peak was partially a consequence of a
sample contamination due to its exposure to air (CO2)
before the analysis, as well as some cracking of the
hydrocarbons in the vacuum system of the electron
microscope under the electron beam.23
Sample 2 was obtained with the subsequent thermal
treatment of Sample 1 without performing any other
mechanical operation. Such a preparation path for
Sample 2 also meant that the ECR for the metals should
be the same as for Sample 1. However, the absolute
values of the Ni, Ce, Sm and O element contents should
increase during the thermal treatment due to a mass loss
when the metal-organic compound in Sample 1 was
transformed into an oxide mixture of NiO and SDC.
After the thermal decomposition, the sample retained
its degree of homogeneity (Figure 7). For that reason,
the X-ray signals for the quantitative EDS analysis of
Sample 2 were repeatedly collected from ROIs of appro-
ximately 0.25 mm2 in size (6–7 regions of interest on
each field of view), similarly to the EDS analysis of
Sample 1. Since the treatment at 900 °C transformed the
sample into a mixture of oxides, the summation of the
Ni, Ce, Sm and O contents was normalized to 100 %.
The EDS spectra of Sample 2 also exhibited a charac-
teristic C peak, which was attributed only to the sample
contamination and subsequently entirely omitted in the
quantitative calculations. The average absolute values for
the Ni, Ce and Sm contents as obtained by the EDS
measurements are given in Table 3. The errors were
estimated on the basis of the variations of the data. The
results of the quantitative EDS analysis of the Ni, Ce and
Sm contents (38.7 %, 39.7 % and 7.0 %, respectively)
were in a relatively good agreement with the expected
calculated values for Sample 2. However, due to the local
inhomogeneity, the measured absolute values of the
element contents may differ substantially between ROIs,
resulting in a relatively high standard deviation (± 4.8 %
for Ni). The calculated ECR for Sample 2 implied that
the Sm content was again estimated too low.
The quantitative XRD phase analysis with the
Rietveld method resulted in (48.3 ± 3.0) % of SDC and
T. SKALAR et al.: DETERMINATION OF THE METAL CONCENTRATIONS IN AN ANODE MATERIAL ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 423–429 427
Figure 5: X-ray surface mapping of Sample 1
Slika 5: Elementna porazdelitev vzorca 1
Table 3: Results of metal mass fractions for Sample 2 obtained with
different analytical methods
Tabela 3: Masni dele`i treh kovin v vzorcu 2, dolo~eni z razli~nimi
analitskimi metodami
Used method w(Ni)/% w(Ce)/% w(Sm)/%
w(Ni) : w(Ce)
: w(Sm)
Calculated value 38.2 34.7 7.6 5.0 : 4.6 : 1.0
Gravimetric
method
38.58 ±
0.09 / / /
Volumetric
method
36.8 ±
0.2 / / /
SEM-EDS 38.7 ±4.8
39.7 ±
4.1 7.0 ± 1.5 5.5 : 5.7 : 1.0
XRD 40.6 ±3.0
32.9 ±
3.0 7.2 ± 2.0 5.6 : 4.6 : 1.0
ICP-OES 37.4 36.5 7.6 4.9 : 4.8 : 1.0
w/% – mass fraction, w/% – masni dele`
Table 4: Results of metal mass fractions for Sample 3 obtained with
different analytical methods
Tabela 4: Masni dele`i treh kovin v vzorcu 3, dolo~eni z razli~nimi
analitskimi metodami
Used method w(Ni)/% w(Ce)/% w(Sm)/%
w(Ni) : w(Ce)
: w(Sm)
Calculated value 38.2 34.7 7.6 5.0 : 4.6 : 1.0
Gravimetric
method
38.5 ±
0.2 / / /
Volumetric
method
36.45 ±
0.09 / / /
SEM-EDS 37.6 ±2.3
38.9 ±
2.1 7.0 ± 0.7 5.4 : 5.6 : 1.0
XRD 40.3 ±2.0
33.1 ±
2.0 7.3 ± 2.0 5.5 : 4.5 : 1.0
ICP-OES 38.0 36.7 7.5 5.1 : 4.9 : 1.0
w/% – mass fraction, w/% – masni dele`
(51.7 ± 3.0) % of NiO. From these results, the mass
fractions of the metals, given in Table 3, are calculated;
they are close to the expected calculated values (within
the experimental error). Sample 3 was prepared by
attritor milling of Sample 2. Such milling results in an
increased homogeneity of the sample. The region of the
one-phase dominance (either a NiO or SDC phase) was
estimated on the basis of the morphological characte-
ristics of Sample 3 and expected to be in the sub-micro-
meter range. Collecting the X-ray spectra of Sample 3
and the subsequent calculations were performed iden-
tically to the previous analyses by investigating several
ROIs (C was omitted in the quantitative calculations).
According to Table 4, the consequence of the increased
homogeneity in Sample 3 was reflected through a
relatively accurate determination of the Ni and Ce
contents with a much smaller deviation, practically
halved in comparison to Sample 2. The key problem of
processing the EDS spectra of the selected system still
remains in the fact that the main Sm peaks overlap with
the Ce peaks, resulting again in a too low Sm-content
and a too high Ce-content estimation and thus making
the EDS quantification heavily dependent on the method
used to deconvolute the spectrum. The quantitative XRD
phase analysis using the Rietveld method of Sample 3
resulted in (48.7 ± 2.0) % of SDC and (51.3 ± 2.0) % of
NiO. From these results, the mass fractions of the metals,
given in Table 4, were calculated, which were close to
the expected calculated values (within the experimental
error). It has to be emphasized that the Rietveld method
gave us the mass fractions for NiO and SDC in the cry-
stalline part of the sample. Consequently, in the case of
the presence of a significant amount of an amorphous
phase the resulted mass fractions may not be represen-
tative for the whole sample (especially if the metal
content in the amorphous phase differs significantly in
comparison to the crystalline part). Rietveld-refinement
results also confirmed that Samples 2 and 3 contained
only NiO and SDC with the desired compositions where
the lattice parameter a for SDC is 0.543 2(1) and 0.543
3(1) nm,22 respectively. The width of the reflections of
Sample 3 was larger in comparison with those of Sample
2. Consequently, the crystallite-size parameters resulting
from Rietveld refinement for both NiO and SDC were
larger in Sample 2 than in Sample 3. This was evidently
caused by the milling in the atritor. On the other hand,
the milling of Sample 3 improved the fit between the cal-
culated and measured curves (Figures 1 and 2) due to
the grinding of larger grains.
The results of the ICP analysis depend on the treat-
ment of the material before the analysis. This includes
homogenization, weighing, dissolving and diluting.
Consequently, if the element contents in the samples
submitted to the ICP analysis were higher than the
recommended concentrations, multiple diluting was
necessary, which eventually caused a deviation from the
calculated values. Furthermore, the standard deviation of
the ICP results for the analyzed samples could not be
defined due to an insufficient number of the analysis
repeats. Such repeats of the ICP analyses would have
increased the costs.
4 CONCLUSION
Based on various analytical approaches to the ana-
lyzed samples used in the process of the final-material
development, the following conclusions can be drawn:
The gravimetry and volumetry methods are rather
time-consuming showing fairly accurate values in the
chosen concentration range; however, both methods can
be used only for determining the Ni content. The gravi-
metry shows better results.
Although the topographies of the prepared samples
were not ideal for an X-ray microanalysis, one of the
T. SKALAR et al.: DETERMINATION OF THE METAL CONCENTRATIONS IN AN ANODE MATERIAL ...
428 Materiali in tehnologije / Materials and technology 47 (2013) 4, 423–429
Figure 7: X-ray surface mapping of Sample 2
Slika 7: Elementna porazdelitev vzorca 2
Figure 6: XRD patterns at low angles of Sample 1 (upper) and of pure
precursor of SDC (bottom)
Slika 6: Rentgenski pra{kovni posnetek pri nizkih kotih: vzorec 1
(zgoraj), ~isti prekurzor SDC-ja (spodaj)
aims of this work was to demonstrate the feasibility of
the quantitative EDS analysis of the selected system.
While the Ni content could be determined accurately, the
Sm content was generally estimated too low and the Ce
content too high due to the overlapping of the Ce and Sm
peaks, making the EDS quantification heavily dependent
on the method used for deconvoluting the spectrum.
The XRD quantitative analysis using the Rietveld
method gave us, in a fast, easy and relatively inexpensive
way, the mass contents for all three metals that are in
agreement with the expected calculated values within the
experimental error. The standard deviations are com-
parable to those of SEM-EDS and higher in comparison
with the volumetry or gravimetry. The advantage of this
method was also that the effect of milling on the
crystallite-size parameter could be observed and the
qualitative-phase analysis checked. The disadvantages of
this method are that it requires the knowledge of the
structures of all the present crystalline phases and that
the obtained mass contents may not be representative for
the whole sample, when a significant amount of an
amorphous phase is present in the sample.
The instrumental ICP-OES method may be consi-
dered to be accurate for all three samples. However, the
possibility of an analytical error becomes greater if the
concentration of the components is higher than 1 % due
to the necessary sample dilution. In addition, ICP-OES is
a rather costly method.
In general, to reliably determine the contents of all
three metals, we recommend a combination of gravi-
metry with the alternative instrumental methods (ICP,
SEM-EDS and XRD) that also determine cerium and
samarium.
Acknowledgement
We thank the Ministry of Education, Science, Culture
and Sport of the Republic of Slovenia for the financial
support provided through grants P1–0175-103,
P1-0A5-103 and the Center of Excellence for Low-
Carbon Technologies (CO NOT).
5 REFERENCES
1 H. Uchida, S. Suzuki, M. Watanabe, Electrochemical and Solid State
Letters, 6 (2003), A174–A177
2 M. Backhus-Ricoult, Solid State Sciences, 10 (2008), 670–688
3 M. Chen, B. H. Kim, Q. Xu, B. G. Ahn, D. P. Huang, Solid State
Ionics, 181 (2010), 1119–1124
4 Z. Wang, M. Mori, T. Itoh, ECS Transactions, 35 (2011), 1631–1640
5 H. Yoshida, H. Deguchi, M. Kawano, K. Hashino, T. Inagaki, H.
Ijichi, M. Horiuchi, K. Kawahara, S. Suda, Solid State Ionics, 178
(2007), 399–405
6 J. B. Wang, J. C. Jang, T. J. Huang, Journal of Power Sources, 122
(2003), 122–131
7 R. Maric, S. Ohara, T. Fukui, T. Inagaki, J. Fujita, Electrochemical
and Solid-State Letters, 1 (1998), 201–203
8 Y. Okawa, Y. Hirata, Journal of the European Ceramic Society, 25
(2005), 473–480
9 I. M. Kolthoff, P. J. Elving, Treatise on analytical chemistry, Part II,
Analytical chemistry of the elements, Vol. 2, Ga-In-Tl, Si, Ge, Fe,
Co, Ni, John Willey & Sons, New York- London 1962, 377–440
10 E. Merck, Complexometric Assay Methods with Titriplex®, 3rd ed.,
Darmstadt 1979, 46–47
11 H. A. Flaschka, EDTA titrations: An introduction to theory and prac-
tice, Pergamon Press, New York 1959
12 J. J. Friel, C. E. Lyman, Microscopy and Microanalysis, 12 (2006),
2–25
13 J. Goldstein, Scanning Electron Microscopy and X-Ray Microanaly-
sis, Springer, New York 2003
14 Topas2.1: Advanced X-Ray solutions, Bruker AXS, Karlsruhe,
Germany 2003
15 G. Brauer, H. Gradinger, Zeitschrift fuer Anorganische und Allge-
meine Chemie, 276 (1954), 209–226
16 N. G. Schmahl, J. Barthel, G. F. Eikerling, Zeitschrift fuer Anorga-
nische und Allgemeine Chemie, 332 (1964), 230–237
17 A. J. Darbandi, T. Enz, H. Hahn, Solid State Ionics, 180 (2009),
424–430
18 V. A. C. Haanappel, J. Mertens, J. Malzbender, Short communica-
tion, Journal of Power Sources, 171 (2007), 789–792
19 M. Stanislowski, J. Froitzheim, L. Niewolak, W. J. Quadakkers, K.
Hilpert, T. Markus, L. Singheiser, Journal of Power Sources, 164
(2007), 578–589
20 L. Saravanan, A. Pandurangan, R. Jayavel, Materials Letters, 66
(2012), 343–345
21 D. A. Skoog, D. M. West, E. J. Holler, S. R. Crouch, Fundamentals
of Analytical Chemistry, 8th ed., Brooks/Cole, USA 2004
22 T. Skalar, J. Ma~ek, A. Golobi~, Journal of the European Ceramic
Society, 32 (2012), 2333–2339
23 D. Lau, A. E. Hughes, T. H. Muster, T. J. Davis, A. M. Glenn, Micro-
scopy and Microanalysis, 16 (2010), 13–20
T. SKALAR et al.: DETERMINATION OF THE METAL CONCENTRATIONS IN AN ANODE MATERIAL ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 423–429 429

F. E. BASTAN, Y. Y. ÖZBEK: PRODUCING ANTIBACTERIAL SILVER-DOPED HYDROXYAPATITE POWDERS ...
PRODUCING ANTIBACTERIAL SILVER-DOPED
HYDROXYAPATITE POWDERS WITH CHEMICAL
PRECIPITATION AND RESHAPING IN A SPRAY DRYER
IZDELAVA S SREBROM DOPIRANEGA PROTIBAKTERIJSKEGA
PRAHU HIDROKSIAPATITA S KEMIJSKIM IZLO^ANJEM IN
PREOBLIKOVANJEM V RAZPR[ILNEM SU[ILNIKU
Fatih Erdem Baþtan, Yýldýz Yaralý Özbek
Sakarya University, Engineering Faculty, Department of Metallurgy and Materials Engineering, 54187 Esentepe-Sakarya, Turkey
febastan@sakarya.edu.tr
Prejem rokopisa – received: 2012-09-28; sprejem za objavo – accepted for publication: 2012-12-18
Hydroxyapatite is used for fixing orthopedic prostheses and it has reached a significant level of clinical application. It is
important to prevent the initial bacterial colonization of the existing colonies. Silver has been known to exhibit a strong
cytotoxicity towards a broad range of microorganisms. In this study, the aim was to produce antibacterial silver-doped
hydroxyapatite powders and to reshape them in a spray dryer. Mole fraction of silver 1 % was added to the hydroxyapatite
structure with the ion-exchange technique. E. coli bacteria were used for investigating the powder antibacterial specimens.
Different temperatures and a pressure of 1.5 bar were used for shaping them in a spray dryer. Scanning electron microscopy
(SEM), X-ray diffraction (XRD), Energy-dispersive X-ray spectroscopy (EDX) and a bacterial test were used to characterize
the powder specimens.
Keywords: hydroxyapatite, spray dryer, silver-doped hydroxyapatite, antibacterial HAP
Hidroksiapatit se uporablja za pritrditev ortopedskih protez in se mno`i~no uporablja na klinikah. Pomembno je prepre~iti
za~etno bakterijsko kolonizacijo obstoje~ih kolonij. Za srebro je poznano, da izkazuje mo~no citotoksi~nost v {irokem spektru
mikroorganizmov. Namen te {tudije je bila izdelava protibakterijskega, s srebrom dopiranega prahu iz hidroksiapatita in njegovo
preoblikovanje v pr{ilnem su{ilniku. Molski dele` srebra 1 % je bil dodan v strukturo hidroksiapatita z metodo izmenjave ionov.
Za preiskavo protibakterijskega prahu je bila uporabljena bakterija E. coli. Za preoblikovanje v pr{ilnem su{ilniku so bile
uporabljene razli~ne temperature in tlak 1,5 bar. Za karakterizacijo vzorcev prahu so bile uporabljene: vrsti~na elektronska
mikroskopija (SEM), rentgenska difrakcija (XRD), energijsko disperzijska rentgenska spektroskopija (EDX) in bakterijski
preizkus.
Klju~ne besede: hidroksiapatit, razpr{ilni su{ilnik, s srebrom dopiran hidroksiapatit, protiobakterijski sinteti~ni hidroksiapatit
1 INTRODUCTION
Inorganic biomaterials based on calcium orthopho-
sphate have a wide range of applications in medicine.
Among them, synthetic hydroxyapatite (HAP,
Ca10(PO4)6(OH)2) is the most promising one because of
its biocompatibility, bioactivity and osteoconductivity.
Hydroxyapatite has been used to fill a variety of bone
defects in orthopedic and maxillofacial surgeries and in
dentistry.1 Hydroxyapatite [HAP, Ca10(PO4)6(OH)2] is the
main mineral constituent of a human bone.2 This mate-
rial does not possess acceptable mechanical properties to
be used as a bulk biomaterial; however, it does demon-
strate a significant potential to be used as a coating on
metallic orthopedic and dental prostheses.3
Post-surgical infections associated with the presence
of implant materials are found with up to 5 % of
patients; the problem usually requires their removal.
Microorganism adhesions on implant surfaces represent
the initial crucial reason for an infection and lead to the
formation of a biofilm, whose microorganisms are more
resistant to antimicrobial agents. To solve the problem of
contaminating the hydroxyapatite, it has been proposed
to use antimicrobial agents such as antibiotics, fluorine
and biocide metal ions. The main concerns with
antibiotics are the development of resistant microorga-
nisms and the fact that the adsorbed antibiotics are
quickly washed out by the body fluids so that they
cannot prevent post-surgical infections in the long term.
Metal ions (Ag+, Cu2+ and Zn2+) are widely used in medi-
cine as antimicrobial agents. Silver ions, in particular,
show an oligodynamic effect with a minimum develop-
ment of the microorganism resistance.4–6
Compared with the other heavy-metal ions, silver has
demonstrated a high antimicrobial activity while main-
taining a relatively low cytotoxicity.6–8 Silver ions are
highly active ions that are strongly bound to the electron
donor groups containing sulphur, oxygen or nitrogen. In
the case of the biological molecule components such as
thio, amino, imidazole, carboxylate and phosphate, the
groups contain these electron donors.7 Studies have
shown that Ag+ ions are able to penetrate a bacterial cell
wall and cause DNA to transform to a condensed form
that reacts with the thiol-group proteins resulting in a
cell death. It was also found that silver ions are able to
interfere with the replication process. It has been demon-
Materiali in tehnologije / Materials and technology 47 (2013) 4, 431–434 431
UDK 577:66.017 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)431(2013)
strated that the higher the level of silver incorporated
into a material, the better is the antimicrobial effect, but
it comes at the cost of an increased cytotoxicity.8 There-
fore, it is a good idea to incorporate a secondary
chemical to alleviate the potential negative effects, while
maintaining optimum antimicrobial properties of Ag.7
Silver-doped ceramics have a high chemical durability
and antibacterial activity.3 Several in vitro studies
reported that the silver ions in the HAP coatings play an
important role in preventing or minimizing the initial
bacterial adhesion.7
Spray drying can be defined as a transformation of a
material from the fluid state into a dried particulate form
by spraying the feed into a hot-drying gas medium.9 The
uniformity of the shape and the diameter of the particles
is the main advantage of spray drying.10 The aims of
spray drying are the granulation of the solid phase in
spherical, monodispersed, high-density aggregates and a
quick, effective recovery of the production.11
In this study, we have produced antibacterial
silver-doped hydroxyapatite powders and reshaped them
in a spray dryer. 1 % silver was added to the hydro-
xyapatite structure with the ion-exchange technique. We
have investigated the effect of adding silver to a hydro-
xyapatite powder.
2 EXPERIMENTAL PROCEDURE
In this study, the chemical-precipitation method was
chosen for producing silver-doped hydroxyapatite.
Calcium nitrate tetra hydrate (Ca(NO3)2.4H2O, extra pure,
Merck), orthophosphoric acid (H3PO4, 85 %, Merck),
ammonium hydroxide (NH4OH, 28–30 %, Merck) and
silver nitrate (AgNO3, crystalline, 99.9+ %, Alfa Aesar)
were used as raw materials. Calcium nitrate, ortho-
phosphoric acid and silver nitrate were dissolved
separately in deionised water. Dissolved silver nitrate
was added into the calcium nitrate solution. Finally,
orthophosphoric acid was added. Ammonium hydroxide
was added to fix the pH to 10–11. The pH affected the
hydroxyapatite crystallinity.12 The solution was mixed
for 24 hours. After the precipitation, hydroxyapatite was
washed with deionised water when filtered from the
solution, then it was dried in a furnace at the temperature
of 105 °C. The powders were sintered at the temperature
of 1050 °C for 1 h under the 10 °C/min sintering regime.
Dried powders were prepared for drying with deionised
water. The inlet temperatures of 175 °C, 190 °C and 205
°C and a pressure of 1.5 bar were used for drying. 1 %
silver was added to the hydroxyapatite structure. The
plate-count test and E. coli colonies were used for the
antibacterial activity.
3 RESULT AND DISCUSSION
3.1 XRD results
It is seen in Figure 1 that tricalcium phosphate and
silver oxide were in the powder structure. Hydro-
xyapatite decomposes when heated to form 	-TCP.11 The
F. E. BASTAN, Y. Y. ÖZBEK: PRODUCING ANTIBACTERIAL SILVER-DOPED HYDROXYAPATITE POWDERS ...
432 Materiali in tehnologije / Materials and technology 47 (2013) 4, 431–434
Figure 1: XRD peaks after sintering (TCP: tricalcium phosphate,
Ag2O: silver oxide)
Slika 1: XRD-vrhovi po sintranju (TCP: trikalcij fosfat, Ag2O:
srebrov oksid)
Figure 2: SEM images of the powders after spray drying at the inlet
temperatures: a) 175 °C, b) 190 °C, c) 205 °C
Slika 2: SEM-posnetki prahov po razpr{ilnem su{enju pri vstopni
temperaturi: a) 175 °C, b) 190 °C, c) 205 °C
secondary phases such as -tricalcium phosphate and
	-tricalcium phosphate were also found in the coatings
to a small extent. The phases that are undesired such as
tetra-tricalcium phosphate (TTCP) and calcium oxide
(CaO) were absent. In the samples that contained Ag2O,
a characteristic Ag2O (101) peak was observed as well as
the peak shifts, indicating an effective incorporation of
Ag2O into the apatite structure.13,14
Metallic silver precipitates silver oxide according to
the potential-pH diagram in an aqueous solution.15 All
the other peaks are hydroxyapatite peaks.
3.2 SEM results
However, the shapes of the powders are not totally
spherical, though they are close to a spherical shape. The
shape is an important phenomenon for the flowability
and apparent density. When the inlet temperature was
205 °C, all the products were dried with the heat energy.
At the other temperatures (175 °C and 190 °C), the
problem was that the powders tended to stick together
because of the moisture. It is seen in Figure 2 that the
temperature is an important factor affecting the particle
size. Using high heat energy, particles were dried very
quickly. Hence, the particle size decreased.
3.3 Particle-size-analysis results
It is seen in Figure 3 that the average particle size
was 37 μm for A and 28 μm for B. The particle size
decreased with the increasing temperature. High energy
provided quick drying and small particles were
produced. Powder diameters after spray drying were
recorded in Table 1.
Table 1: Mean diameters after spray drying (analyzed with a Micro-
trac S3500)
Tabela 1: Srednji premer zrn po razpr{ilnem su{enju (analizirano z
Microtrac S3500)
Mean diameters 190 °C inlettemperature
205 °C inlet
temperature
d10 19.89 μm 16.91 μm
d50 37.08 μm 28.52 μm
d90 67.13 μm 65.40 μm
It is clear from Table 1 that the higher temperature
led to a smaller particle size for all the mean diameters
(d10, d50, d90).
3.4 Antibacterial-test results
Figure 4a shows that there are 63+ bacterial colonies
in the medium and there are no bacterial colonies in the
medium in Figure 4b. The hydroxyapatite powders with
a 1 % silver addition show a 100 % antibacterial activity.
Silver inhibited the growth of bacterial colonies. On the
other hand, the amount of silver is important with respect
to biocompatibility as an increased amount of silver
F. E. BASTAN, Y. Y. ÖZBEK: PRODUCING ANTIBACTERIAL SILVER-DOPED HYDROXYAPATITE POWDERS ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 431–434 433
Figure 4: After the antibacterial test: a) no additive to hydroxyapatite,
b) 1 % mol fraction Ag added to hydroxyapatite
Slika 4: Po protibakterijskem preizkusu: a) hidroksiapatit brez do-
datkov, b) hidroksiapatit z dodanim molskim dele`em 1 % Ag
Figure 3: Analysis of the particle-size distribution after spray drying
(inlet temperatures: a) 190 °C, b) 205 °C)
Slika 3: [u{enje po razpr{evanju. Analiza razporeditve velikosti zrn
(vstopna temperatura: a) 190 °C, b) 205 °C).
causes a decrease in biocompatibility. A high content of
silver leads to a poor biocompatibility because the
alkaline phosphatase activity of osteoblasts increases.16,17
It was reported that the toxicity of Ag ions affected
the basic metabolic cellular functions common to all the
specialized mammalian cells. A concentration- and
time-dependent depletion of the intracellular ATP
content was attributed to the presence of Ag ions, there-
by compromising the cell energy charge that precedes
the cell death. Therefore, it is prudent to incorporate a
minimum amount of Ag on the implant surfaces to
adequately reduce the bacterial adhesion as well as
minimizing the tissue cytotoxicity.18
4 CONCLUSION
HA coating materials have been well studied as
osteogenic enhancing materials and are effectively used
in several medical and dental applications. In this study,
the silver-doped hydroxyapatite was produced and
reshaped in a spray dryer. The shaping is very important
in the coating application because the shape affects the
fluidity.
The hydroxyapatite structure decomposed into
hydroxyapatite and tricalcium phosphate phases at a high
temperature. It was seen that the hydroxyapatite powders
including silver show a 100 % antibacterial effect. So,
we were able to effectively enhance the coating with
sustainable, long-term antimicrobial properties, while
minimizing the negative cytoplasmic effects on the
osteoblast cells.
5 REFERENCES
1 H. F. Lin, L. Chun-Jen, C. Ko-Shao, Thermal reconstruction behavior
of the quenched hydroxyapatite powder during reheating in air,
Materials Science and Engineering: C, 13 (2000) 1–2, 97–104
2 C. F. Koch, S. Johnson, D. Kumar, Pulsed laser deposition of hydro-
xyapatite thin films, Materials Science and Engineering: C, 27
(2007) 3, 484–494
3 S. Johnson, M. Haluska, R. J. Narayan, In situ annealing of hydro-
xyapatite thin films, Materials Science and Engineering: C, 26
(2006) 8, 1312–1316
4 V. Stani}, D. Jana}kovi}, S. Dimitrijevi}, S. B. Tanaskovi}, M. Mi-
tri}, M. S. Pavlovi}, A. Krsti}, D. Jovanovi}, S. Rai~evi}, Synthesis
of Antimicrobial Monophase Silver-Doped Hydroxyapatite Nano-
powders for Bone Tissue Engineering, Applied Surface Science, 257
(2011) 9, 4510–4518
5 L. Harris, R. Richards, Staphylococci and implant surfaces: a review,
Injury, 37 (2006), S3–S14
6 A. Montali, Antibacterial coating systems, Injury, 37 (2006),
S81–S86
7 G. A. Fielding, M. R. Bandyopadhyay, A. S. Bose, Antibacterial and
biological characteristics of silver containing and strontium doped
plasma sprayed hydroxyapatite coatings, Acta Biomaterialia, 8
(2012) 8, 3144–3152
8 K. Gross, Hydroxyapatite – Synthesis of Hydroxyapatite Powders,
Journal of Biomedical Materials Research, 62 (2002) 2, 228–236
9 V. Stanic, D. Janackovic, S. Dimitrijevic, Synthesis of antimicrobial
monophase silver-doped hydroxyapatite nanopowders for bone tissue
engineering, Applied Surface Science, 257 (2011) 9, 4510–4518
10 H. Jeon, S. Yi, S. Oh, Preparation and antibacterial effects of
Ag-SiO2 thin films by sol-gel method, Journal of Materials Science,
42 (2007) 24, 10085–10089
11 A. L. R. Rattes, W. P. Oliveira, Spray drying conditions and encap-
sulating composition effects on formation and properties of sodium
diclofenac microparticles, Powder Technology, 171 (2007) 1, 7–14
12 J. M. Gac, L. Gradon, A distributed parameter model for the spray
drying of multicomponent droplets with a crust formation, Advanced
Powder Technology, 24 (2013) 1, 324–330
13 M. Serantoni, P. Piancastelli, A. L. Costa, L. Esposito, Improvements
in the production of Yb:YAG transparent ceramic materials: Spray
drying optimisation, Optical Materials, 34 (2012) 6, 995–1001
14 A. K. Nayak, Hydroxyapatite Synthesis Methodologies: An Over-
view, International Journal of ChemTech Research, 2 (2010) 2,
903–907
15 V. V. Skorokhod, S. M. Solonin, V. A. Dubok, L. L. Kolomiets,
Decomposition Activation of Hydroxyapatite in Contact with
-Tricalcium Phosphate, Powder Metallurgy and Metal Ceramics, 49
(2010), 324–329
16 M. Diaz, F. Barba, M. Miranda, F. Guitian, R. Torrecillas, J. S.
Moya, Synthesis and Antimicrobial Activity of a Silver-Hydroxy-
apatite Nanocomposite, Journal of Nanomaterials, (2009), doi:
10.1155/2009/498505
17 J. Qu, X. Lu, D. Li, Silver/hydroxyapatite composite coatings on
porous titanium surfaces by sol-gel method, Journal of Biomedical
Materials Research Part B: Applied Biomaterials, 97B (2011) 1,
4048
18 W. Chen, Y. Liu, H. S. Courtney, M. Bettenga, C. M. Agrawal, In
vitro anti-bacterial and biological properties of magnetron co-sput-
tered silver-containing hydroxyapatite coating, Biomaterials, 27
(2006) 32, 5512–5517
F. E. BASTAN, Y. Y. ÖZBEK: PRODUCING ANTIBACTERIAL SILVER-DOPED HYDROXYAPATITE POWDERS ...
434 Materiali in tehnologije / Materials and technology 47 (2013) 4, 431–434
D. KEK MERL et al.: TRIBOCORROSION DEGRADATION OF PROTECTIVE COATINGS ON STAINLESS STEEL
TRIBOCORROSION DEGRADATION OF PROTECTIVE
COATINGS ON STAINLESS STEEL
TRIBOKOROZIJSKA DEGRADACIJA ZA[^ITNIH PREVLEK NA
NERJAVNEM JEKLU
Darja Kek Merl, Peter Panjan, Miha ^ekada, Peter Gselman, Sre~ko Paskvale
Jo`ef Stefan Institute, Jamova 39, 1000 Ljubljana, Slovenia
darja.kek@ijs.si
Prejem rokopisa – received: 2012-10-12; sprejem za objavo – accepted for publication: 2012-12-18
Wear, erosion, corrosion and other forms of material deterioration lead to a significant decrease in the performance of industrial
production and thus increase the production cost. The PVD protective coatings based on nitrides, carbides and oxides of
transitional metals have been developed in order to reduce these problems. Further progress in this field, however, depends on
the understanding of the surface and interface effects, especially in the case of tribocorrosion when degradation of materials
results from a combination of tribological and electrochemical processes.
This work focuses on the methodology that allows one to conduct the sliding wear and corrosion tests simultaneously and to
follow in situ the degradation processes of the coating systems by controlling (in real time) the friction coefficient, corrosion
current and corrosion potential. Using practical examples from the area of biomedical applications, i.e., the stainless-steel 316L
substrate and the stainless-steel substrate protected with the TiAgN and TiSiN hard coatings, it is shown that tribocorrosion
experiments can be used to design and predict the properties of new coatings with an enhanced performance and stability.
Keywords: tribocorrosion, PVD coatings, stainless steel, ball-on-disc, electrochemical method
Obraba, erozija, korozija in druge oblike degradacije materiala vodijo do manj{e produktivnosti in s tem do ve~jih stro{kov
proizvodnje. Da bi zmanj{ali omenjene te`ave, so bile razvite razli~ne PVD-prevleke. Nadaljnji napredek na tem podro~ju je
odvisen od razumevanja procesov na povr{inah in faznih mejah, {e posebej v primeru tribokorozije, kjer gre za degradacijo
materiala zaradi tribolo{kih in elektrokemijskih procesov.
V delu se osredinjamo na metodologijo, ki omogo~a hkrati merjenje parametrov drsne obrabe (npr. koeficienta trenja) in
elektrokemijskih parametrov (korozijski potencial, korozijski tok). Tako lahko in-situ spremljamo degradacijo prevlek ali
pasivne plasti na pasivnih kovinah. Na prakti~nih primerih tribokorozijskih preizkusov (dveh prevlek TiAgN in TiSiN ter
podlage 316L) smo pokazali, da je mogo~e tribokorozijske preizkuse uporabljati za na~rtovanje in napovedovanje lastnosti
prevlek z optimalnimi lastnostmi.
Klju~ne besede: tribokorozija, PVD-prevleke, nerjavno jeklo, pin-on-disc, elektrokemijske metode
1 INTRODUCTION
Wear and corrosion of materials are the most
important failure mechanisms in industry.1 Significant
efforts have been made to combat wear and corrosion.
Hard coatings such as nitrides and carbides are success-
ful examples in protecting surface against dry wear.1,2
Passivation techniques were shown to improve signi-
ficantly the corrosion properties of passive materials.3,4
However, when wear and corrosion are involved simul-
taneously, the synergistic action of wear and corrosion
may deteriorate the performance of the materials. For
instance, hard coatings may perform poorly in a corro-
sive medium due to defect growth.5–7 On the other hand,
corrosion-resistant materials, such as stainless steels and
titanium alloys, may lose their corrosion resistance and
corrosive-wear conditions as the passive layer conti-
nuously abrades away by wear.8–10
In general, tribocorrosion is defined as a degradation
of materials that results from a combination of tribo-
logical and electrochemical processes.11,12 The material-
degradation rate, whether expressed as a loss in the
component performance or a loss in the material, is very
often higher than the sum of the rates of corrosion and
wear acting separately. The synergy of corrosion and
wear is still poorly understood due to the complexity of
the problem and a lack of appropriate experimental
techniques. The material resistance to tribocorrosion is
very often evaluated using the conventional wear testers
such as pin-on-disc and simply immersing the contact in
a specific electrolyte. These testers can be used to rank
the materials but not to investigate the mechanism
responsible for tribocorrosion. Until recently, no effec-
tive analytic methods have been available to investigate
the separate roles of the wear and corrosion in the
corrosive wear, which are crucial for understanding the
corrosive-wear synergy. Currently, in industry, the selec-
tion of the material to resist corrosive wear is mainly
based on experience and trial-and-error testing. There-
fore, it is useful to develop the guidelines for materials
selection and design.
In this paper, we present the characterization of
material resistance to tribocorrosion. Electrochemical
techniques are combined with tribological testing in
order to assess the roles of wear and corrosion. Addition-
ally, microhardness was measured.
Materiali in tehnologije / Materials and technology 47 (2013) 4, 435–439 435
UDK 620.193:620.197:621.793 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)435(2013)
2 EXPERIMENTAL WORK
Tribocorrosion experiments were conducted using a
CSM tribometer with a pin-on-disc configuration. A
special configuration of a cell (Figure 1) allowed us
simultaneous measurements of the wear (the friction
coefficient) and electrochemical parameters (open-circuit
potential – OCP). The ball-sample contact was fully
immersed in the test electrolyte. The sample served as
the working electrode and its potential was measured
with respect to a standard Ag/AgCl electrode. A plati-
num wire was used as the counter electrode.
The test electrolyte was Hank’s solution which
simulates the body fluid environment with the compo-
sition of NaCl 8 g/L, Na2HPO4 0.0475 g/L, NaHCO3
0.35 g/L, KCl 0.5 g/L, KH2PO4 0.06 g/L, MgCl2 · 6H2O
0.1 g/L, MgSO4 · 7H2O 0.10 g/L, CaCl2 0.18 g/L and
1 g/L glucose at pH = 7.2. As the test samples stain-
less-steel AISI 316L discs were used as the substrates
with two types of coatings (TiAgN and TiSiN) deposited
on the AISI 316L substrates.
The coatings (with the thickness of 1.6 μm for TiSiN
and 1.9 μm for TiAgN ) were deposited with the
plasma-beam-sputtering technique in a Sputron (Balzers)
apparatus. The polished stainless-steel samples (Ra = 25
nm) were used as substrates. The vacuum chamber was
evacuated to the base pressure of approximately 2 mPa.
The nitrogen (with the purity of 99.995 %) flow
remained constant at 5 sccm. During the deposition the
bias voltage was –30 V and the temperature not did not
exceed 130 °C. For enhancement an adhesion layer of
pure titanium (around 100 nm) was sputtered on the
substrate at the beginning of the deposition.
The wear tests were conducted with an alumina ball
sliding on the plate of a sample at a fixed normal load of
8 N. The sliding was performed for 1800 laps at a sliding
radius of 2 mm. The corresponding Hertz contact
pressure was around 1.5 GPa.
The hardness was measured using a Fischerscope
H100 indenter with the loads of 10 mN to 1000 mN.
3 RESULTS AND DISCUSSION
Tribocorrosion involves tribological and corrosion
processes. Due to the synergistic interaction between
them, the material loss can be larger than the sum of the
losses due to the wear and corrosion acting separately.
The total mass loss is expressed as:
Mtot = Mmech + Mc + Msyn (1)
where Mtot is the total mass loss, Mmech is the mass loss
due to the pure mechanical wear in the absence of
corrosion, Mc is the mass loss due to the static corrosion
in the absence of wear and Msyn is the mass loss due to
the synergy effect. Friction and wear change the corro-
sion behavior of a material and vice-versa; corrosion
may change the condition of friction.13 Therefore, Msyn
can be expressed as the sum of the two components,
Mc-w and Mw-c, where Mc-w is the increase in the
corrosion loss due to a tribological action, called the
wear-enhanced corrosion, and Mw-c is the increase in the
mechanical wear due to the corrosion, called the
corrosion-enhanced wear. Therefore, equation (1) can be
expressed as:
Mtot = Mmech+ Mc + Mc-w + Mw-c (2)
Mw-c could be attributed to the inhomogeneous
distribution of strain, defect and surface irregularities
caused by the mechanical wear. In such a way micro-
electrodes (cathodic and anodic places) are generated at
the surface. This leads to an increase in the material
dissolution. A destruction of the passive film by wear
can also lead to the material dissolution. Mc-w is more
relevant in coated substrates, where the electrolyte
penetrates through the pores and defects in the coatings.
These can lead to a reduction of the tribological pro-
perties of the substrate/coating system.
The tribocorrosion experiments were performed on
the bare stainless-steel AISI316L substrate and on the
samples coated with two nitride-based coatings. The
D. KEK MERL et al.: TRIBOCORROSION DEGRADATION OF PROTECTIVE COATINGS ON STAINLESS STEEL
436 Materiali in tehnologije / Materials and technology 47 (2013) 4, 435–439
Figure 2: Evolution of the corrosion potential before, during and after
the wear of the bare substrate and both coatings (TiAgN, TiSiN)
Slika 2: Spreminjanje korozijskega potenciala (OCP) pred obrabo,
med njo in po njej za samo podlago in obe prevleki (TiAgN, TiSiN)
Figure 1: Scheme of the experimental cell that allowed us simul-
taneous measurements of wear (friction coefficient) and electro-
chemical parameters (open-circuit-potential – OCP)
Slika 1: Shema eksperimentalne celice, ki omogo~a hkratno merjenje
obrabnih in elektrokemijskih parametrov
measurement consisted of monitoring the OCP before,
during and after the sliding wear. The drop in the
potential can be the result of the destruction of the
protective passive layer due to the wear mechanism.
After the sliding was stopped the corrosion potential
increased. Therefore, the repassivation of the passive
layer occurred.
Figure 2 shows the corrosion potential (OCP) evo-
lution of the two coated stainless-steel samples and of
the bare stainless-steel substrate in Hank’s solution. For
the bare substrate, the OCP dropped gradually to the
lowest value of –375 mV vs Ag/AgCl after almost 1700
cycles of sliding. This gradual drop in the potential
suggests a gradual removal of the oxide passive layer.
When the sliding ceased, the OCP increased and pro-
gressively returned to its original value. Thus, the
repassivation of the passive layer occurred. For the
coated samples, the OCP dropped to reach the potential
values of –80 mV vs Ag/AgCl for the TiAgN and of
–110 mV vs Ag/AgCl for the TiSiN coatings. The drop
in the potential on the bare substrate is the result of the
destruction of the protective passive layer due to the
wear mechanism. After the sliding was stopped the
corrosion potential increased.
The explanation for the drop and increase in the
potential for the coated substrates is not as straight-
forward as in the case of the bare substrate. In this case
there is no formation of a passive layer on the inert
nitride coating. The drop in the potential suggests a
reduction in the corrosion properties of the substrate/
coating system due to the penetration of the electrolyte
through the pores and defects in the coatings, as
explained before in equation (2) (Mc-w part). However,
more data are needed to support this statement.
The evolution of the friction coefficient was analyzed
during the sliding in Hank’s solution for both coatings
and the bare substrate. Figure 3 shows the representative
curves of the observed trends. For the bare substrate,
there was an increment in the values of the friction
coefficients at the beginning of the sliding process (250
laps). After the initial running-in period, the value of the
friction coefficient remains stable at around 0.6. The
fluctuation in the value of the friction coefficient
corresponds to the removal and regrowth of the passive
layer.13 When the substrate was coated with TiSiN or
TiAgN, the friction coefficient dropped by about five
times to about 0.15. However, it should be pointed out,
that for the TiAgN coatings the friction coefficient did
not converge to a stable value during 1800 laps. An
increase in the friction coefficient during the sliding was
observed (Figure 3) for the TiAgN coating. Such
behavior may be due to the formation of the corrosion
products and wear debris during the sliding process, as
also reported in literature.13 In general, the tribocorrosion
parameters are not the material properties, but the pro-
perties of the system. Therefore, the fiction coefficient
depends on the sliding processes (i.e., parameters, elec-
trolyte).
After the tribocorrosion test, the wear track was
analyzed by a profilometer (Figures 4 and 5). Figure 4
shows the average depths of the worn profiles on both
coatings, TiAgN and TiSiN, and on the bare substrate.
The depth of the worn profile was estimated and com-
pared to the thickness of the coatings. The wear-track
depth is around 0.70 μm for the bare substrate, 0.33 μm
for the TiSiN coatings and 0.15 μm for the TiAgN
coatings. Based on the results, we can find out if the
coatings were removed completely or not after the
tribocorrosion test. In the case of the TiSiN coatings 20
% was removed, while only 8 % of the TiAgN coatings
were removed. This shows that the current systems
(when the coating and the substrate are considered)
exhibit good tribocorrosion properties.
Hardness is one of the key mechanical properties of
the protective coatings. However, the result obtained
with the measurement strongly depends on the load
D. KEK MERL et al.: TRIBOCORROSION DEGRADATION OF PROTECTIVE COATINGS ON STAINLESS STEEL
Materiali in tehnologije / Materials and technology 47 (2013) 4, 435–439 437
Figure 4: Average depth profile after the tribocorrosion experiment of
the bare substrate and both coatings (TiAgN, TiSiN)
Slika 4: Povpre~ni globinski profil podlage in obeh prevlek (TiAgN,
TiSiN) po tribokorozijskem preizku{anju
Figure 3: Evolution of the friction coefficient during the sliding in
Hank’s solution of the bare substrate and both coatings (TiAgN,
TiSiN)
Slika 3: Spreminjanje koeficienta trenja med razenjem v Hankovi raz-
topini za podlago in obe prevleki (TiAgN, TiSiN)
applied, as well as on the sample preparation. At smaller
loads there is no substrate influence on the hardness
(when the indentation depth is less than about 10 % of
the coating thickness). Therefore, only the coating pro-
perties are measured. At higher loads, on the other hand,
the substrate influence comes into sight. In our case, the
measurements were taken at the applied loads from 10
mN to 100 mN. It should be noted, that, at a lower load,
the error of the measurement grows because of various
surface affects (roughness, defects, inhomogeneities, tip
irregularities, etc.).
Figure 6 presents the hardness in dependence of the
relative indentation depth (RID) for both coatings in the
load range from 10 mN to 100 mN. The relative inden-
tation depth is computed by dividing the indentation
depth by the coating thickness of the corresponding
system.14 The hardness of the bulk substrate was deter-
mined to be 200 HV. The hardness, measured at the 10
mN load reached the value of around 3800 HV for the
TiSiN coating, and 2900 HV for the TiAgN coatings.
Thus, with a coating deposition the surface hardness
increases by almost 20-fold. However, the highest value
of the surface hardness does not imply the best wear
resistance. One possible reason why TiSiN exhibits a
lower wear resistance than TiAgN is the presence of the
hard Si3N4 particles that increase the wear.
4 CONCLUSION
In this paper, the hardness and tribocorrosion charac-
terization of the TiAgN and TiSiN coatings and bare
stainless-steel substrate was performed. Electrochemical
techniques were combined with friction tests to under-
stand the synergy between wear and corrosion.
For the bare substrate, the drop in the potential is the
result of the destruction of the protective passive layer
due to the wear mechanism. After the sliding was
stopped the corrosion potential increased. Therefore, a
repassivation of the oxide layer occurred. The friction
coefficient of the bare substrate and the depth of the wear
track exhibit the wear resistance that is five times lower
than that of the coated substrate.
For the TiAgN and TiSiN coatings, the drop in the
potential during the wear suggests a reduction in the
corrosion properties of the substrate/coating system due
to the penetration of the electrolyte through the pores
and defects in the coatings. Both coatings improved the
tribocorrosion and mechanical stability compared to the
bare stainless-steel substrate. The TiAgN coating exhi-
bits a better wear resistance than the TiSiN coating,
although the surface hardness of the TiSiN coating is
higher than that of the TiAgN coatings.
D. KEK MERL et al.: TRIBOCORROSION DEGRADATION OF PROTECTIVE COATINGS ON STAINLESS STEEL
438 Materiali in tehnologije / Materials and technology 47 (2013) 4, 435–439
Figure 6: Vickers hardness as a function of the relative indentation
depth (RID) for both coatings at the loads from 10 mN to 100 mN
Slika 6: Trodota po Vikersu v odvisnosti od relativnega globinskega
profila (RID) za obe prevleki (TiAgN, TiSiN) pri obremenitvi od 10
mN do 100 mN
Figure 5: 3D micrographs of the wear tracks after tribocorrosion
experiments for the bare substrate and both coatings
Slika 5: 3D globinski profil raze podlage in obeh prevlek (TiAgN,
TiSiN) po tribokorozijskem preizku{anju
Acknowledgements
The work was supported by the Slovenian Research
Agency – ARRS (project L2-4173 and research program
P2-0082).
5 REFERENCES
1 P. Panjan, M. Cekada, D. K. Merl, M. Macek, B. Navinsek, A. Jesih,
A. Zalar, M. Jenko, Vacuum, 71 (2003), 261
2 P. Panjan, I. Boncina, J. Bevk, M. Cekada, Surf. Coat. Tech., 200
(2005), 133
3 Z. Bou-Saleh, A. Shahryari, S. Omanovic, Thin Solid Films, 515
(2007), 4727
4 A. Shahryari, F. Azari, H. Vali, S. Omanovic, Acta Biomater., 6
(2010), 695
5 D. K. Merl, P. Panjan, M. Panjan, M. Cekada, Plasma Process
Polym., 4 (2007), S613
6 P. Gselman, T. Boncina, F. Zupanic, P. Panjan, D. K. Merl, M.
Cekada, Mater. Tehnol., 46 (2012) 4, 351–354
7 D. K. Merl, I. Milosev, P. Panjan, F. Zupanic, Mater. Tehnol., 45
(2011) 6, 593–597
8 Y. Yan, A. Neville, D. Dowson, S. Williams, Tribol. Int., 39 (2006),
1509
9 A. Igual Munoz, J. L. Casaban, Electrochimica acta, 55 (2010), 5428
10 N. Papageorio, S. Mishler, Tribology Letter, 48 (2012), 271
11 J. P. Celis, P. Ponthiaux, F. Wenger, Wear, 261 (2006), 939
12 D. Landolt, S. Mischler, M. Stemp, S. Barril, Wear, 256 (2004), 517
13 S. Mischler, Tribol. Int., 41 (2008), 573
14 D. Beegan, S. Chowdhury, M. T. Laugier, Thin Solid Films, 516
(2008), 3813
D. KEK MERL et al.: TRIBOCORROSION DEGRADATION OF PROTECTIVE COATINGS ON STAINLESS STEEL
Materiali in tehnologije / Materials and technology 47 (2013) 4, 435–439 439

V. GLIHA et al.: BEHAVIOUR OF SHORT CRACKS EMANATING FROM TINY DRILLED HOLES
BEHAVIOUR OF SHORT CRACKS EMANATING FROM
TINY DRILLED HOLES
VEDENJE KRATKIH RAZPOK, KI NASTANEJO IZ MAJHNIH
IZVRTIN
Vladimir Gliha1, Pavlo Maruschak2, Toma` Vuherer1
1University of Maribor, Faculty of Mechanical Engineering, Smetanova 17, 2000 Maribor, Slovenia
2Ternopil Ivan Pul’uj National Technical University, Ruska 56, 46001 Ternopil, Ukraine
vladimir.gliha@um.si
Prejem rokopisa – received: 2012-10-15; sprejem za objavo – accepted for publication: 2012-11-27
Specimens with martensitic microstructure were defected by tiny drilled holes with the existing local residual stresses induced
by drilling and without them. The objective of this research was to determine the cyclic stress level for the crack initiation and
fatigue limit of the defected and smooth specimens’ dependence upon the residual-stress field. Compressive residual stresses
retarded the crack initiation. Immediately after the crack initiation, residual stresses decelerated the short-crack propagation, but
later, when the residual-stress sign was changed, they accelerated it. Non-propagating short-crack size and fatigue limit also
depend on the residual-stress field.
Keywords: small defect, drilled hole, crack initiation, crack propagation, short crack, long crack, anomalous fast-crack
propagation
Pri preizku{ancih z martenzitno mikrostrukturo smo naredili majhne izvrtine z obstoje~imi lokalnimi zaostalimi napetostmi
zaradi vrtanja in brez njih. Namen te raziskave je bilo dolo~iti odvisnost cikli~ne napetosti, ki je potrebna za nastanek razpoke,
in odvisnost trajne dinami~ne trdnosti z izvrtino oslabljenih in gladkih preizku{ancev, od polja zaostalih napetosti. Tla~ne
zaostale napetosti so zavrle nastanek razpoke. Takoj po nastanku razpoke so zaostale napetosti upo~asnile {irjenje kratke
razpoke, pozneje, ko se je predznak napetosti spremenil, pa so ga pospe{ile. Velikost kratke razpoke, ki se ne {iri ve~, in trajna
dinami~na trdnost sta prav tako odvisni od polja zaostalih napetosti.
Klju~ne besede: majhne napake, izvrtina, nastanek razpoke, {irjenje razpoke, kratka razpoka, dolga razpoka, nenormalno hitro
{irjenje razpoke
1 INTRODUCTION
The phenomenon of crack initiation from small
defects in metals caused by cyclic stress at a level lower
than the fatigue limit, in connection with the initial
fatigue-crack propagation to the size when it becomes a
non-propagating crack is now well understood.1–3 The
hardness and the defect size are crucial for the fatigue
limit of the metals with small defects.4,5 The measure of
the defect size is the square root of its projection onto the
plane of cyclic stress, i.e., the parameter area. The
measure of hardness is the HV number.
However, most of the experiments confirming the
above-mentioned relation have been performed on
specimens with small, three-dimensional, artificial
surface defects. The effects of the eventual residual
stresses due to the artificial defect preparation have not
been taken into account. Of course, the local residual
stresses induced by drilling, scratching, impressing, etc.,
affect the crack initiation and the initial fatigue-crack
propagation. The latter is known as the short-crack pro-
pagation. Acceleration or retardation of a crack initiation
and assistance or obstruction of the short-crack propaga-
tion are reflected in the sizes of the largest non-propagat-
ing cracks emanating from small defects.5,6
The influence of the local residual stresses caused by
Vickers-pyramid indenting in an artificially prepared
martensitic microstructure on the crack initiation and
subsequent propagation was the matter of our previous
paper.7 We have found that these stresses play an impor-
tant role in the crack initiation from Vickers indentations
and its initial propagation. Actually, at the very beginn-
ing, the tensile residual stresses around both exposed
corners of an indentation accelerate the initiation of two
cracks along the edges. Later on, the compressive resi-
dual stresses around the top of the indentation retard the
short-crack propagation along the edges in the direction
towards the top. When both cracks are linked and
become a single crack, which is larger than individual
short cracks, it starts to behave as a long crack. Now, the
tensile residual stresses around the corners accelerate the
crack propagation at the surface. This large crack also
assists the crack penetration into the depth. As a result,
the indentations with present local residual stresses seem
to be smaller than expected.4,7
The experimental results of the crack initiation from
tiny drilled holes, instead of Vickers indentations, as well
as the fatigue limits of smooth and defected specimens
with identical microstructures are shown in the present
article. The stress level for the crack initiation was expe-
rimentally assessed. On the specimens in the as-drilled
Materiali in tehnologije / Materials and technology 47 (2013) 4, 441–446 441
UDK 539.42 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)441(2013)
condition, crack sizes at different stress levels were
measured in order to analyse the effects of the residual-
stress field on the short-crack-propagation rate.
2 EXPERIMENTAL WORK
2.1 Samples and specimen preparation
Samples of the material with martensite microstruc-
ture were used for experimental work. The chemical
composition and the mechanical properties of this mate-
rial are given in Tables 1 and 2.
Table 1: Chemical composition of the material (mass fractions, w/%)
Tabela 1: Kemijska sestava materiala (masni dele`i, w/%)
C Si Mn P S Cr Ni Cu Mo Al
0.18 0.22 0.43 0.012 0.028 1.56 1.48 0.15 0.28 0.023
Table 2: Mechanical properties of the material
Tabela 2: Mehanske lastnosti materiala
Rp02/
MPa
Rm/
MPa
A5/
%
Z/
%
Impact
toughness
(at +20 °C)/J
Hardness/
(HV10)
1042 1431 15 56 72 463
A two-step heat treatment was used to create the
microstructure of the samples. It is schematically shown
in Figure 1 (the solid line). The first stage of the heat
treatment is long-lasting high-temperature annealing and
water quenching (mark b), while the second stage is
hardening by water quenching (d). The grains have
grown during the annealing. The martensite transfor-
mation was a result of water quenching (Figure 2).
Cylinders were machined from the as-delivered steel
17CrNiMo7 (mark a). In the first sequence of the
specimen preparation high-temperature annealing and
quenching were applied (b), and then the cylinders were
turned to produce grooved cylinders (c). The grooved
area was fine polished.
The first sequence of the specimen preparation was
the same for all the specimens. In the second sequence of
the specimen preparation, three different types of
grooved cylinders were produced, i.e., type-I, type-II and
type-III specimens.
I. The grooved cylinders were hardened (d), while the
surface at the bottom of the groove remained intact.
The type I-specimens were smooth and in the
residual-stress-free condition.
II. Before hardening the cylinders were hole-drilled at
the bottom of the groove. The diameter of the hole
was 90 μm, while the depth was 45 μm. The local
residual stresses induced by hole-drilling were
relaxed due to heating over the Ac3 temperature
during the hardening (d). The type-II specimens were
defected in the residual-stress-free condition.
III. The hardening (d) was followed by groove polishing
and hole drilling. Therefore, the local residual-stress
field induced by drilling existed in the specimens.
The type-III specimens were defected and were in the
as-drilled condition.
The cylinders were slightly distorted after the second
sequence of the specimen preparation. In order to avoid
any possible influence on the stress level during the
loading, the cylinders were turned, in the third sequence
of the specimen preparation, to the final size and shape
(mark e). The axis of the specimens was adjusted to the
centre of the existing groove. The final stress-concen-
tration coefficient was 1.73.8 At the end of the specimen
preparation the grooves were cleaned to make the surface
suitable for observations.
The specimens were loaded, at room temperature,
onto the rotary bending machine with a ratio R = –1. The
levels, at which the crack initiated from the tiny drilled
hole and began to propagate, were determined experi-
mentally. The bottom of the groove was analysed after
every stop using a light microscope mounted on the
loading machine. SEM was used on the dismounted spe-
cimens. The fatigue limit was defined as 107 cycles
strength.
V. GLIHA et al.: BEHAVIOUR OF SHORT CRACKS EMANATING FROM TINY DRILLED HOLES
442 Materiali in tehnologije / Materials and technology 47 (2013) 4, 441–446
Figure 2: Microstructure of the material samples
Slika 2: Mikrostruktura vzorcev materiala
Figure 1: Procedure for the specimen preparation: a) cylinder machin-
ing, b) high-temperature annealing and quenching, c) thinning and
machining a groove with 
 = 2 mm, d) hardening, e) specimen
aligning by thinning and shortening
Slika 1: Postopek priprave preizku{ancev: a) izdelava valja, b) visoko-
temperaturno `arjenje in kaljenje, c) tanj{anje in izdelava `leba z 
 = 2
mm, d) kaljenje, e) poravnavanje preizku{ancev s tanj{anjem in kraj-
{anjem
2.2 Cracks registration
The type-I specimens were smooth at the bottom of
the groove, i.e., without small defects:
A photograph of small cracks in the type-I specimen
at the stress level lower than the fatigue limit is shown in
Figure 3a. The cracks are non-propagating. Their size is
2a = 50–60 μm.
A photograph of larger non-propagating cracks at the
stress level higher than the former one and equal to the
fatigue limit is shown in Figure 3b. A net of small indi-
vidual cracks can be seen. The size of the largest one is
2a  300 μm. All the other ones are minor cracks.
The type-II specimens were defected by tiny holes.
They were in the residual-stress-relaxed condition:
A photograph of the specimen with the drilled hole at
the bottom of the groove at the stress level lower than the
level for the crack initiation is shown in Figure 4a.
There is no crack initiated.
A photograph of the specimen at the stress level
above the level for the crack initiation is shown in
Figure 4b. The cracks on both sides of the hole are
non-propagating. The cracks including the hole behave
as a single short, non-propagating crack. Its size is 2a =
165 μm.
The type-III specimens were in the as-drilled condi-
tion, i.e., with the existing residual stresses:
A photograph of the initiated crack is shown in
Figure 5a. Only one extremely short crack is registered.
The crack including the hole behaves as a single short,
non-propagating crack. Its length is 2a = 109.5 μm.
A photograph of the cracks at the stress level slightly
higher than the fatigue limit is shown in Figure 5b. Two
cracks were initiated at each side of the hole. The cracks
including the hole behave as a single, but propagating
crack. In this moment its size is 2a = 202 μm. After-
wards, the loading was continued. The specimen finally
fractured before 107 cycles.
2.3 Short-crack analysis
In order to register the short-crack propagation in the
specimens with the existing residual-stress field induced
by hole-drilling and evaluate short-crack-propagation
rate, some of the specimens were tested in detail at the
stress levels between the levels for the crack initiation
and the fatigue limit. One of the specimens was tested a
V. GLIHA et al.: BEHAVIOUR OF SHORT CRACKS EMANATING FROM TINY DRILLED HOLES
Materiali in tehnologije / Materials and technology 47 (2013) 4, 441–446 443
Figure 4: Two different stages at the bottom of the groove in the
type-II specimens after 107 cycles: a) no cracks at the stress level of
357 MPa, b) non-propagating cracks at the stress level of 405 MPa
Slika 4: Dve razli~ni stanji na dnu `leba pri preizku{ancih tipa II: a)
brez razpok na nivoju napetosti 357 MPa, b) razpoke, ki se ne {irijo na
nivoju napetosti 405 MPa
Figure 3: Non-propagating short cracks in the type-I specimens after
107 cycles: a) individual cracks at the stress level of 708 MPa, b) a net
of cracks at the stress level of 760 MPa
Slika 3: Kratke razpoke, ki se ne {irijo, pri preizku{ancih tipa I po 107
ciklih: a) posamezne razpoke na nivoju napetosti 708 MPa, b) mre`a
razpok na nivoju napetosti 760 MPa
bit above the fatigue limit. The crack lengths at different
stress levels and at different numbers of stress cycles
were measured using microscopes.
The registered crack lengths, i.e., the sum of the
single crack and the hole diameter or the sum of two
cracks and the hole diameter are presented in Figure 6
depending on the specific number of the cycles.
3 RESULTS AND DISCUSSION
3.1 Fatigue limit
The summarized data on the fatigue limit, the stress
level for the crack initiation and the non-propagating
crack size registered during the testing as well as the
parameter area of the drilled holes and the information
about the local residual-stress field are listed in Table 3.
The surface stress at the bottom of the groove corres-
ponds to the nominal bending stress multiplied by the
stress-concentration coefficient.
The fatigue limit of defected specimens is lower than
the fatigue limit of smooth specimens. Tiny holes help
crack initiation. Cracks appear at lower stress levels. The
reason for it is the stress concentration caused by a hole.
Table 3: Test results
Tabela 3: Rezultati preizku{anja
Specimen
Fatigue
limit
(MPa)
Stress
level for
initiation
(MPa)
Non-prop
agating
crack size
(μm)
Parameter
area
(μm)
Residual
stress
Type I 760 691 >150 0 none
Type II 706 381 136 56.2 none
Type III 678 517 112 58.0 present
V. GLIHA et al.: BEHAVIOUR OF SHORT CRACKS EMANATING FROM TINY DRILLED HOLES
444 Materiali in tehnologije / Materials and technology 47 (2013) 4, 441–446
Figure 6: Dependences of crack lengths, a, upon the number of stress
cycles, N, at different stress levels, : a) 697 MPa, b) 678 MPa, c) 677
MPa, d) 522 MPa
Slika 6: Odvisnost dol`in razpoke a od {tevila ciklov napetosti N na
razli~nih nivojih napetosti : a) 697 MPa, b) 678 MPa, c) 677 MPa, d)
522 MPa
Figure 5: Two different stress levels applied to the type-III specimens:
a) non-propagating crack at the stress level of 565 MPa after 107
cycles, b) propagating cracks at the stress level of 685 MPa after 1.27
106 cycles
Slika 5: Dva razli~na nivoja napetosti pri preizku{ancih tipa III: a)
razpoka, ki se ne {iri, na nivoju napetosti 565 MPa po 107 ciklih, b)
razpoki, ki se {irita, na nivoju napetosti 685 MPa po 1,27 ·106 ciklih
Figure 7: Course of the circular stress induced by drilling , against
the distance, r (schematically)
Slika 7: Potek kro`ne napetosti , ki je nastala z vrtanjem, glede na
razdaljo r (shematsko)
In respect to the residual-stress field induced by drill-
ing the defected specimens behave differently. Namely,
the residual stresses are compressive adjacent to the hole.
Residual stresses are self-balancing forces. Compressive
stresses are in the equilibrium with tensile ones. There-
fore, the compressive residual stress decreases with the
distance and turns to the tensile stress. The tensile stress
attains its maximum at a certain distance and then it
decreases. The sketch of the residual stresses induced by
hole-drilling is shown in Figure 7.
The stress level for the crack initiation is the lowest
for the type-III specimens. The largest non-propagating
crack is the shortest in the type-III specimens, too. The
reason can only be the existing residual-stress field.
Compressive stresses retard the crack initiation.
3.2 Anomalous fast short-crack propagation
After the initiation, the short-crack propagation is
anomalous fast, much faster than the propagation of the
long cracks close to the threshold.
Between two successive moments during the cyclic
loading of the specimens when the crack-length measu-
rement was performed (Figure 6), the average crack-
propagation rate was calculated as:
d
d
a
N
a a
N N
i i
i i
=
−
−
−
−
1
1
(1)
V. GLIHA et al.: BEHAVIOUR OF SHORT CRACKS EMANATING FROM TINY DRILLED HOLES
Materiali in tehnologije / Materials and technology 47 (2013) 4, 441–446 445
Figure 9: Dependences of the crack-propagation rate, da/dN, upon the
stress-intensity-factor range, K, for different stress levels, : a) 697
MPa, b) 678 MPa, c) 677 MPa, d) 522 MPa
Slika 9: Odvisnost hitrosti {irjenja razpoke da/dN od razpona faktorja
intenzitete napetosti K pri razli~nih nivojih napetosti : a) 697 MPa,
b) 678 MPa, c) 677 MPa, d) 522 MPa
Figure 8: Dependences of the crack-propagation rate, da/dN, upon the
crack size, a, for different stress levels, : a)  = 697 MPa, b)  = 678
MPa, c)  = 677 MPa, d)  = 522 MPa
Slika 8: Odvisnost hitrosti {irjenja razpoke da/dN od velikosti razpoke
a pri razli~nih nivojih napetosti : a) 697 MPa, b) 678 MPa, c) 677
MPa, d) 522 MPa
Figure 10: Comparison of short-crack and long-crack propagation
rates
Slika 10: Primerjava hitrosti {irjenja kratkih in dolgih razpok
The results of the calculations shown against the
crack size are presented in Figure 8.
As expected, the short-crack-propagation rate de-
creases with the crack length. At the stress levels lower
than the fatigue limit, the cracks sooner or later stop pro-
pagating. The crack at the stress level equal to the fatigue
limit is the largest non-propagating crack.
However, the driving force for the crack propagation
is the stress-intensity-factor range, K:4,6
Δ ΔK a= ⋅ ⋅ ⋅065.  π (2)
The crack-propagation rates against the stress-inten-
sity-factor range are plotted in Figure 9.
The comparison of all the short-crack-propagation
rates with the long-crack-propagation rate7,8 close to the
threshold value for the long-crack propagation and above
it are shown in Figure 10. The largest non-propagating
crack in the type-III specimens is 112 μm long (the sum
of the cracks and the hole diameter is 224 μm).
4 CONCLUSIONS
Immediately after the initiation the driving force for
the short-crack propagation is not equal to the whole
stress-intensity-factor range, K, but to its effective
range, Keff. The reason is the crack-closure effect. As a
short crack approaches the grain boundary, the pro-
pagation rate is gradually retarded because Keff
decreases. If the stress level of the applied cyclic stress is
lower than the fatigue limit, i.e., Keff is smaller than the
threshold stress-intensity-factor range, the crack stops
propagating and becomes a non-propagating crack.9 In
contrast, if Keff is bigger than the threshold value, the
crack will continue to propagate as a long crack after
retardation during the short-crack propagation.
The largest non-propagating crack in the type-I
specimens is longer than in the type-II specimens. Both
specimen types are in the residual-stress-free condition.
A drilled hole makes a crack initiation easy. The regi-
stered crack in a type-I specimen is the largest crack
spreading along the whole circumference of the speci-
men, while the crack in a type-II specimen was forced to
emanate from the drilled hole. It is expected that the
condition for the unimpeded short-crack propagation is
not so convenient in the type-II specimens because
short-crack propagation depends on the grain compo-
sition (size, orientation etc).
The reason why the non-propagating cracks in the
type-III specimens are not as large as the ones in the
type-II specimens is the stress field induced by hole-
drilling. At the stress levels much lower than the fatigue
limit, a short crack begins to propagate through the
domain of compressive residual stresses. These levels
lower Keff as well as the short-crack propagation rate.
Due to the decreasing compressive residual stress during
the crack propagation, the crack closure decreases and
Keff increases. The result is the quickly increasing
short-crack propagation. Despite the decelerated initial
crack propagation due to the compressive residual-stress
field, the crack, sooner or later, propagates to the domain
of the tensile residual stress. Then the crack propagates
faster. The largest non-propagating crack in the type-III
specimens is therefore smaller than in the type-II speci-
mens.
5 REFERENCES
1 K. J. Miller, The behavior of short fatigue cracks and their initiation
1. A review of 2 recent books, Fatigue & Fracture of Engineering
Materials & Structures, 10 (1987) 1, 75–91
2 K. J. Miller, The behavior of short fatigue cracks and their initiation
2. A general summary, Fatigue & Fracture of Engineering Materials
& Structures, 10 (1987) 2, 93–113
3 Y. Murakami, M. Endo, Effects of defects, inclusions and inhomo-
geneities on fatigue-strength, International Journal of Fatigue, 16
(1994) 3, 163–182
4 Y. Murakami, M. Endo, Quantitative evaluation of fatigue strength of
metals containing various small defects or cracks, Engineering
Fracture Mechanics, 17 (1983), 1–15
5 Y. Murakami, M. Endo, Effect of hardness and Crack geometries on
Kth of small cracks estimating from small defect, In K. J. Miller, E.
R. de los Rios, editors, Proceedings, The behaviour of short fatigue
cracks, Mechanical Engineering Publications, London 1986,
275–294
6 Y. Murakami, K. J. Miller, What is fatigue damage? A view point
from the observation of low cycle fatigue process, International
Journal of Fatigue, 27 (2005) 3, 991–1005
7 T. Vuherer, L. Milovi}, V. Gliha, Behaviour of small cracks during
their propagation from Vickers indentations in coarse-grain steel: An
experimental investigation, International Journal of Fatigue, 33
(2011) 12, 1505–1513
8 T. Vuherer, Analysis of influence of micro defects on fatigue strength
of coarse grain HAZ in welds, Doctor Thesis, University of Maribor,
Faculty of Mechanical Engineering, 2008
9 S. Suresh, Fatigue of materials, second edition, Cambridge Univer-
sity Press, Cambridge 1998, 343–345, 541–569
V. GLIHA et al.: BEHAVIOUR OF SHORT CRACKS EMANATING FROM TINY DRILLED HOLES
446 Materiali in tehnologije / Materials and technology 47 (2013) 4, 441–446
L. BALANOVI] et al.: INTERNAL-OXIDATION KINETICS OF Ag-Cd ALLOYS
INTERNAL-OXIDATION KINETICS OF Ag-Cd ALLOYS
KINETIKA NOTRANJE OKSIDACIJE ZLITIN Ag-Cd
Ljubi{a Balanovi}1, Vladan ]osovi}2, Nade`da Talijan2, Dragana @ivkovi}1
1University of Belgrade, Technical Faculty, Bor, Serbia
2Institute of Chemistry, Technology and Metallurgy, University of Belgrade, Belgrade, Serbia
ljbalanovic@tf.bor.ac.rs
Prejem rokopisa – received: 2012-10-15; sprejem za objavo – accepted for publication: 2012-12-11
The oxidation kinetics of the Ag-Cd alloys with different Cd contents in mass fractions (9.5, 11, 12 and 16) % at (670, 750 and
795) °C during the production of Ag-CdO electrical-contact materials was investigated. It was found that the internal oxidation
of the investigated Ag-Cd alloys generally follows the parabolic-rate law for all the applied oxidation times (5, 9.5, 20.5, 36 and
48) h. The parabolic-rate-law constant for the measured oxide-thickness gain was evaluated at every examined temperature and
the Arrhenius-type equations were determined in order to describe the temperature dependence of the rate constants.
Microstructures of the obtained Ag-CdO electrical-contact materials were investigated using a SEM-EDS analysis and the
microstructure of a sample produced under the optimum process conditions is illustrated with the corresponding micrographs
and chemical compositions.
Keywords: Ag-Cd oxidation kinetics, internal oxidation, electrical-contact materials, microstructure
Med proizvodnjo materialov Ag-CdO za elektri~ne kontakte je bila preiskovana kinetika oksidacije zlitin Ag-Cd z razli~no
vsebnostjo masnih dele`ev Cd (9,5, 11, 12 in 16) % pri (670, 750 in 795) °C. Ugotovljeno je bilo, da se notranja oksidacija
preiskovanih zlitin Ag-Cd ujema z zakonom paraboli~ne odvisnosti hitrosti pri vseh ~asih oksidacije (5; 9,5; 20,5; 36 in 48) h.
Konstanta paraboli~ne odvisnosti hitrosti nara{~anja debeline oksida je bila dolo~ena pri vsaki temperaturi preizkusa, prav tako
je bila dolo~ena Arrheniusova vrsta ena~be za opis temperaturne odvisnosti konstant hitrosti. Mikrostruktura dobljenega
materiala Ag-CdO za elektri~ne kontakte je bila preiskovana s SEM-EDS-analizami. Prikazana je bila mikrostruktura in
narejena kemijska analiza vzorca, izdelanega v optimalnih razmerah.
Klju~ne besede: kinetika oksidacije Ag-Cd, notranja oksidacija, materiali za elektri~ne kontakte, mikrostruktura
1 INTRODUCTION
In the electrical industry most of the contact appli-
cations use the materials based on silver that include a
combination of pure metals, alloys and metal powders.
Silver is also used as a plated, brazed, or mechanically
bonded overlay on copper and copper-based materials
due to its high thermal and electrical conductivities.
When considering that electrical-contact materials
are used in diverse service conditions and also that no
metal has all the properties required to accomplish the
objectives of different contact applications, in addition to
the silver-type contacts, other types of contacts based on
platinum-group metals, tungsten, molybdenum, copper,
copper alloys and mercury can be found in the electrical
industry.
The usefulness of an electrical-contact material also
depends on a variety of electrical and mechanical
properties, service life, load conditions and economical
reasons.1–5
The Ag-Cd groups of the contact alloys are widely
used in tonnage quantities in the electrical and elec-
tronics industries because of their high electrical and
thermal conductivities, high resistance to arcing, high
welding-adhesion resistance, low-contact resistance,
high hardness and strength. Cadmium improves the
arc-quenching ability of silver and also increases its
resistivity and mechanical strength.6–8
Silver cadmium oxide can be obtained using internal
oxidation and powder metallurgy, and the manufacturing
method has a significant influence on the properties and
microstructure of this material.9
In the case of the internal-oxidation process, oxygen
diffuses into an alloy and causes a sub-surface precipi-
tation of the oxides of one or more alloying elements,
which was the subject of the reviews by Birks et al.10
The process of internal oxidation occurs in the
following manner: oxygen dissolves in the base metal
and diffuses inward through the metal matrix containing
previously precipitated oxide particles. The critical
activity product for the nucleation of the precipitates is
established at the reaction front (parallel to the specimen
surface) with the inward-diffusing oxygen and the out-
ward diffusion of the solute. When the rate-controlling
step in the oxidation process is the diffusion of ions
through a compact barrier layer of the oxide with the
chemical potential gradient as the driving force, the
parabolic-rate law is usually observed.11 As the oxide
grows thicker, the diffusion distance increases and the
oxidation rate slows down.
The results of the investigation of the oxidation-
layer-thickness dependence on the temperature and time
of the oxidation for the Ag-Cd alloys with different
contents of Cd are presented in this paper. Previous
investigations have shown that the oxidation process in
this alloy has a parabolic behavior according to the
Materiali in tehnologije / Materials and technology 47 (2013) 4, 447–452 447
UDK 66.094.3:669.018.5 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)447(2013)
theoretical relation between the parabolic-rate constant
of the metal oxidation and the self-diffusion coefficients
in the oxide layer growing on the metal surface, given by
Wagner.12 But, Wagner’s general equation in its original
form could only be applied to the cases, in which the
dependence of the self-diffusion coefficient on the
equilibrium partial pressure of the oxidizer was known a
priori.
Therefore, since the thickness of the oxide layer
depends on the oxidation rate and the oxidation rate is
directly correlated with the temperature, the oxidation
time and the oxygen pressure, the Arrhenius equation,
being widely used to describe the temperature depen-
dence of the kinetic constants, was used for determining
the activation energy for the oxidation processes in the
studied Ag-Cd alloys.
2 EXPERIMENTAL PROCEDURE
Silver cadmium alloys with mass fractions (9.5, 11,
12 and 16) % of Cd were prepared using the Ag and Cd
metals of 99.99 % purity. The alloys were first cast into
ingots and then plastically deformed, at low tempera-
tures, into the strips with a thickness of 2 mm. The
prepared silver cadmium alloys were heated in an
electro-resistive oven in air atmosphere under natural
convection at the temperatures of 670 °C, 750 °C and
795 °C during (5; 9.5; 20.5; 36 and 48) h. The micro-
structure observation and measurements of the thickness
of the oxidized layer were done on the polished
cross-sections of the investigated samples using a
scanning electron microscope JEOL-JSM-6610LV with
an energy dispersive spectrometer (EDS) and a metallo-
graphic microscope REICHART of the POLYVAR –
MET type.
3 RESULTS AND DISCUSSION
3.1 Oxidation rate
The samples were measured before and after the
internal oxidation. The graphic dependence of the
relative change in the weight after (5, 9.5, 20.5, 36 and
48) h of the oxidation of the silver cadmium alloys with
mass fractions (9.5, 11, 12 and 16) % of Cd is presented
in Figure 1a. The inner oxidation of the Ag alloys with
different Cd contents generally follows the parabolic-rate
law for the examined time range up to 48 h at the tempe-
rature of 750 °C. In Figure 1b it can be seen that 60 %
of Cd was oxidized after a 20.5 h oxidation of the silver
cadmium alloy with 9 % Cd.
3.2 Oxidation kinetics
During the internal oxidation of the silver cadmium
alloy, the cadmium species become depleted in the zones
when the oxygen front moves into the silver cadmium
alloy. As the oxidation front moves from the surface of
the strip toward the center, the concentration of the cad-
mium species becomes increasingly dilute as compared
to the original composition. Hence, after the oxidation is
completed, the cross-section will display a significant
oxide-deficient or oxide-depleted zone in the center of
the contact body. For some applications the presence of
the depletion zone is detrimental, requiring its removal
or displacement from the center. There are two common
methods to achieve such a result. The first method takes
account of the fact that an oxidation barrier, such as a
ceramic glaze, is applied to one surface so that the
oxidation can proceed only from one side. The second
method is to laminate two silver cadmium sheets of the
same size and to form a package by welding them along
all four edges. After the oxidation, the sheets are sepa-
rated. The oxide-deficient zone will appear on one side
(the inner side of the package) of each sheet.5,12
Figure 2 shows the dependence of the thickness of an
oxidized layer for different Cd contents in the Ag-Cd
alloys. After comparing the changes in the thickness of
the oxidized layer with the temperature and time, it is
obvious that the oxidation rate is the highest for the alloy
with the smallest amount of Cd and the lowest for the
alloy with largest amount of Cd.
L. BALANOVI] et al.: INTERNAL-OXIDATION KINETICS OF Ag-Cd ALLOYS
448 Materiali in tehnologije / Materials and technology 47 (2013) 4, 447–452
Figure 1: a) Dependence of the relative mass change during the
oxidizing for the Ag-Cd alloys with different Cd contents at 750 °C;
b) extent of the Cd oxidation for the Ag-Cd alloys with 9 % and 16 %
Cd at 750 °C
Slika 1: a) Odvisnost relativne spremembe mase med oksidacijo zlitin
Ag-Cd z razli~no vsebnostjo Cd pri 750 °C, b) dele` oksidacije Cd v
Ag-Cd zlitinah z 9 % in 16 % Cd pri 750 °C
A mass content of 9–12 % of CdO in the Ag-CdO
alloys successfully creates a decrease in the deformation
and devastation of electrical contacts. With the
concentrations of CdO above 12 %, an inner break is
possible due to an intension caused by an increase in the
grain diameter after the oxidation. The main difference is
that the elevated temperatures promote ionic diffusion
and, thus, the oxide formation can proceed to a much
greater extent than at the low temperatures where only
thin layers are formed.
The dependence of the scale thickness on the time
can be used for determining the kinetic law for the
oxidation rate. As the oxidized layer grows thicker, the
diffusion distance increases and the oxidation rate slows
down. The rate (dx/dt) is inversely proportional to the
oxide thickness (x) or:
d
d
px
t
k
x
= (1)
Equation (1) can be rewritten as:
x x k tpd d= (2)
Upon an integration of equation (2):
x x k t
x
x x
t
t t
d dp
=
=
=
=
∫ ∫=
0 0
(3)
On integration, the parabolic equation is obtained:
L. BALANOVI] et al.: INTERNAL-OXIDATION KINETICS OF Ag-Cd ALLOYS
Materiali in tehnologije / Materials and technology 47 (2013) 4, 447–452 449
Figure 2: Dependence of the thickness of an oxidized layer for Ag-Cd
alloys with different Cd contents; inner oxidation at: a) 670 °C, b) 750 °C
and c) 795 °C
Slika 2: Odvisnost debeline oksidirane plasti pri zlitinah Ag-Cd z
razli~no vsebnostjo Cd; notranja oksidacija pri: a) 670 °C, b) 750 °C
in c) 795 °C
Figure 3: Logarithm of the oxide thickness versus the logarithm of the
oxidation time for the Ag-Cd alloys with different Cd contents at: a)
670 °C, b) 750 °C and c) 795 °C
Slika 3: Logaritem debeline oksida v primerjavi z logaritmom ~asa
oksidacije za zlitine Ag-Cd z razli~no vsebnostjo Cd pri: a) 670 °C,
b) 750 °C in c) 795 °C
x k t c2 2= +p (4)
and the parabolic-rate constant, kp, can be determined
from the experiment in the units of cm2 s–1 or a similar
length squared per time units:12,13
If one takes the logarithm of both sides of equation
(4), the following is obtained:
lg x2 = lg kp + lg t (5)
A plot of lg x2 versus lg t can be done linearly and the
parabolic-rate constants can be determined for the
reaction carried out at several temperatures. The inter-
cept of the line is lg kp.
The logarithm of the oxide thickness as a function of
the logarithm of the oxidation time for the Ag-Cd alloys
with different Cd contents at the investigated tempera-
tures is presented in Figure 3, while the dependence of
the parabolic-rate constant as a function of the inverse
temperature oxidation of the Ag-Cd alloys with different
Cd contents is shown in Figure 4.
This indicates that the thermal diffusion process is
rate controlling, with oxygen or cations or both diffusing
through a compact layer. The rate of diffusion through an
oxide film depends on a number of factors, such as the
temperature, oxygen partial pressure and structure of the
oxide.12–14
At the temperatures higher than the melting point of a
metal, the lattice diffusion dominates through the
crystalline oxide formed on the metal and, at the lower
temperatures, the diffusion via the oxide-grain boun-
daries is predominant. In this case, the rate of oxidation
of a metal or alloy depends on the oxide-grain size,
which is often dictated by the substrate-grain orientation,
surface pretreatment, etc.9 A deviation from the para-
bolic-oxidation behavior is often observed and can be a
result of the oxide-grain-size change with time and at a
particular temperature. In this case, the number of
oxide-grain-boundary šeasy diffusion paths’ decreases
with time, causing an apparent decrease in the oxidation
rate. When the rate-controlling step in the oxidation
process is a diffusion of ions through a compact barrier
layer of oxide with the chemical potential gradient as the
driving force, the parabolic-rate law is usually observed15
and then the change in the oxidation rate with the
temperature will follow the Arrhenius equation:11,16–18
k Ae
E
RT
p
a
=
−
(6)
where k is the rate constant, A is the frequency factor (or
pre-exponential factor), R is the molar gas constant, Ea
is the activation energy, and T is the temperature (K). If
we take the natural logarithm of both sides and the
rearrangement, this equation can be put in the linear
form:
ln lnk A
E
RTp
a= − (7)
so that the constants can be calculated using a linear
regression. This type of linearization, which allows
fitting a straight line into the transformed experimental
data, has been used for nearly a century.19–23 Most of the
constants reported in the literature were obtained using
this method. And therefore, a plot of ln kp versus 1/T can
be made, or a linear regression performed, after the rate
constants have been determined for a reaction carried
out at several temperatures. The Arrhenius plot is shown
in Figure 5. For a particular reaction, the following rate
constants were obtained when the reaction was studied
at a series of temperatures.
Using the data presented in Figure 4 and the
Arrhenius plot of the parabolic-rate constants for the
oxidation of the Ag-Cd alloy, shown in Figure 5, the
activation energy was calculated as 36 ± 2 kJ/mol and
shown in Table 1.
L. BALANOVI] et al.: INTERNAL-OXIDATION KINETICS OF Ag-Cd ALLOYS
450 Materiali in tehnologije / Materials and technology 47 (2013) 4, 447–452
Figure 5: Logarithm of the parabolic-rate constants versus the inverse
temperature for the oxidation of the Ag-Cd alloys
Slika 5: Logaritem konstant paraboli~nih hitrosti v primerjavi z
inverzno vrednostjo temperature pri oksidaciji zlitin Ag-Cd
Figure 4: Dependence of parabolic-rate constants versus temperature
oxidation of the Ag-Cd alloys with different Cd contents
Slika 4: Odvisnost konstant paraboli~nih hitrosti od temperature
oksidacije zlitin Ag-Cd z razli~no vsebnostjo Cd
Table 1: Parabolic-rate constants and activation energy of the
oxidation of the Ag-Cd alloys with different Cd contents
Tabela 1: Konstante paraboli~ne hitrosti in aktivacijska energija zlitin
Ag-Cd z razli~no vsebnostjo Cd
Material
compositi-
on in mass
fractions
K/(mm2 h)
Activation
energy of
oxidation
kJ/mol670 °C 750 °C 795 °C
16 % Cd 1.52 · 10–3 3.66 · 10–3 4.23 · 10–3 38
12 % Cd 2.90 · 10–3 5.83 · 10–3 8.00 · 10–3 36
11 % Cd 3.41 · 10–3 7.20 · 10–3 9.41 · 10–3 38
9.5 % Cd 4.02 · 10–3 8.40 · 10–3 10.92 · 10–3 36
Considering that the anti-welding behavior and wear
resistance of the electrical-contact materials can be
improved with a uniform dispersion of the metal oxide
particles in a soft silver matrix, the ultimate goal of the
investigations is an optimization of the internal-oxidation
process parameters that promote the formation of the
optimum microstructure, i.e., obtaining of a good and
homogeneous oxide dispersion within the silver matrix.
As an illustration of the optimum microstructure of
the Ag-CdO electrical-contact materials obtained in the
process investigating the internal oxidation, a characte-
ristic SEM micrograph of the Ag-CdO composite
obtained with the internal oxidation of the Ag-11 % Cd
alloy (Ag-CdO11) is given in Figure 6.
The results of a further microstructural analysis using
EDS are presented in Table 2 and the selected points
used for the analysis are marked in Figure 7.
Table 2: Results of the EDS analysis of the Ag-11 % Cd alloy
(Ag-CdO11) for the electrical-contact material
Tabela 2: Rezultati EDS-analize zlitine Ag-11 % Cd (Ag-CdO11) za
elektri~ne kontakte
Spectrum w(Ag)/% w(Cd)/ % w(O)/ %
1 0 83.55 16.45
2 4.99 78.76 16.25
3 97.92 2.08 0
4 97.91 2.09 0
Figures 6 and 7 display a microstructure of the
material upon the completion of the oxidation process,
which consists of the CdO particles embedded in a silver
matrix. It should be noted that in order to provide a good
bond between the contact material and the contact holder
by soldering, about 25 % of the total thickness of the
electrical-contact material must not be oxidized, as
illustrated in Figure 6. The presented EDS results (Table
2) corroborate the observed microstructure and illustrate
the Ag and CdO-rich regions.
4 CONCLUSION
The present study on the oxidation behavior of the
Ag-Cd alloys with different contents of Cd, carried out
in the air environment at the temperatures of 670 °C, 750
°C and 795 °C during (5, 9.5, 20.5, 36 and 48) h using a
L. BALANOVI] et al.: INTERNAL-OXIDATION KINETICS OF Ag-Cd ALLOYS
Materiali in tehnologije / Materials and technology 47 (2013) 4, 447–452 451
Figure 7: SEM micrograph of a cross-section of the Ag-11 mass %
Cd alloy (Ag-CdO11) with the annotated points of the EDS analysis
Slika 7: SEM-posnetek preseka zlitine Ag-11 % Cd (Ag-CdO11) z
ozna~enimi to~kami, kjer je bila izvr{ena EDS-analiza
Figure 6: SEM micrograph of an Ag-CdO composite obtained with
the internal oxidation of the Ag-11 % Cd alloy (Ag-CdO11); a) pure
Ag-Cd and an oxide layer of Ag-CdO; b) oxide layer of Ag-CdO
Slika 6: SEM-posnetek kompozita Ag-CdO, dobljenega z notranjo
oksidacijo zlitine Ag-11 % Cd (Ag-CdO11); a) ~isti Ag-Cd in oksidna
plast Ag-CdO; b) oksidna plast Ag-CdO
microstructure observation and measurements of the
thickness of the oxidized layer revealed the following:
The thickness of the oxidized layer during the
heating depends on the following process conditions: the
temperature and duration of oxidation, the amount of
cadmium in the Ag-Cd alloys and the oxygen pressure
on the surface of the alloy.
The oxidation kinetics of the alloys follows the
parabolic-rate law suggesting a diffusion-controlled
growth of the oxide film.
The activation energy, Ea, and the pre-exponential
factor were determined using the Arrhenius equation,
fitting ln kp vs 1/T at the temperatures of 670 °C, 750 °C
and 795 °C. The activation-energy value of 36 ± 2
kJ/mol indicates a mixed effect of diffusion and tem-
perature on the kinetics of the mentioned process of the
internal oxidation of Ag-Cd alloys. This is in accordance
with the observation given in5 that the temperature has
the highest influence on the process of an inner oxida-
tion5 because, at higher temperatures, the inner oxidation
is faster according to Wagner’s expression of the para-
bolic-rate constant for an internal oxidation where the
solute element is part of the denominator.12
The aim of this study is to gain a better basic under-
standing of the controlling processes in the coarsening of
the cadmium oxide phase in the internally oxidized silver
cadmium alloys.
Acknowledgement
This paper was done under the financial support of
the Ministry of Education and Science, Republic of
Serbia – project No 172037.
5 REFERENCES
1 J. R. Davis, ASM Materials Engineering Dictionary, ASM Inter-
national, Materials Park, OH 1992
2 Y. S. Shen, P. Lattari, J. Gardner, H. Wiegard, ASM Handbook,
Powder Metal Tehnologies and Applications, 7 (1998), 1025
3 R. Holm, Electrical Contacts - Theory and applications, 4th ed.,
Springer, 2000
4 M. Braunovi}, N. K. Myshkin, V. V. Konchits, Electrical contacts –
Fundamentals, Applications and Technology, CRC Press, Taylor and
Francis Group, 2007
5 N. Talijan, J. Staji}-Trosi}, V. ]osovi}, A. Gruji}, E. Romhanji, Elec-
trical and mechanical properties of silver based electical contact
materials, Proceedings of the 4th Balkan Conference on Metallurgy,
Zlatibor, 2006, 445
6 P. G. Slade, ed., Electrical Contacts: Principles and Applications,
CRC Press, 1999, 684
7 N. Talijan, J. Staji}-Trosi},V. ]osovi}, A. Gruji}, E. Romhanji, Com-
posite electrical contacts based on silver, Proceedings of the 10th
International Research/Expert Conference Trends in the Develop-
ment of Machinery and Associated Technology, Barcelona, 2006,
1339
8 V. ]osovi}, N. Talijan, D. @ivkovi}, D. Mini}, @. @ivkovi}, Com-
parison of properties of silver-metal oxide electrical contact
materials, J. Min. Metall. Sect. B-Metall., B 48 (2012), 131–141
9 ASTM Standards, Standard Guide for Silver-Cadmium Oxide
Contact Material, B781–93a, 1999
10 N. Birks, G. H. Meier, F. S. Pettit, Introduction to the High Tem-
perature Oxidation of Metals, Cambridge University Press, New
York 2006
11 J. R. Davis, Heat-resistant materials, ASM International, Handbook
Committee, Materials Park, OH 1997
12 C. Wagner, Z. Phys. Chem. B, (1933), 21
13 D. J. Young, High Temperature Oxidation and Corrosion of Metals,
Elsevier, Cambridge 2008
14 J. H. Moore, N. D. Spencer, Encyclopedia of Chemical Physics and
Physical Chemistry, Volumes 1–3, Institute of Physics, 2001
15 A. M. Huntz, Journal of materials science letters, 18 (1999),
1981–1984
16 M. J. Graham, R. J. Hussey, Oxid. Met., 44 (1995), 339
17 S. Arrhenius, Zur Theorie der chemischen Reaktionsgeschwindig-
keiten, Z. Physik. Chem., 7 (1899)
18 W. Jaenicke, S. Leistikow, A. Sadler, J. Electrochem. Soc., 111
(1964), 1031
19 N. B. Pilling, R. E. Bedworth, J. Inst. Met., 29 (1923), 529
20 W. K. Appleby, R. F. Tylecote, Corros. Sci., 10 (1970), 325
21 J. Stringer, Werkstoffe Korros., 23 (1972), 747
22 N. H. Chen, R. Aris, Determination of Arrhenius constants by linear
and nonlinear fitting, AIChE J., 38 (1992) 4, 626–628
23 J. [esták, A. Kozmidis-Petrovi}, @. @ivkovi}, Crystallization kinetics
accountability and the correspondingly developed glass-forming
criteria – A personal recollection at the forty years anniversaries, J.
Min. Metall. Sect. B-Metall. B, 47 (2011) 2, 229–239
L. BALANOVI] et al.: INTERNAL-OXIDATION KINETICS OF Ag-Cd ALLOYS
452 Materiali in tehnologije / Materials and technology 47 (2013) 4, 447–452
F. NEMATZADEH, S. K. SADRNEZHAAD: EFFECTS OF THE AGEING TREATMENT ON THE SUPERELASTIC BEHAVIOR ...
EFFECTS OF THE AGEING TREATMENT ON THE
SUPERELASTIC BEHAVIOR OF A NITINOL STENT FOR
AN APPLICATION IN THE ESOPHAGEAL DUCT: A
FINITE-ELEMENT ANALYSIS
U^INEK STARANJA NA SUPERELASTI^NO VEDENJE
NITINOLNE OPORE PRI UPORABI V CEVI PO@IRALNIKA:
ANALIZA KON^NIH ELEMENTOV
Fardin Nematzadeh1, Seyed K. Sadrnezhaad2
1Materials and Energy Research Center, (MERC), P.O. Box 14155-4777, Tehran, Iran
2Department of Materials Science and Engineering, Sharif University of Technology, P.O. Box 11155-9466, Tehran, Iran
fardinnematzadeh@gmail.com
Prejem rokopisa – received: 2012-10-23; sprejem za objavo – accepted for publication: 2012-12-18
The effects of design parameters and material properties obtained with the ageing treatment on the mechanical performance of a
z-shaped esophageal-duct nitinol wire stent under the crushing tests for clinical applications are investigated with a
finite-element simulation. With 90 % crushing, low chronic outward force, high radial resistive strength and favorable
superelastic behavior are attained at the segment angle of 65° and the Af temperature of 24 °C. The performance of the stent is
seen to drastically vary with a change of only 1° in the segment angle.
Keywords: finite-element analysis, esophageal duct, nitinol stent, ageing treatment, mechanical performance
Z metodo kon~nih elementov so bili preu~evani u~inek konstrukcijskih parametrov in mehanske lastnosti materiala pri tla~nih
preizkusih starane nitinolne `i~ne opornice z-oblike za klini~no uporabo v cevi po`iralnika. Z 90-odstotnim tla~enjem, nizko
stalno silo navzven, veliko radialno silo upornosti in `eleno superplasti~no vedenje je bilo dose`eno pri segmentu s kotom 65° in
temperaturi Af 24 °C. Opazi se, da se vedenje opore ob~utno spremeni `e pri spremembi kota segmenta za 1°.
Klju~ne besede: analiza kon~nih elementov, cev po`iralnika, nitinolna opora, staranje, mehansko delovanje
1 INTRODUCTION
Gastrointestinal disease is an important cause of
death these days. Renteln et al. have referred to the
esophageal cancer as a worldwide source of gastro-
intestinal malignancies.1 The majority of patients with
esophageal cancer suffers from dysphagia and undergoes
death.1 The primary aim of treating these patients is to
reduce dysphagia with a minimum morbidity and
mortality and to improve the quality of their lives.1 With
respect to this type of illness, Renteln et al.2 have stated
that an endoscopic stenting is better than a surgical
repair. A stent placement has, therefore, been one of the
main remedies for these types of diseases in the last
decade.
The use of a stent has two main objectives: (1) a
short-term effect of avoiding intimal dissection and an
elastic recoil and (2) a long-term effect of avoiding
restenosis caused by neointimal hyperplasia mentioned
by Stoeckel et al.3 A nitinol-stent placement was deve-
loped to introduce a behavioral modality of the palliation
of malignant dysphagia. Nitinol stents for the esophageal
duct are easily implanted with a low risk of a severe
complication. They provide a successful palliation of
malignant esophageal obstructions, too. They also help
relieve dysphagia for the patients with inoperable carci-
noma esophagus.1
A stent placement has been the major approach to
solving gastrointestinal diseases like esophageal malig-
nancy during the past decade. Nitinol superelastic stents
have been considered a solution to the problems such as
restenosis after an implantation, a low twisting ability, an
unsatisfactory radial mechanical strength and improper
dynamic behaviors associated with the ducts.
Owing to a good retrievability and flexibility, the
z-shaped wire stents are most widely used in the stent
designs.4 They can be used to fabricate custom-made
stents with preselected values exerting the radial forces
for the clinical use. The Z-shaped models are also desi-
rable as, even in a laboratory, they can be easily manu-
factured by hand. The Z-shaped models also allow
various designs with different amounts of radial forces as
described by Patrick et al.5 Important parameters like the
length, diameter of the wire, number of bends, segment
angle, stent inner diameter and radial contraction should
be determined with analytical equations associated with
the geometry of a stent.
The first report about a numerical study of the fatigue
behavior of a nitinol stent was presented by Whitcher et
al.6 A good agreement of their numerical results, calcu-
Materiali in tehnologije / Materials and technology 47 (2013) 4, 453–459 453
UDK 519.61/.64:620.17:66.017 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)453(2013)
lated for the crushing tests, with the other researchers’
experimental data was assessed by Petrini et al.7, Klein-
streuer et al.8 have offered a computational analysis of
different nitinol stent-graft combinations for the support
of the abdominal aortic aneurysm (AAA). Beule et al.9
have developed the strategies to investigate and optimize
the mechanics of braided stents. Silber et al.10 have
shown the effect of changing geometrical characteristics
on the mechanical properties of the nitinol wire stents.
The finite-element method is useful for designing the
knitted nitinol meshes utilized in a prospective external
vein reinforcement.11
Liu et al.12 have performed critical testing for deter-
mining the pseudoelastic behavior and transformation
temperatures of the near-equiatomic NiTi shape-memory
alloys. Duerig et al.4 have compared the mechanical
performance of self-expanding stents usable for the treat-
ment of a vascular disease. It has been shown that the
NiTi Af temperature has a great influence on the mecha-
nical performance of self-expandable stents. Yeung et
al.13 have shown that the Af temperature of a nitinol alloy
can be controlled by manipulating the heat-treatment
parameters. This is important for the production of the
nitinol stents with favorable superelastic behavior. Patel
et al.14 have studied the effect of the active Af tempe-
ratures on the fatigue properties of nitinol. They have
shown that a low Af temperature has a short fatigue-life
consequence. Liu et al.15 have evaluated the effect of age
treatment on the shape-setting and superelasticity of the
Ti-50.7 % Ni stents. They have shown that the optimum
ageing temperature for the production of the expected
shape of a stent is 500 °C with an ageing time of no
more than 60 min resulting in an excellent non-linear
superelasticity with the maximum recoverability at the
body temperature. To achieve the superelastic behavior
of a stent, Af should be lower than the body temperature.
Its optimum value depends on the stiffness needed for a
removal of a duct obstruction. A low Af causes problems
like high stiffness and augmented radial forces that can
lead to an injury and scars on the duct.
Although there is some information available on the
nitinol wire stents, to the best of our knowledge, the
z-shaped wire stents for esophageal ducts have not been
studied before. The purpose of this investigation is,
hence, an application of the finite-element method for an
elucidation of the effects of the design parameters and
material properties obtained from the ageing treatment
on the mechanical performance of the nitinol stents
designed for esophageal ducts. The effects of important
parameters are predicted via the model calculations and
are verified with the experimental information available
in the literature. The model is developed on the basis of
the nonlinear 3D finite-element method. An application
of the results can help diminish dysphagia, improving the
quality of patients’ life.
2 MODELING AND METHODS
2.1 Geometric models
Micro-CT can be used for obtaining the optimum
geometry of the commercial stents available on the
market. This instrument is, however, not useful for
designing innovative new designs. The geometry of the
stent used in this research was, therefore, generated by a
computer-aided three-dimensional interactive application
(Catia v.5 (Dassault Systèmes, USA)) and transformed
into a finite-element-analysis code. The geometric para-
meters of the esophageal stent are sketched in Figure 1
and quantified in Table 1. The geometric parameters
shown in Figure 1 and Table 1 were based on the age
treatment of the Ti-50.7 % Ni alloy at 500 °C for
respective durations of 60 min and 30 min according to
the clinical reports given in the literature.1,15
Table 1: Geometric values for the stents of this research (unit of
length is mm)
Tabela 1: Geometrijske vrednosti za opore v tej raziskavi (enota
dol`ine je mm)
Sample
Internal
diameter
of stent
(D)
Radius of
curvature
(
)
Segment
length
(L)
Diameter
of wire
(d)
Bending
angle
(
)
Sample 1 25.0 0.3 10.03 0.2 66°
Sample 2 25.0 0.3 10.80 0.2 65°
2.2 Material properties
The properties of the nitinol stents considered for
esophageal ducts are evaluated with a crushing test.
Superelastic wire properties are merely considered in an
evaluation of the clinical behavior of the nitinol stents
used in this research. The model first developed by
Auricchio et al.16 and Lubliner et al.17 and then
significantly extended by Rebelo et al.18 is used for the
simulation of the nitinol superelastic behavior. The
model is based on a generalization of the theory of
plasticity and the second law of thermodynamics written
in terms of Helmholtz free energy. In this theory, the
strain is decomposed into two components:
Δ Δ Δ  = +el tr (1)
F. NEMATZADEH, S. K. SADRNEZHAAD: EFFECTS OF THE AGEING TREATMENT ON THE SUPERELASTIC BEHAVIOR ...
454 Materiali in tehnologije / Materials and technology 47 (2013) 4, 453–459
Figure 1: Geometric parameters of a Z-shaped esophageal stent
Slika 1: Geometrijski parametri Z-oblike opornice za cev po`iralnika
in which Δ el is the elastic strain and Δ tr is the trans-
formation strain. The transformation of austenite to
twinned martensite occurs due to the motion of the
shear forces taking place in the stress-threshold range of
the superelastic material according to the following
equations:
F F FS F≤ ≤ (2)
in which F is the transformation potential and S and F
denote the martensite transformation start and finish,
respectively;
Δ Δ 

tr = a
F∂
∂
(3)
Δ Δ = f F( ,σ (4)
F p CT= − 	tan + (5)
in which a is the coefficient of strain,  is the martensite
percentage,  is the von Misses stress,  is the von
Misses equivalent stress, p is the pressure, 	 and C are
the material constants and T is the temperature. A
similar approach is applied to define the reverse
transformation which takes into account different stress
thresholds. Equations 3 and 4 define the transformation
intensity. Any change in the stress direction generates a
martensite reorientation with a negligible additional
attempt. Furthermore, the model includes a linear shift-
ing of the stress thresholds with the temperature. In
view of the fact that there is a volume increase asso-
ciated with the transformation, less stress is needed to
generate a transformation in tension and, specifically, in
compression. This effect is modeled with the linear
Drucker-Prager approach for the transformation
potential shown in equation 5 and referred to by Rebelo
et al.18 The parameters required in the Abaqus 6.10
(Dassault Systèmes, Providence, RI, USA) user-material
subroutine for the nitinol material based on the
Auricchio model for opening the esophageal duct are
summarized in Table 2.12,15,19 The basis for the mate-
rial-property selection of the stents is the age treatment
results of the Ti-50.7 % Ni alloy at 500 °C for 30 min
and 60 min. The data given for the sample No. 1 from
Table 2 is based on the age treatment of the Ti-50.7 %
Ni alloy at 500 °C for 30 min. The data for the sample
No. 2 from Table 2 is based on the age treatment of the
Ti-50.7 % Ni alloy at 500 °C for 60 min. A typical
nitinol superelastic behavior employed in this research
is illustrated in Figure 2. Before testing, a cubic
element of nitinol is considered and its results are
compared with the experimental data. Based on the
comparative experimental and mathematical curves
plotted in Figure 3, it is recognized that the Auricchio
model is in reasonable agreement with the empirical
data and, consequently, hereafter, the properties of the
materials are defined in a subroutine based on the
aforementioned numerical model.
2.3 Meshing and boundary conditions
Hypermesh (Altair® Hypermesh® v. 6.0) is a high-
performance finite-element preprocessor that provides
highly interactive software for the performance analysis
of a product design. It is noticeable that due to the mesh-
ing problems caused by a small section of the nitinol
wire and a relatively complex geometry of the stents, the
hypermesh software is used to mesh the samples. The
F. NEMATZADEH, S. K. SADRNEZHAAD: EFFECTS OF THE AGEING TREATMENT ON THE SUPERELASTIC BEHAVIOR ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 453–459 455
Figure 2: Typical behavior of superelastic nitinol
Slika 2: Zna~ilno vedenje superelasti~nega nitinola
Table 2: Material properties used in the simulation of an esophageal-
duct opening with a nitinol stent. The data are based on the Auricchio
model.12,15,19
Tabela 2: Lastnosti materiala, uporabljenega v simulaciji odpiranja
nitinolne opore po`iralnika. Podatki temeljijo na Auricchiovem mo-
delu.12,15,19
Symbol Description Sample No.1
Sample No.
2
EA Austenite elasticity 24100 MPa 20700 MPa
A
Austenite Poisson’s
ratio 0.33 0.33
EM Martensite elasticity 17800 MPa 11700 MPa
M
Martensite Poisson’s
ratio 0.33 0.33
L Transformation strain 0.054 0.055
∂
∂

T
⎛
⎝
⎜ ⎞
⎠
⎟
L
stress/temperature
ratio during loading
5.32
MPa T–1
5.32
MPa T–1
L
S Start of transformation
loading 390 MPa 344 MPa
L
E End of transformation
loading 401 MPa 363 MPa
T0 Reference temperature 37 °C 37 °C
∂
∂

T
⎛
⎝
⎜ ⎞
⎠
⎟
U
Stress/temperature
ratio during unloading
5.32
MPa T–1
5.32
MPa T–1
U
S Start of transformation
unloading 112 MPa 58 MPa
U
E End of transformation
unloading 93 MPa 42 MPa
CL
S Start of transformation
stress in compression - -
V
L Volumetric
transformation strain 0.054 0.055
Af
Austenite finish
transformation
temperature
22 °C 24 °C


Segment bending
angle 66° 65°
mesh parameters for the esophageal stent during crush-
ing are listed in Table 3. Using the Abaqus contact
module, only the contact between the outer stent surface
and the inner plane surfaces is activated. In this contact
algorithm, the master surface for the rigid planes and the
slave surface for the stent are concurrently applied. A
penalty interaction property is employed to enforce the
impermeable boundaries. Two inflexible similar plates
are applied for the pressure on the stent. The distance
between the two planes is 25 mm (equal to the stent
nominal diameter value in the expanded configuration).
The rigid-flexible contact pairs are formed between the
planes and the surfaces of the stent and no friction is
considered for any contact pair. According to Petrini et
al.,7 the lower plane is fixed in all directions and the
higher plane can just move in the Y direction. Four stent
nodes close to the planes are fixed in the X and Z direc-
tions. As a result, the stent can be compressed just in the
Y direction. Displacement loading is thus applied on the
higher plane: the higher plane compresses the stent and
reduces the distance between the two planes to 7.5 mm
(a 70 % reduction in the diameters of the stents) and 2.5
mm (a 90 % reduction in the diameters of the stents).
The higher plane then moves back to its first location and
recovers the stent. The reaction force on the plane is
collected during the compression and recovers the
process. The boundary conditions for the crushing of the
esophageal-duct stent are shown in Figure 4. To de-
crease the calculation time and use the benefits of the
axis symmetry, only one-quarter of the geometrical
model is analyzed. The temperature is adjusted to 37 °C
(the body temperature) during the simulation.
3 RESULTS AND DISCUSSION
Duerig et al.4 have shown a favorable clinical perfor-
mance of the nitinol stents used for esophageal appli-
cation due to a complete superelastic hysteresis loop, the
lowest chronic outward force (COF), the highest radial
resistive force (RRF) and the long plateau stress. Gideon
et al.20 have shown that a lower stress and a higher strain
on the critical points of a stent allow a suitable clinical
performance. COF and RRF are related to the super-
elastic behavior of a nitinol stent. A typical superelastic
stress–strain curve for a self-expanding stent is shown in
Figure 5.15 The stent is crushed by two rigid planes (path
a–b) and later deployed, reaching stress equilibrium with
the duct at point c. The force against the duct is
controlled with the unloading curve (COF) and the force
resisting deformation is controlled with the loading
curve (RRF). Generally, stent designers aim for as high
an RRF as possible, with as low a COF as possible.4
Furthermore, two particular points should be considered
so that the stents are never in the elastic region. In
addition, the stents are known for having a satisfactory
superelastic behavior. Moreover, the stents should be in
the failure-safe domain from mechanical strength and
strain.21 Beule et al.9 have shown that an increase in the
segment angles, radial contraction of the stent and
decrease in the length of the segments improve the
performance of the stent. Their study indicates, however,
that due to the transformation-temperature (Af) effects, in
the case of a lower angle and longer segments, a better
superelastic performance is gained with the nitinol stent.
F. NEMATZADEH, S. K. SADRNEZHAAD: EFFECTS OF THE AGEING TREATMENT ON THE SUPERELASTIC BEHAVIOR ...
456 Materiali in tehnologije / Materials and technology 47 (2013) 4, 453–459
Figure 5: RRF and COF of a typical superelastic hysteresis loop15
Slika 5: RRF in COF zna~ilne superelasti~ne histerezne zanke15
Figure 3: Comparison of the Auricchio-model calculations with the
experimental data obtained for the superelastic nitinol samples No. 1
(66°) and 2 (65°) from Table 2
Slika 3: Primerjava izra~unov po Auricchiovem modelu z eksperimen-
talnimi podatki za superplasti~ne vzorce nitinola {t. 1 (66°) in {t. 2
(65°) iz tabele 2
Figure 4: a) Schematic representation of the boundary conditions for
the crushing of the esophageal stent, b) exact position of the stent, for
which numerical calculations have been performed
Slika 4: a) Shematski prikaz robnih pogojev za tla~enje opore po`iral-
nika, b) to~en polo`aj opore, kjer so bili izvr{eni numeri~ni izra~uni
This advantage is consistent with the observations of
Beule et al.9 Every nitinol stent with a high Af tempe-
rature (close to the body temperature) will be in the
lower COF range with a better fatigue life. A nitinol
stent of an acceptable performance does not require a
high RRF for a low obstruction of esophageal ducts.
With a high obstruction of esophageal ducts, the Af
temperature of nitinol should be much lower than the
body temperature due to the need for a high RRF.
With respect to the nitinol stents illustrated in Figure
1 and quantified in Table 1, this paper first discusses the
results of the 70 % standard crushing according to
Petrini et al.7 and then the results of the 90 % crushing.
According to Table 4, the maximum stress, strain and
displacement during the 70 % crushing show an elastic
behavior of the esophageal stents. A comparison of
sample No. 1 from Table 2 and sample No. 2 from Table
2 indicates that a decrease in the maximum stress from
269.9 MPa to 229.9 MPa results in an insignificant
reduction in the maximum strain from 0.00862 to
0.00850. The maximum radial displacement increases
from 0.00642 to 0.011 under the same conditions. This
increasing ratio is about 71.3 %. Also, these stents,
owing to the non-superelastic behavior, are not suitable
for esophageal applications. According to the stent
shown in Figure 1 and Table 1 (sample No. 1 from
Table 2), the minimum stress to initiate a martensitic
transformation is 390 MPa. The stress level of this stent
is not sufficient for achieving the superelastic behavior.
Consequently, this stent is not appropriate for an
esophageal application. Therefore, discussing other test
results of the stent has no meaning. As a result, the
desired superelastic behavior is not achieved. Finally,
considering the non-superelastic behavior, the stent
would not be suitable for an esophageal-stent appli-
cation. With the material properties of sample No. 2
from Table 2, the minimum stress to initiate a martensite
transformation is 344 MPa. Discussing the test results
for the stents with unacceptable stress levels is not
necessary, even if they can provide a normal spring
behavior. As a result, the desired superelastic behavior is
not achieved with the stents that have very low stresses.
According to Table 5, a comparison of sample No. 1
from Table 2 and sample No. 2 from Table 2 indicates
that a decrease in the maximum stress from 486.8 MPa
to 373.5 MPa results in a insignificant reduction in the
maximum strain from 0.015 2 to 0.014 6. The maximum
radial displacement increases from 0.002 46 to 0.003 40
under the same conditions. This increasing ratio is about
38.2 %. The maximum stress on the internal curvature of
the stent of sample No. 2 from Table 2 is lower than that
of sample No. 1 from Table 2. The former is, hence,
preferred to the latter. The maximum strain on the
internal curvature of the stent of sample No. 1 from
Table 2 is slightly higher than that of sample No. 2 from
Table 2. Consequently, the former helps the dynamic
motion of the stent to be in harmony with the duct. In
summary, despite the fact that the standard crushing level
is 70 %, the stents showed no superelastic behavior at
this crushing level. An increase in the crushing from
70 % to 90 % results in an appearance of the superelastic
behavior and produces a favorable condition for fabri-
cating esophageal stents. This indicates that the designed
stents are capable of withstanding higher strain forces of
up to 90 % while preserving the superelastic behavior. In
addition, a failure of a stent with this value is safe.21 With
90 % crushing, the stent shown in Figure 1 and quan-
tified in Table 1 (sample No. 2 from Table 2) exhibits a
more acceptable mechanical performance as far as the
material properties are concerned. Considering the pro-
perties of the materials used in Table 2 (samples 1 and 2)
at 22 °C and 24 °C, respectively, there should be a
different range of loading and unloading plateau stress.
A lower Af temperature of nitinol and extensive loading
and unloading stress levels of the mechanical hysteresis
are related to the superelastic behavior. Duerig et al.4
have shown that a difference of 7 °C in the Af tempera-
ture results in about a 50 % variation in the stress level.
Patel et al.14 have shown that a decreasing Af temperature
results in a higher upper plateau stress. Figure 6 shows
an increase in the upper plateau stress due to the
difference between the Af temperatures of 22 °C and
24 °C. For a difference of 2 °C, the expected 15 %
variation has been confirmed with the experimental and
numerical results4. The angular difference between the
stent geometries illustrated in Figure 1 and Table 1 is
F. NEMATZADEH, S. K. SADRNEZHAAD: EFFECTS OF THE AGEING TREATMENT ON THE SUPERELASTIC BEHAVIOR ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 453–459 457
Table 3: Mesh parameters for the esophageal stent during crushing
Tabela 3: Parametri mre`e za oporo po`iralnika med stiskanjem
Material Element type Number ofelements
Number of
nodes
Stent C3D8I 3600 14400
Rigid plane R3D4 20000 20402
Table 4: Results for stress, strain and displacement of the esophageal
stent under 70 % crushing
Tabela 4: Rezultati napetosti, raztezka in radialnega raztezka opore
po`iralnika pri 70-odstotnem stiskanju
Stent model Maximumstress (MPa)
Maximum
strain
Radial
displacement
Sample 1 from
Table 2 269.9 0.00862 0.00642
Sample 2 from
Table 2 229.9 0.00850 0.011
Table 5: Results for stress, strain and displacement of the esophageal
stent under 90 % crushing
Tabela 5: Rezultati napetosti, raztezka in radialnega raztezka opore
po`iralnika pri 90-odstotnem stiskanju
Stent model Maximumstress (MPa)
Maximum
strain
Radial
displacement
Sample 1 from
Table 2 486.8 0.0152 0.00246
Sample 2 from
Table 2 373.5 0.0146 0.00340
only 1 degree. The basis for the aforementioned geo-
metric design results from the age treatment of the
Ti-50.7 % Ni alloy at 500 °C and the time intervals of 30
min and 60 min.15 However, the stent shown in Figure 1
and Table 1 (sample No. 2 from Table 2) reveals a more
acceptable mechanical and practical performance in
comparison with that illustrated in Figure 1 and Table 1
(sample No. 1 from Table 2). Nevertheless, according to
Figure 1 and Table 1 (sample No. 1 from Table 2) and
considering the short segments of the stent and a wider
angle between the stent segments, better geometric
conditions for the crushing test are achieved and a more
reasonable mechanical performance is expected as
described by Beule et al.9 The acceptable mechanical
performance of the stent from Figure 1 and Table 1
(sample No. 2 from Table 2) is mainly rooted in its
material properties obtained by the ageing treatment of
the Ti-50.7 % Ni alloy at 500 °C, the time interval of 60
min and the optimum angle of 65°.15 The results of the
numerical calculations performed for the curve of the
stent in Figure 4b obtained with crushing under a high
pressure are demonstrated in Figure 6. Consequently,
based on the desired mechanical standards of the stents
and according to Figure 6, the stent shown in Figure 1
and quantified in Table 1 (sample No. 2 from Table 2)
with a low COF, high RRF, superelastic behavior related
to the hysteresis loop and no stress concentration at the
internal curvature of the stent is preferred to the other
types.
4 LIMITATIONS
This simulation is complex because of the contacts,
non-linear geometry, the material nonlinearity, large
deformation, additional buckling and bending of the
stent. On the other hand, during crushing, the self-con-
tact phenomenon (the contact between the edges of the
stent) is anticipated and it can cause additional stress at
the contact points. However, according to Wu et al.,22
owing to the superelasticity behavior of the stent, these
types of stress can be neglected. The crushing tests have
only been performed in the longitudinal orientations. It is
obvious that more experiments and simulation results
related to the stenosis degree of ducts are needed to
reach to a comprehensive conclusion.
5 CONCLUSIONS
The effects of design parameters and material pro-
perties obtained from the ageing treatment on the
mechanical performance of the z-shaped nitinol stents
for esophageal ducts are investigated. A nitinol stent
shows a better mechanical and practical performance
owing to only a 1-degree angular difference between the
segments of the stent. The material is the Ti-50.7 % Ni
alloy, age treated at 500 °C for a 30 min to 60 min time
interval. The nitinol stent with the segment angle of 65°,
crushing of 90 % and the Af temperature of 24 °C having
a low COF and high RRF shows a favorable superelastic
behavior and a low stress concentration on the internal
curvature. The superelasticity and hysteresis behavior of
nitinol are proven to be promising properties for the
z-shaped wire stents used for fabricating esophageal
ducts.
Acknowledgments
The authors would like to thank Dr. Amir R. Khoei,
professor at Civil Engineering Department, Sharif
University of Technology and Mr. Ehsan Haghighat,
research assistant at McMaster University for assisting
with the implementation of the simulation program to
model the samples. We would also like to thank Mr.
Seyed M. Seyed Salehi, PhD student at Material Science
and Engineering Department, Sharif University of
Technology for many helpful discussions on the nume-
rical analysis of the stents.
6 REFERENCES
1 K. Hameed, H. Khan, R. Din, J. Khan, A. Rehman, M. Rashid,
Self-Expandable metal stents in palliation of malignant esophageal
obstruction, Gomal J. Med. Sci., 8 (2010), 39–43
2 D. Renteln, B. Walz, B. Riecken, T. Kayser, K. Caca, Endoscopic
management of acute esophageal dissection by using a covered
self-expanding metal stent, Gastro. Endoscopy, 69 (2009), 577–580
3 D. Stoeckel, A. R. Pelton, T. Duerig, Self-expanding Nitinol stents:
material and design considerations, Eur. Radio, 14 (2004), 292–301
4 T. Duerig, D. Tolomeo, M. Wholey, An overview of superelastic
stent Design, Min. Invas. Ther. and Allied Technol., 9 (2000),
235–246
5 B. Patrick, B. Snowhill, L. John, L. Randall, H. Frederick, Charac-
terization of Radial Forces in Z Stents, Inves. Radiology, 36 (2001),
521–530
6 F. D. Whitcher, Simulation of in vivo loading conditions of Nitinol
vascular stent structures, Comput. Struct., 64 (1997), 1005–1011
F. NEMATZADEH, S. K. SADRNEZHAAD: EFFECTS OF THE AGEING TREATMENT ON THE SUPERELASTIC BEHAVIOR ...
458 Materiali in tehnologije / Materials and technology 47 (2013) 4, 453–459
Figure 6: Comparison of the superelastic behavior of the 90 %
crushed esophageal-duct stent with the material properties of sample
No. 1 (66°) and 2 (65°) from Table 2
Slika 6: Primerjava superelasti~nega vedenja 90-odstotno stisnjene
opornice po`iralne cevi z lastnostmi materiala vzorca {t. 1 (66°) in {t.
2 (65°) iz tabele 2
7 L. Petrini, F. Migliavacca, P. Massarotti, S. Schievano, G. Dubini, F.
Auricchio, Computational studies of shape memory alloy behavior in
biomedical applications, J. Biomech. Eng., 127 (2005), 716–725
8 C. Kleinstreuer, Z. Li, C. Basciano, S. Seelecke, M. Farber, Com-
putational mechanics of Nitinol stent grafts, J. Biomech., 41 (2008),
2370–2378
9 M. Beule, S. Cauter, P. Mortier, D. Loo, R. Impec, P. Verdonck, B.
Verhegghe, Virtual optimization of self-expandable braided wire
stents, Med. Eng. Phys., 31 (2009), 448–453
10 G. Silber, M. Alizadeh, A. Aghajani, Finite element analysis for the
design of Self-expandable Nitinol stent in an artery, I. J. Energ.
Tech., 2 (2010), 1–7
11 H. V. D. Merwe, B. D. Reddy, P. Zilla, D. Bezuidenhout, T. A. Franz,
Computational study of knitted Nitinol meshes for their prospective
use as external vein reinforcement, J. Biomech., 41 (2008),
1302–1309
12 Y. Liu, S. P. Galvin, Criteria for pseudoelasticity in near-equiatomic
NiTi shape memory alloys, Acta Materialia, 45 (1997) 11,
4431–4439
13 K. Yeung, K. Cheung, W. Lu, C. Chung, Optimization of thermal
treatment parameters to alter austenitic phase transition temperature
of NiTi alloy for medical implant, Mater. Sci. Eng. A, 383 (2004),
213–218
14 M. Patel, D. Plumley, R. Bouthot, The Effects of Varying Active Af
temperatures on the Fatigue Properties of Nitinol Wire, Proc. of
ASM Conf., Boston, 2005, 1–8
15 X. Liu, Y. Wang, D. Yang, M. Qi, The effect of ageing treatment on
shape-setting and superelasticity of a Nitinol stent, Mater. Charact.,
59 (2008), 402–406
16 F. Auricchio, R. Taylor, Shape-memory alloys: Modeling and nume-
rical simulations of the finite-strain superelastic behavior, Comput.
Methods Appl. Mech. Engrg., 143 (1997), 175–194
17 J. Lubliner, F. Auricchio, Generalized plasticity and shape memory
alloy, I. J. Solid. Struct., 33 (1996), 991–1003
18 N. Rebelo, N. Walker, H. Foadian, Simulation of implantable stents,
Abaqus user’s conf., 143 (2001), 421–434
19 A. G. Prince, G. L. Quarini, J. E. Morgan, J. Finlay, Thermomecha-
nical response of 50.7 % Ni-Ti alloy in the pseudoelastic regime,
Mater. Sci. Tech., 19 (2003), 561–565
20 V. Gideon, P. Kumar, L. Mathew, Finite Element Analysis of the
Mechanical Performance of Aortic Valve Stent Designs, Trends
Biomater. Artif. Organs, 23 (2009), 16–20
21 A. R. Pelton, V. Schroeder, M. Mitchell, X. Gong, M. Barneya, S.
Robertson, Fatigue and durability of Nitinol stents, J. Mech. Behav.
Biomed. Mater., 1 (2008), 153–164
22 W. Wu, M. Qi, X. Liu, D. Yang, W. Wang, Delivery and release of
Nitinol stent in carotid artery and their interactions: a finite element
analysis, J. Biomech., 40 (2007), 3034–3040
F. NEMATZADEH, S. K. SADRNEZHAAD: EFFECTS OF THE AGEING TREATMENT ON THE SUPERELASTIC BEHAVIOR ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 453–459 459

D. ASEFI et al.: INVESTIGATION OF THE EFFECT OF SKIN-PASS ROLLING ON THE FORMABILITY ...
INVESTIGATION OF THE EFFECT OF SKIN-PASS
ROLLING ON THE FORMABILITY OF LOW-CARBON
STEEL SHEETS
PREISKAVA U^INKA DRESIRNEGA VALJANJA NA
PREOBLIKOVALNOST MALOOGLJI^NE JEKLENE PLO^EVINE
Davoud Asefi, Hussein Monajatizadeh, Amir Ansaripour, Ahmadreza Salimi
Department of Materials Engineering, Najafabad Branch, Islamic Azad University, P.O.Box 517, Isfahan, Iran
h_monajati@iaun.ac.ir
Prejem rokopisa – received: 2012-10-26; sprejem za objavo – accepted for publication: 2012-12-11
In the present survey, the effects of skin-pass elongation and cold-rolling reduction on the formability parameters are studied.
For this reason, three sets of samples, including those from industry and laboratory, were evaluated. In the first set, samples with
different hot-rolling conditions were submitted to the same amount of skin reduction and their final mechanical properties were
compared. In the second set, samples with similar conditions were tested with different amounts of laboratory cold reduction
after full annealing and, finally, in the third set, fully annealed samples went through different amounts of industrial elongations
in a skin-pass line. The results showed that skin reduction has a significant effect on the work-hardening coefficient and it
should be kept at the minimum value required to prevent the yield-drop phenomenon. This value was observed to be around 0.5
% for St14 sheets with a thickness of 0.7 mm.
Keywords: ferritic steels, plastic behavior, destructive testing, cold rolling
Preiskovan je bil u~inek raztezka in redukcije prereza pri hladnem preoblikovanju na parameter preoblikovanja. Zato so bile
ocenjene tri skupine industrijskih in laboratorijskih vzorcev. V prvi skupini so bili v razli~nih razmerah vro~e valjani vzorci
dresirno valjani z enako redukcijo, primerjane pa so bile tudi njihove kon~ne mehanske lastnosti. V drugi skupini so bili
preizku{eni laboratorijski vzorci z razli~nimi stopnjami hladne redukcije po predhodnem mehkem `arjenju, v tretji skupini pa
mehko `arjeni vzorci izpostavljeni razli~nim stopnjam raztezka na industrijski liniji za dresiranje. Rezultati so pokazali, da ima
redukcija pri dresiranju pomemben vpliv na koeficient hladnega utrjevanja, ki naj bi imel najmanj{o vrednost, da bi se prepre~il
pojav zmanj{anja plasti~nosti. Ugotovljeno je bilo, da je pri plo~evini iz jekla St14 z debelino 0,7 mm ta vrednost okrog 0,5 %.
Klju~ne besede: feritna jekla, plasti~nost, poru{ni preizkusi, hladno valjanje
1 INTRODUCTION
Complex stamping is an important stage in forming
automotive body parts and various other products. The
steel industry has made many attempts to enhance the
quality of their products so that they would lead to easier
and faster metal-forming processes. The formability is
defined as the resistance of a material to necking and
thinning in the thickness direction. The former is known
to be related to the work-hardening exponent (n) and the
latter to the average plastic-strain ratio (r).1,2 Drawability,
on the other hand, is the ability of a material to easily
flow in the plane of the sheet without thinning in the
thickness direction, and stretchability represents the
ability of the material to resist localized necking.3,4 In a
combined forming operation, r, the normal anisotropy
and n, the work-hardening coefficient can be represen-
tatives of the drawability and stretchability, respectively.
They are obtained with a uniaxial tensile test. Industrial
steel sheets with typical values of (r) in the range
1.6–2.0 and n in the range 0.22–0.24 are known as
deep-drawing-quality (DDQ) products.5
In recent years, the effects of process conditions and
metallurgical microstructure on the quality of low-car-
bon steel sheets have been studied by mathematical
models that use experimental observations as their data-
base. The formability is affected by all the process
parameters, such as hot rolling, annealing and the
skin-pass conditions.6,7 Numerous studies have been
published regarding the influence of the processing
parameters on the microstructure and formability of
low-carbon steel.8–10 However, little has been published
on the effect of the skin-pass parameters on the formabi-
lity.11
Skin pass or temper rolling is the last stage in thin-
strips production, introducing a small plastic strain in the
sheet to bypass the non-uniform deformation region and
erase the yield-point jogs. It also improves the flatness of
the strip and makes a certain surface roughness and
smoothness.12 The reduction percentage of temper rol-
ling has a significant effect on its mechanical properties.
Therefore, the preset of the reduction percentage is
restricted not only by the flatness and the surface rough-
ness of the strip steel, but also by the mechanical pro-
perties. The mechanical properties of low-carbon steels
are influenced by a small strain processing treatment.13,14
But the effects of a little cold strain and the reduction
percentage of skin pass on the formability parameters of
steel sheets still need to be investigated. This important
issue is the purpose of the present study.
Materiali in tehnologije / Materials and technology 47 (2013) 4, 461–466 461
UDK 621.771:669.14 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)461(2013)
2 EXPERIMENTAL PROCEDURE
2.1 Materials
Three kinds of samples were used in this study. The
chemical composition of each is shown in Table 1. The
first set, prefixed by A from 1 to 5, had a different
Al-to-N ratio and coiling temperature (CT), as shown in
Table 1. All of them received the same amount of 0.7 %
industrial skin reduction and their properties were mea-
sured before and after the skin-pass process.
The effect of the skin-pass strain on the final
formability of the sheets was studied in set B. For this
reason, different amounts of cold reduction were applied
using a laboratory rolling machine on 40 cm × 60 cm
samples obtained from a longitudinal section of a coil
after an industrial box anneal and near each other. Again,
their properties were measured before and after the skin
pass (Table 2).
It should be noted that there are two operations in
skin pass rolling: a stretching that is applied using a
tension leveler and a compression applied from rolls. As
it was not possible to apply the stretching load on labora-
tory skin rolling, a separate set of samples, set C, was
imposed on a different amount of reductions using an
industrial skin machine to investigate the effect of the
real skin process on the final properties.
2.2 Experiments
Standard uniaxial tensile tests were carried out on the
samples in three directions, i.e., 0°, 45° and 90°, corres-
ponding to the rolling direction, to obtain the strength
and ductility of the specimens as well as the r and n
coefficients. The normal anisotropy was calculated
according to the equation:
r = (r0 + 2r45 + r90)/4 (1)
n was also calculated in a similar way to reach an ave-
rage value.2
The average grain sizes (d) of the samples were
measured according to ASTM E112-96 (Jeffries method)
and the optical micrograph of the samples was prepared
by a standard metallographic technique. As the grain
structure in low-carbon steel sheets after cold rolling and
box annealing is generally pancaked, the aspect ratio of
the elongated grains was also evaluated using measure-
ments of the large and small diameters of the grains,
separately.
3 RESULTS AND DISCUSSION
3.1 Effects of the process parameters on the final
properties under similar skin-pass conditions
The magnitudes of r and n for the samples of set A,
before and after the skin pass, are illustrated in Figure 1.
Although r does not demonstrate any large change
before and after the skin pass, the variations of n are
drastic, clearly showing a significant decrease after the
treatment.
In accordance with Figure 1a, the skin pass does not
have any significant effect on r. This parameter is mostly
affected by grain orientation or texture.2 It has been
shown that high r values are displayed by materials that
have a high proportion of grains oriented with {111}
planes parallel to the sheet plane. The {111} fiber texture
is particularly beneficial for imparting good deep draw-
ability.15 It has also been shown that the morphology of
cementite particles and the precipitation of fine AlN pre-
cipitates control the texture of low-carbon steel
D. ASEFI et al.: INVESTIGATION OF THE EFFECT OF SKIN-PASS ROLLING ON THE FORMABILITY ...
462 Materiali in tehnologije / Materials and technology 47 (2013) 4, 461–466
Table 1: Chemical composition of the investigated samples in mass fractions, w/%
Tabela 1: Kemijska sestava preiskovanih vzorcev v masnih dele`ih, w/%
NUM w(C) w(Si) w(Mn) w(Al) w(N)/(μg/g) w(Al)/w(N) CT/°C
A1 0.045 0.009 0.209 0.049 19 25.7 560
A2 0.049 0.01 0.204 0.047 30 15.6 560
A3 0.039 0.009 0.209 0.049 19 25.7 620
A4 0.05 0.01 0.0220 0.055 58 9.4 625
A5 0.05 0.01 0.220 0.055 58 9.4 625
B 0.036 0.012 0.27 0.043 25 17.2 –
C 0.034 0.009 0.215 0.045 35 12.8 –
CT – Coiling Temperature
Table 2: Normal anisotropy, work hardening values, applied reduction and grain size and aspect ratio for the samples in set B
Tabela 2: Normalna anizotropija, vrednosti utrjevanja pri preoblikovanju, uporabljena redukcija, velikost zrn in razmerje {irina proti debelini
vzorcev iz skupine B
NUM Cold reduction (%) r n Grain sized/(μm) Aspect ratio
B0 0 1.60 0.24 26 2.26
B1 2.5 1.60 0.19 24 2.9
B2 3.8 1.62 0.19 26 3.1
B3 4.8 1.59 0.15 27 3.6
sheets.15,16 The most important process parameters
affecting these microstructural features are hot and cold.
The rolling parameters and subsequent annealing
parameters, such as the heating rate and soaking time are
important.6,10 The most important microstructural effect
exerted by the limited strain in skin-pass rolling, except
the very small strain near the surface, is releasing dislo-
cations from short-range local locks of carbon and nitro-
gen atoms and eliminating the yield drop and hetero-
geneous deformation.8,9 Therefore, cold deformation due
to the skin-pass process has a slight effect on the volume
fraction of {111} fiber and its influence on r is almost
negligible.1
Figure 1b implies that for all samples of set A, n is
decreased considerably after the skin-pass process.
The variation percentage of n in each sample is
shown in Figure 2. It can also be seen that the drop value
is not the same for all the samples, varying in the range
of 8.7 % to 30 % for different samples. This reduction is
due to the direct relationship between the work-harden-
ing coefficient and the microstructure variations. The
relationship between the microstructure and the work-
hardening coefficient can be explained by Eq 1, which
was first expressed by Antoine:8,9
n = − + + + +0 450 0 001. . ( )    P SA GB PCT F0.2 (2)
The stresses in this equation are the strengthening
contribution of Peierls (P), the solid solution (SA) (both
affected by the chemical composition), the precipitates
(PCT), the grain boundaries (GB) and the dislocation
density (F0.2). The above equation shows that all the
mentioned microstructural factors affect n directly.
Among them, the dislocation density and the grain-
boundary characters (in the form of the shape and the
aspect ratio) may be changed during the skin-pass
rolling, while the chemical composition’s and the preci-
pitates’ contributions remain unchanged.
The study of the grain characteristics in these sam-
ples implied a very negligible difference between the
grain elongations. Therefore, the only influential micro-
structure factor would be the variation of the dislocation
density in the various samples.
Among the microstructural parameters that may
affect the dislocation density during cold work, the
precipitates are of paramount importance. They can play
the role of an obstacle preventing dislocation movement
and increase the work-hardening rate. Hence, the work-
hardening coefficient decreases. This can be explained
better by referring to the results of Antoine’s work.
Figure 3 illustrates the dislocation density calculated for
the samples with a different volume percentage of TiC
precipitates after 0.02 % strain. These data are derived
from the results of Antoine’s work. It can be seen from
the figure that by increasing the volume fraction of the
precipitates, the dislocation density is increased and n is
decreased.8,9
The difference in the amount of work hardening
coefficient for the samples in set A, therefore, may be
attributed to the volume fraction of the precipitates
before the skin pass, which in the case of St14 steels, and
in absence of carbide-forming elements like Ti, are
mostly AlN particles. The volume fraction and the size
of the AlN precipitates are influenced by various pro-
D. ASEFI et al.: INVESTIGATION OF THE EFFECT OF SKIN-PASS ROLLING ON THE FORMABILITY ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 461–466 463
Figure 1: Variation of: a) r and b) n before and after the skin-pass
rolling for samples in set A
Slika 1: Spreminjanje: a) r in b) n pred dresirnim valjanjem vzorcev
iz skupine A in po njem
Figure 2: Variation percentage of n after skin pass in the samples of
set A
Slika 2: Spreminjanje dele`a n pri dresirnem valjanju vzorcev iz sku-
pine A
duction parameters, with the most important ones being
the cooling rate and the coiling temperature after hot
rolling,10 and the heating rate during annealing.1 The
reported CTs for the samples in set A are shown in Table
1. It is clear that A3, A4 and A5 experienced the highest
CT.
Coiling above 600 °C can promote the precipitation
of AlN precipitates and affect the microstructural
features during the following annealing process.10 As all
the samples in set A have experienced the same anneal-
ing conditions, it can be concluded that the differences in
the CT have led to a different amount of work-hardening
coefficient due to the different volume fraction of AlN
under the same conditions of the skin-pass process.
3.2 Effect of experimental cold reduction on the for-
mability parameters of fully annealed samples
Figure 4 shows the dependences of r and n on the
percentge of cold reduction (after full annealing) for the
samples in set B. It is implied that while increasing the
cold reduction the percentage does not have any obvious
effect on r in the examined range,1 it decreases n in a
discernible manner. For instance, in the case of the sam-
ple without cold work, the amount of n is almost 0.25,
whereas it reaches 0.15 after 4.8 % cold reduction. As
mentioned in the previous section, the variations of n
with the percentage of cold work may be related to
variations in the dislocation density. On the other hand,
the slip distance of dislocations, which is in relation to
the grain size, is expected to affect n due to the increas-
ingly homogeneous strain region.8,9 To examine this
parameter, the variations of the grain size with applied
elongation were characterized and illustrated in Figure 5.
As can be seen, the variation of grain size with in-
creasing cold-work percentage in the examined range is
not considerable. However, the grains’ aspect ratio is
increased in proportion to cold reduction percentage, due
to the elongation along the rolling direction. This con-
firms that the most influential factor in decreasing n with
cold reduction is increasing the dislocation density. It
suggests that the limitation of cold reduction in the skin-
pass rolling can prevent an excess drop of n . However,
the amount of the reduction should be in such a range
that would guarantee the elimination of the heteroge-
neous deformation region (yield drop phenomenon).
D. ASEFI et al.: INVESTIGATION OF THE EFFECT OF SKIN-PASS ROLLING ON THE FORMABILITY ...
464 Materiali in tehnologije / Materials and technology 47 (2013) 4, 461–466
Figure 4: Variation of r and n with cold reduction in samples of set B
Slika 4: Spreminjanje r in n z redukcijo v hladnem vzorcev iz skupine B
Figure 3: The dependence of: a) dislocation density on volume
fraction of TiC precipitates and b) work-hardening coefficient on the
dislocation density, calculated and drawn from Antoine’s results9,10
Slika 3: Odvisnost: a) gostote dislokacij od volumenskega dele`a
izlo~kov TiC in b) koeficient utrjevanja pri hladni predelavi na gostoto
dislokacij, izra~unano in vzeto iz rezultatov Antoina9,10
Figure 5: The relation of grain sizes and their aspect ratio to cold
reduction in the samples of set B
Slika 5: Odvisnost velikosti zrn in njihovega razmerja {irina-debelina,
pri hladni redukciji vzorcev iz skupine B
3.3 Effect of skin-pass elongation on the formability
properties of industrial samples
The effects of the skin-pass elongation on the r and n
values in the samples of set C are shown in Figure 6.
These samples had the same chemical composition
(Table 1) and the same processing conditions, except the
amount of elongation in the skin-pass step.
As can be seen in Figure 6, and as can be expected
from the results of the previous section, the elongation
percentage of the skin pass does not have any influence
on n , whereas it severely decreases n . It should be noted
that although the strain path in experimental skin rolling
(compression from rolls and the tension applied on the
whole line between the feeding and coiling rolls) is
different from the laboratory rolling experiments (just
compression from the rolls) in the previous section, it
does not have any influence on the variation of the
work-hardening coefficient. This is due to the fact that n
is mostly affected by the dislocation density, which is
similar in the experimental range of strain in both
laboratory and industrial experiments.
Figure 6 shows that the elongation percentage of
0.5 % in the skin pass has led to the highest amount of
work-hardening coefficient. Laboratory tests showed that
lower amounts of elongation percentage do not prevent
the occurrence of yield-drop phenomenon. It may be
concluded that the most suitable elongation percentage
in the skin-pass step for the sheet thickness studied (0.7
mm) is 0.5 %.
4 CONCLUSIONS
In this study, low-carbon steel sheets were investi-
gated from a formability point of view.
The following results were obtained. The work-hard-
ening coefficient is mostly affected by the dislocation
density. Any microstructural feature that can increase the
dislocation density can decrease n . While the skin-pass
reduction after full annealing does not have any large
effect on r, it may change n drastically. Therefore, the
limitation of skin reduction to a minimum value is
required to prevent yield drop, which is an important
process parameter that can preserve n at an appropriate
level after it reaches its maximum value during the
annealing process. This value was observed in this study
to be around 0.5 % for St14 sheets with a thickness of
0.7 mm.
Acknowledgements
This research was conducted under a contract with
the Mobarakeh steel company, Isfahan, Iran. The support
from the research & development management, MPT,
and skin-pass and product laboratory personnel are
gratefully appreciated.
5 REFERENCES
1 Y. Jia, H. Guo, R. Li, H. Li, Effects of Rolling Processes on Yield
Ratio and Formability of Hot Rolled Gas Cylinder Steel, Journal of
Iron and Steel Research, International, 19 (2012) 3, 52–55
2 M. Safavi, S. M. Abbasi, R. Mahdavi, Influence of Aluminum Con-
tent on Mechanical Properties and Cold Workability of Fe-33Ni-
15Co Alloy, Journal of Iron and Steel Research, International, 19
(2012) 2, 67–72
3 M. A. Akoy, E. S. Kayali, H. Cimenoglu, The influence of micro-
structural features and mechanical properties on the cold formability
of ferritic steel sheets, ISIJ Int, 44 (2004), 422–428
4 H. Monajati, D. Asefi, A. Parsapour, S. Abbasi, Analysis of the
effects of processing parameters on mechanical properties and
formability of cold rolled low carbon steel sheets using neural
networks, Comput Mater Sci, 49 (2010), 876–81
5 C. Capdevila, C. Garcia-Mateo, F. G. Caballero, C. Garcia de
Andres, Neural network analysis of the influence of processing on
strength and ductility of automotive low carbon sheet steels, Comput
Mater Sci, 38 (2006), 192–201
6 P. Antoine, S. Vandeputte, J. B. Vogt, Effect of microstructure on
strain-hardening behavior of a Ti-IF steel grade, ISIJ Int, 45 (2005),
399–404
7 P. Antoine, S. Vandeputte, J. B. Empirical model predicting the
value of the strain-hardening exponent of a Ti-IF steel grade, Mater
Sci Eng A, 433 (2006), 55–63
8 C. Capdevila, J. P. Ferrer, F. G. Caballero, C. Garcia De Andres,
Influence of processing parameters on the recrystallized micro-
structure of extra-low-carbon steels, Metal Mater Trans A, 37 (2006),
2059–2068
9 H. Monajati, D. Asefi, Investigation of the effect of Work Hardening
coefficient on Formability of low carbon Steel sheets and the
improvement methods for its enhancement, Technical Report,
Mobarakeh Steel Co, 2008
10 H. Kijima, N. Bay, Influence of tool roughness and lubrication on
contact conditions in skin-pass rolling, J Mater Proc Tech, 125
(2009) 48, 38–41
11 Q. L. Ma, W. D. Cheng, L. H. Min, Effect of temper rolling on
tensile properties of low-Si Al-killed sheet steel, Journal of Iron and
Steel Research, International, 16 (2009) 3, 64–67
12 E. J. Hoggan, R. I. Scott, M. R. Barnett, P. D. Hodgson, Mechanical
properties of tension leveled and skin passed steels, J Mater Proc
Tech, 125 (2002), 55–63
13 R. K. Ray, J. J. Jonas, R. E. Hook, Cold rolling and annealing
textures in low carbon and extra low carbon steels, Int Mater Rev, 39
(1994), 129–71
D. ASEFI et al.: INVESTIGATION OF THE EFFECT OF SKIN-PASS ROLLING ON THE FORMABILITY ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 461–466 465
Figure 6: Variation of r and n with elongation percentage in skin-pass
rolling for samples in set C
Slika 6: Spreminjanje r in n z dele`em raztezka pri dresirnem valjanju
vzorcev iz skupine C
14 T. De Cock, C. Capdevila, F. C. Caballero, Global recrystallization
model of low carbon sheet steels with different cementite contents,
Mater Sci Eng A, 519 (2009), 9–18
15 J. Pero-Sanz, M. Ruiz-Delgado, V. Martinez, J. I. Verdeja, Annealing
textures for drawability: influence of the degree of cold rolling
reduction for low-carbon and extra low-carbon ferritic steels, Mater
Charact, 43 (1999), 303–309
16 J. P. Ferrer, T. D. Cock, C. Capdevila, F. G Caballero, C. Garcia de
Andres, Comparison of the annealing behavior between cold and
warm rolled ELC steels by thermoelectric power measurements,
Acta Mater, 55 (2007), 2075–83
D. ASEFI et al.: INVESTIGATION OF THE EFFECT OF SKIN-PASS ROLLING ON THE FORMABILITY ...
466 Materiali in tehnologije / Materials and technology 47 (2013) 4, 461–466
A. SACHDEVA et al.: SYNTHESIS, CHARACTERIZATION AND SENSING APPLICATION OF A SOLID ALUM/FLY ASH ...
SYNTHESIS, CHARACTERIZATION AND SENSING
APPLICATION OF A SOLID ALUM/FLY ASH
COMPOSITE ELECTROLYTE
SINTEZA, KARAKTERIZACIJA IN MO@NOST UPORABE
TRDNEGA KOMPOZITNEGA ELEKTROLITA GALUN-LETE^I
PEPEL
Amit Sachdeva1, Roja Singh2, Pramod K. Singh2, Bhaskar Bhattacharya2
1Department of Nanotechnology, School of Bioscience and Biotechnology, Lovely Professional University, Phagwara-144411, Punjab, India
2Material Research Laboratory, School of Engineering and Technology, Sharda University, Greater Noida-201310, India
b.bhattacharya@sharda.ac.in
Prejem rokopisa – received: 2012-11-16; sprejem za objavo – accepted for publication: 2013-01-03
A new composite system using low-cost materials, potash alum and fly ash, was prepared and characterized using various
techniques. Complex impedance spectroscopy was used for the conductivity measurement. A conductivity enhancement due to
an addition of fly ash was observed and the maximum value was obtained with the 65 : 35 compositions. The maxima in
conductivity can be explained with the percolation theory. Infrared spectroscopy (IR) as well as X-ray diffraction (XRD)
confirmed the composite nature of the samples. Scanning electron microscopy (SEM) shows a good homogeneous mixture of
the composite material which is essential for the conductivity enhancement. Based on the maximum-conductivity composition,
a humidity sensor was fabricated which showed a good sensing behavior.
Keywords: composite, conductivity, IR, XRD, humidity sensor
Pripravljen in karakteriziran z razli~nimi tehnikami je bil nov kompozitni sistem iz poceni kalijevega galuna in lete~ega pepela.
Za meritve prevodnosti je bila uporabljena kompleksna impedan~na spektroskopija. Opa`eno je bilo maksimalno pove~anje
prevodnosti zaradi dodatka lete~ega pepela pri razmerju sestave 65 : 35. Maksimum prevodnosti lahko razlo`imo s teorijo
pronicanja. Infrarde~a spektroskopija (IR) in rentgenska difrakcija (XRD) sta potrdili kompozitno osnovo vzorcev. Vrsti~na
elektronska mikroskopija (SEM) je pokazala homogeno me{anico kompozitnih materialov, ki je klju~na za pove~anje
prevodnosti. Na osnovi sestave z najve~jo prevodnostjo je bil izdelan senzor vlage, ki je pokazal dobro ob~utljivost.
Klju~ne besede: kompozit, prevodnost, IR, XRD, senzor vlage
1 INTRODUCTION
Composite materials are well known materials
playing a dominant role in the industrial areas such as
sport, aerospace, automotive industry and transportation.
Due to the pioneering work of Liang and the subsequent
efforts of many researchers, solid-composite electrolytes
have been widely studied for their attractive behavior.1 A
significant enhancement in the conductivity has been
observed in the multiphase materials with very small
amounts of dispersoids. Fly ash, one of the dispersoids,
has led to a multifold increase in the ionic conductivity
when together with alumina. Several theoretical appro-
aches have also been made to explain such a behavior. A
comparison of a pure system and a two-phase mixture
reveals that the two-phase mixtures in general exhibit a
higher conductivity than those of the starting mate-
rials.2–5 In this paper we report on an increase in the
conductivity for a new composite using two low-cost
materials, i.e., potash alum and fly ash.
Fly ash is a waste product produced from coal-fired
thermal-power stations during the combustion of coal. It
is an alkaline grey powder which causes serious envi-
ronmental problems. Due to the environmental regu-
lations, new ways of utilizing fly ash have to be explored
in order to safeguard the environment and provide useful
ways for its disposal. Hence, there is a considerable
interest in utilizing fly ash as a raw material. More
recently, materials scientists and engineers have
suggested and devised various methods of using this
waste product for the synthesis of some useful
composites.6–8 Raw fly ash consists of quartz and mullite
as the crystalline phases and an amount of a glassy
phase.9 Efforts have been made to understand the
electrical conductivity and dielectric behavior of fly ash
and it was observed that this material possesses a very
high relative dielectric constant of the order of 104. Such
a high dielectric constant is one of the important
parameters in a capacitor fabrication and microwave
absorption applications.8–14 The common alum is a
double sulphate of potassium and aluminum,
K2Al2(SO4)4 · 24H2O, which is a white crystalline
powder readily soluble in water.
Through this work we have made a successful effort
in preparing potash alum/fly ash composites and develop
a humidity sensor which gives newer ways of a better
utility of fly ash.
Materiali in tehnologije / Materials and technology 47 (2013) 4, 467–471 467
UDK 66.017:669.018.95:544.6.018.4 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)467(2013)
2 EXPERIMENTS
2.1 Preparation of the sample
Potash alum was purchased and used without any
further purification while fly ash of an unknown purity
and composition was collected from a local supplier.
Appropriate amounts of potash alum and fly ash were
weighed and thoroughly mixed in an agate mortar and
pestle (2 h) and this was followed by a pulverization
and pelletization in a nickel-plated steel die at the
pressure of 2.5 t using a hydraulic pelletizer machine.
The circular-shaped pellets thus obtained were 0.15 cm2
in area and 4–6 mm in thickness. Infrared spectroscopy
(Perkin Elmer 883) was carried out to study the
composite nature and the functional groups present in the
composite electrolyte. For further confirmation, X-ray
diffraction was obtained using a Rigaku D/max-2500 in
the range of 2 = 20–55° with a scan rate of 1° min–1.
Scanning electron microscopy (SEM) was used to check
the surface morphology. SEM micrographs were re-
corded using SEM (Hitachi S-570). For all the electrical
measurements, a conductive silver paste was coated onto
both surfaces of the pellets and dried under room envi-
ronment. The complex impedance spectroscopy was
used to calculate the bulk electrical conductivity of the
polymer electrolyte films. In our laboratory we used
pressure-contacted stainless-steel electrodes and connec-
ted them with a CH instruments electrochemical work-
station (Model CHI604D) with a frequency range of
0.00001 Hz to 100 kHz. The electrical conductivity ()
was evaluated using the following formula:
 = Rb (l/A)
where  is the ionic conductivity, Rb is the bulk resi-
stance, l is the thickness of the pellet and A is the area
of a given sample.
3 RESULTS AND DISCUSSION
3.1 Complex impedance spectroscopy
The ionic conductivity of pure potash alum and the
alum doped with fly ash pellets were measured using the
complex impedance spectroscopy. The bulk resistance
(Rb) was determined from an intersection at a high
frequency with the real axis in the impedance plots. The
complex impedance plot of a typical potash alum/fly ash
composition is shown in Figure 1. The ionic conducti-
vity values were calculated from the impedance spectro-
scopic data and are listed in Table 1 and plotted in
Figure 2.
It is observed that initially the electrical conductivity
of the composite increases with the increasing content of
fly ash. It attains its maximum at 65 % fly ash and then
decreases gradually. To interpret such variations in
conductivity, several theoretical models have been pro-
posed in the literature. The percolation model considers
the highly conducting paths that were created along the
interface between the host electrolyte and the dispersoid.
In the present system the conductivity maxima could be
explained with the percolation model.9 The maximum is
obtained for the composition, with which the percolation
threshold is achieved. The water of crystallization
A. SACHDEVA et al.: SYNTHESIS, CHARACTERIZATION AND SENSING APPLICATION OF A SOLID ALUM/FLY ASH ...
468 Materiali in tehnologije / Materials and technology 47 (2013) 4, 467–471
Figure 2: Variation in conductivity with the concentration of fly ash in
a potash alum/fly ash composite system
Slika 2: Spreminjanje prevodnosti s koncentracijo lete~ega pepela in
kalijevega galuna v kompozitnem sistemu
Figure 1: Complex impedance plot of a typical potash alum/fly ash
solid electrolyte system
Slika 1: Kompleksno impedan~no podro~je zna~ilnega trdnega elek-
trolita kalijev galun-lete~i pepel
Table 1: Room-temperature electrical conductivity of a potash alum/
fly ash composite system
Tabela 1: Elektri~na prevodnost kompozitnega sistema kalijev galun-
lete~i pepel pri sobni temperaturi
Composition, mass fraction,
w/%
Potash alum/fly ash
Electrical conductivity
/(S/cm)
95 : 5 3.2 × 10–6
85 : 15 5.3 × 10–6
75 : 25 9.2 × 10–6
65 : 35 1.5 × 10–5
60 : 40 1.2 × 10–5
40 : 60 5.4 × 10–6
10 : 90 4.6 × 10–6
present in the alum gets adsorbed on the surface of the
composite, in which the movement of H+ and OH– ions is
responsible for an increase in the conductivity. The
decrease in the conductivity after the maximum is simi-
lar to those obtained in similar composite systems.15–17
3.2 Temperature dependence of conductivity
Figure 3 shows the variation in conductivity with the
temperature of the maximum conductivity composition
of the composite. The variation in conductivity indicates
that the physisorbed water present in potash alum plays
an important role in the conductivity enhancement. The
conductivity rises with an increase in the temperature up
to 50 °C and then begins to fall; after this a constant
value is attained. This is because the physisorbed water
present in potash alum is lost between 45 °C to 50 °C
leading to a loss of the H+ and OH- ions that were
responsible for the conductivity of the composite sample.
3.3 Scanning electron microscopy
Scanning electron microscopy (SEM) was used to
check the surface morphology of composite electrolytes.
We have recorded the SEM micrographs using a SEM
instrument (SEM, Hitachi S-570) and the micrographs
are shown in Figure 4.
It is clear from this figure that potash alum shows a
needle-like morphology (Figure 4a) while fly ash (of an
unknown purity) shows a more or less irregular structure
when seen under SEM (Figure 4c). From the potash
alum/fly ash composite SEM micrograph it is clear that
the composite system shows a mixed morphology where
the white-colored fly ash is uniformly mixed with the
light-blackish potash-alum grains (Figure 4b).
3.4 Infrared spectroscopy
Infrared spectroscopy (Perkin Elmer 883) was carried
out to study the composite nature and the functional
groups present in the composite electrolyte. Figure 5
shows the recorded IR spectra of pure potash alum, pure
fly ash and the potash alum doped with fly ash (the
maximum  composition).
It is obvious that the IR spectrum of the composite
material (Figure 5c) contains the peaks related to either
pure potash alum (Figure 5a) or to alum (Figure 5b).
An absence of any new peaks in the composite
electrolyte other than those of the host materials clearly
confirms the composite nature. The IR spectrum of
potash alum (Figure 4a) based on one monovalent and
one trivalent cation has been published.18 Ross gave an
interpretation of the infrared spectrum of potassium alum
as (981, 1200, 1105, 618 and 600) cm–1 for (SO4)2–. The
water stretching modes were reported at 3400 cm–1 and
3000 cm–1, the bending modes at 1645 cm–1 and the
vibrational modes at 930 cm–1 and 700 cm–1.19
A. SACHDEVA et al.: SYNTHESIS, CHARACTERIZATION AND SENSING APPLICATION OF A SOLID ALUM/FLY ASH ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 467–471 469
Figure 5: IR spectra of: a) pure potash alum, b) pure fly ash, c) potash
alum/fly ash composite
Slika 5: IR-spektri: a) ~isti kalijev galun, b) ~isti lete~i pepel, c) kom-
pozit kalijev galun-lete~i pepel
Figure 3: Variation in conductivity with temperature (maximum con-
ductivity)
Slika 3: Spreminjanje prevodnosti s temperaturo (maksimalna prevod-
nost)
Figure 4: SEM micrographs of: a) pure potash alum, b) potash
alum/fly ash composite and c) pure fly ash
Slika 4: SEM-posnetki: a) ~isti kalijev galun, b) kompozit kalijev
galun-lete~i pepel in c) ~isti lete~i pepel
Figure 5b shows the IR spectrum of fly ash, in which
the characteristic IR peak observed at 1094 cm–1 may be
attributed to the presence of silica.17 This peak confirms
the highest % of silica in fly ash, which is confirmed by
a chemical analysis.8 Several absorption peaks are
observed in the range from 1400 cm–1 to 1300 cm–1,
which correspond to the various metal oxides present in
fly ash. A characteristic broad band at 3422 cm–1 in the
spectrum of the composite is due to the N-H stretching
of aniline (Figure 4b). The absorptions at (1297.49 ± 10,
1139.71 ± 10, 1089.48 ± 10 and 794.81 ± 20) cm–1
correspond to the fly-ash constituents in the potash
alum/fly ash composite (Figure 5c).
3.5 X-ray diffraction
To further investigate the composite nature, we
recorded the XRD patterns of the composite electrolytes
along with the host materials and the recorded patterns
are shown in Figure 6. It was noted that the peaks that
appear in the XRD pattern of the composite material
(Figure 6c) also appear in the XRD patterns of fly ash
(Figure 6a) or potash alum (Figure 6b) affirming the
composite nature that we have already found with the IR
data.
4 APPLICATIONS
4.1 Humidity sensor
On the basis of the maximum electrical conductivity
sample we tried to fabricate a humidity sensor in our
laboratory. The optical photograph of the sensor is
shown in Figure 7. To develop it we made a finger-type
electrode on the surface of the composite pellet using
vacuum-coated silver paint as the electrode material and
deposited the pattern using the vacuum coating unit
(Hind High Vacuum, India) at a pressure of 1.33 × 10–3
mbar. A conducting copper wire was used for the contact
and sensing behavior. To measure the response of the
humidity sensor we also developed a constant-humidity
chamber in our laboratory as designed by Chandra et al20.
A very simple approach was adopted for creating diffe-
rent constant humidities for an in-situ measurement
inside the chamber. It is known that the water-vapor
pressure in over-saturated solutions of šdifferent salts’
give different relative humidities. We applied different
humidities inside the close chamber developed by us and
measured the sensor response. The response of the
sensor (voltage vs. humidities) is shown in Figure 8. It is
clear from this figure that the developed sensor shows a
good sensing behavior with a quick response. Within a
short period the developed sensor shows an exponential
decrease in the voltage with an increase in the humidity
level.
5 CONCLUSION
A solid-state composite based on potash alum/fly ash
has been developed and well characterized using various
techniques. The electrical-conductivity measurement
shows that by adding fly ash the electrical conductivity
enhances the attained maxima at 65 % of fly ash in the
composition to the conductivity value of 1.5 × 10–5 S/cm
A. SACHDEVA et al.: SYNTHESIS, CHARACTERIZATION AND SENSING APPLICATION OF A SOLID ALUM/FLY ASH ...
470 Materiali in tehnologije / Materials and technology 47 (2013) 4, 467–471
Figure 8: Response of the humidity sensor based on the potash alum/
fly ash composite (the maximum conductivity sample)
Slika 8: Odziv na vlago senzorja na osnovi kompozita kalijev galun-
lete~i pepel
Figure 6: XRD patterns of: a) pure fly ash, b) potash alum and c)
potash alum/fly ash composite
Slika 6: XRD-posnetki: a) ~isti lete~i pepel, b) kalijev galun in c)
kompozit kalijev galun-lete~i pepel
Figure 7: Photograph of developed humidity sensor based on the
potash alum/fly ash composite
Slika 7: Posnetek razvitega senzorja vlage, ki temelji na kompozitu
kalijev galun-lete~i pepel
and this is followed by a decrease. IR as well as XRD
confirmed the composite nature, while SEM affirmed the
homogeneous mixing of potash alum/fly ash. A humidity
sensor has been fabricated (with the maximum conduc-
tivity sample) and it shows a stable and good perfor-
mance.
Acknowledgments
This work was supported by the DST project
(SR/S2/CMP-0065/2010) of the government of India.
6 REFERENCES
1 C. C. Liang, J. Electrochem. Soc., 120 (1973), 1289
2 T. Jow, J. B. Wagner, J. Electrochem. Soc., 126 (1979), 1963
3 R. C. Agrawal, R. K. Gupta, J. Mat. Sci., 34 (1999), 1131
4 N. Vaidehi, R. Akila, A. K. Shukla, K. T. Jacob, Mater. Res. Bull., 21
(1986), 909
5 N. J. Dudney, Annu. Rev. Mater. Sci., 19 (1989), 103
6 B. B. Owens, P. M. Skarstad, Solid State Ionics, 53–56 (1992), 665
7 K. K. Nanda, S. N. Sahu, Adv. Mater., 13 (2001), 280
8 D. N. Bose, D. Majumdar, Bull. Mater. Sci., 6 (1984), 223
9 A. Bunde, W. Dietrich, E. Roman, Phys. Rev. Lett., 55 (1985), 5
10 K. Ahmad, W. Pan, S. L. Shi, Appl. Phys. Lett., 89 (2006), 133122
11 D. Rout, V. Subramanian, K. Hariharan, V. Sivasubramanian, V. R.
K. Murthy, J. Phys. Chem. Solids, 67 (2006), 1629
12 J. Maier, J. Phys. Chem. Solids, 46 (1985), 309
13 J. Maier, Mater. Chem. Phys., 17 (1987), 485
14 B. K. Money, K. Hariharan, Solid State Ionics, 179 (2008), 1273
15 P. Kumar, P. K. Singh, B. Bhattacharya, Ionics, 17 (2011), 721
16 P. K. Singh, B. Bhattacharya, R. M. Mehra, H. W. Rhee, Current
Applied Physics, 11 (2011), 616
17 M. Sotelo-Lerma, M. A. Quevedo-Lopez, R. A. Orozco-Teran, R.
Ramirez-Bon, F. J. Espinoza-Beltran, J. Phys. Chem. Solids, 59
(1998), 145
18 P. K. Singh, P. Kumar, T. Seth, H. W. Rhee, B. Bhattacharya, J. Phys.
Chem. Solids, 73 (2012), 1159
19 S. D. Ross, Inorganic Infrared and Raman Spectra, McGraw-Hill
Book Company Ltd, London 1972
20 N. Srivastava, S. Chandra, European Polymer Journal, 36 (2000),
421
A. SACHDEVA et al.: SYNTHESIS, CHARACTERIZATION AND SENSING APPLICATION OF A SOLID ALUM/FLY ASH ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 467–471 471

J. LÓPEZ GARCÍA et al.: DIFFERENT SOURCE ATELOCOLLAGEN THIN FILMS: PREPARATION, PROCESS ...
DIFFERENT SOURCE ATELOCOLLAGEN THIN FILMS:
PREPARATION, PROCESS OPTIMISATION AND ITS
INFLUENCE ON THE INTERACTION WITH
EUKARYOTIC CELLS
RAZLI^EN IZVIR ATELOKOLAGENSKIH TANKIH PLASTI:
PRIPRAVA, OPTIMIZACIJA PROCESA IN VPLIV NA
INTERAKCIJE Z EVKARIONTI^NIMI CELICAMI
Jorge López García1, Petr Humpolí~ek1, Marián Lehocký1, Ita Junkar2, Miran Mozeti~2
1Centre of Polymer Systems, Tomas Bata University in Zlín, nám. T. G. Masaryka 5555, 760 01 Zlín, Czech Republic
2Department of Surface Engineering, Plasma Laboratory, Jo`ef Stefan Institute, Jamova cesta 39, 1000 Ljubljana, Slovenia
lehocky@post.cz
Prejem rokopisa – received: 2012-11-16; sprejem za objavo – accepted for publication: 2013-01-09
Collagen thin films were prepared via bovine atelocollagen matrices. The film casting was carried out by using different culture
dishes, concentrations, equipment, drying processes and periods of time. In order to optimise the repeatability and
reproducibility, microscopic analyses were utilised to explore the film quality and topographical patterning. In addition, the
human immortalised non-tumorigenic keratinocyte cell line (HaCaT) was seeded onto the obtained specimens, and the cell
proliferation was determined by using the MTT assay. These results indicated how the substrate, its concentration and
processing conditions influence the cellular response. The attempted technique shows itself to be an excellent procedure for
continuous collagen film preparation with optimal cell-proliferation rates, which may potentially be used in tissue engineering
or wound-healing applications.
Keywords: atelocollagen thin films, film optimisation, film quality, surface topography, eukaryotic cell response
Kolagenska tanka plast je bila narejena z govejo atelokolagensko matriko. Priprava plasti je potekala na razli~nih goji{~ih, pri
razli~nih koncentracijah, opremi, postopkih su{enja in pri razli~nih ~asih. Za doseganje optimalne ponovljivosti in obnovljivosti
so bile uporabljene mikroskopske analize, s katerimi je bila ugotovljena kakovost plasti in njihov topografski vzorec. Poleg tega
so bile na vzorce nanesene imortalizirane nekancerogene celice linije keratinocita (HaCaT), njihova proliferacija pa je bila
ugotovljena z MTT-preizkusom. Rezultati prikazujejo, kako podlaga, njena koncentracija ter procesne razmere vplivajo na
biolo{ki odziv. Tovrstna tehnika je edinstven postopek za izdelavo kontinuirnih kolagenskih plasti za optimalno proliferacijo
celic, ki je potencialno uporabna v tkivnem in`enirstvu ali za celjenje ran.
Klju~ne besede: atelokolagenske tanke plasti, optimizacija plasti, kvaliteta plasti, topografija povr{ine, odziv evkarionti~nih celic
1 INTRODUCTION
Collagen is a fibrous protein that is present in nearly
all mammalian tissues. It constitutes approximately 25 %
of the whole-body protein content1 and its abundance is
mainly centred on connective tissues, such as tendons,
ligaments and cartilage. Skin also contains this protein,
which is involved in prime biological functions, such as
tissue formation, cell attachment and proliferation.2
Around 19 proteins are classified as collagen. Moreover,
there are several proteins that have collagen domains.3
Although the 20 standard amino acids are encoded in
collagen biosynthesis, the general sequence is (X-Y-Gly-
cine)n, where proline is recurrently in the X-position, and
4-hydroxyproline, which is almost unique to collagen, in
the Y-position of the sequence (Figure 1a).4,5
This protein is made up of three polypeptide strands,
each chain is a left-handed helix, and the three chains are
coiled around each other in a right-handed super-helix.6,7
As a biomaterial for industrial applications, collagen has
been widely used in many fields, such as cell cultures,
cosmetic and food products and medical devices.8
As for medical applications, this protein is consi-
dered as a primary source in biomedical applications and
one of the most useful biomaterials because of its
excellent biocompatibility, immunogenicity and biode-
gradability.9–11 Unfortunately, native collagen has some
difficulties with its tractability. For example, it may not
be processed by injection moulding or any conventional
extrusion technique.12
Collagen treated either by enzymatic digestion or by
salt/acid extraction is so-called atelocollagen. This treat-
ment leads towards crosslinks breaking amongst the
collagen molecules, resulting in soluble triple helices
lacking of telopeptides, which has the same physical
properties as untreated collagen (Figure 1b).13
Atelocollagen possesses enormous assets; for in-
stance, it is soluble in an acid pH, and its liquid form is
tractable and free of telopeptides, which guaranties a low
immunogenicity. Thus, it is employed for a wide range
of purposes, including wound healing, vessel prostheses,
haemostatic agent and tissue engineering.14–18
Despite the plethora of collagen-related publications,
and to the best of our knowledge, there is no general
Materiali in tehnologije / Materials and technology 47 (2013) 4, 473–479 473
UDK 577:532.6 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)473(2013)
description of the preparation of atelocollagen thin films.
Therefore, the aim of this contribution is the preparation
and optimisation of collagen thin films. These findings
seek to enlighten some of the issues involved in collagen
sheet forming, film casting and the use of atelocollagen
as a biomaterial.
2 PREPARATION OF ATELOCOLLAGEN FILMS
Selecting a proper collagen source is the most im-
portant issue before preparing collagen films. Four sorts
of atelocollagen were employed across the trial, i.e., an
emulsion of atelocollagen from bovine Achilles tendon,
which contents 1.4 % of atelocollagen with a pH of 3.5;
a gel of atelocollagen from bovine splits, 16.2 % of
atelocollagen, pH 5.16 (Splits 1); a gel of atelocollagen
from bovine splits, 11.8 % of atelocollagen, pH 3.7
(Splits 2) and an emulsion of atelocollagen from bovine
tendons, 3.6 % of atelocollagen, pH 3.3. These starting
materials were provided by Vipo A. S., and all of them
were produced by trypsin enzymatic digestion. Three
main factors were taken into account: atelocollagen
concentration, processing factors such as mixing and
casting solution volume. Each atelocollagen source was
solubilised in a aqueous acetic acid solution 0.1M to
finally obtain the mass fractions (0.5, 0.25, 0.1 and
0.05) %. A solution of acetic acid 0.1M was used as a
solvent throughout this study; on account of it being
recommended and employed in previous researches.19–21
It is because pH values within 1.0 and 4.0 have the best
collagen-solubility yields.22,23 The measured pH of the
experimental solution was between 3.5 and 4.0.
The tested samples were homogenised and casted in
culture dishes by using stirring machines with distinct
rotational frequencies and pouring volumes. An IKA
RCT stirring machine (IKA® works, Inc, Germany) and
Merci 1500 (Merci S. R. O, Czech Republic) were used
for the experiments.
A rotational frequency of 1000 r/min was set as the
appropriate one, since this speed provided an efficient
flock reduction. Higher frequencies may cause sample
spilling, whereas lower frequencies are insufficient for
the appropriate preparation of films within a reasonable
processing time. The mixture is referred as the state
formed by a complex of two or more ingredients, whilst
mixing is an operation intended to reduce the
non-uniformity of a mixture and it may be achieved by
inducing the physical motion of the ingredients. The
mixing time was a crucial factor; ensuring 4 h of mixing,
the solution was homogeneous and without atelocollagen
flocks. After an hour of mixing, medium-size flocks
were visible in the solution, and those were progressively
reducing their sizes. Nevertheless, the turbidimetry assay
that measures transmitted light, and is directly propor-
tional to the concentration and depth of the dispersion,
indicates that approximately one hour is enough for
serving, since the transmittance started to be constant
(Figure 2).
This information may be corroborated either with
Scanning Electron Microscopy (SEM) images, which
illustrate the relation between surface-quality and mixing
time (Figure 3), or in Table 1, which lists the percentage
of transmittance before beginning the enquiry.
Table 1: Mass fractions of transmittance with respect to concentration
(w/%)
Tabela 1: Dele` prepustnosti glede na koncentracijo v masni dele`ih
(w/%)
Source 0.5 % 0.25 % 0.1 % 0.05 %
Atelocollagen from
bovine splits 1 80.1 73.4 51.2 28.5
Atelocollagen from
bovine splits 2 91.2 73.6 56.5 32.0
Atelocollagen from
bovine tendons 87.7 72.7 50.4 27.8
J. LÓPEZ GARCÍA et al.: DIFFERENT SOURCE ATELOCOLLAGEN THIN FILMS: PREPARATION, PROCESS ...
474 Materiali in tehnologije / Materials and technology 47 (2013) 4, 473–479
Figure 2: Turbidimetry assay of different atelocollagen solutions,
studied concentration 0.1 %
Slika 2: Preizkus turbidimetrije razli~nih atelokolagenskih raztopin;
raziskovana koncentracija 0,1 %
Figure 1: a) Collagen typical primary sequence; b) atelocollagen via
enzymatic digestion
Slika 1: a) Zna~ilno primarno zaporedje kolagena; b) atelokolagen z
encimsko razgradnjo
The turbidimetry assay was carried out on a Helios
-spectrophotometer (Thermo scientific, USA) at 422
nm.
All the atelocollagen solutions were disposed into the
culture dishes (TPP, Switzerland) 30 min after mixing, to
avoid bubbles coming from the solution stirring, which
alter surface texture. When the solutions were casted
immediately, the subsequent films had a rough appear-
ance, along with collagen congregates, which indeed
deteriorate the surface quality (Figure 4a).
The employed culture dishes were of (9.0, 6.0 and
4.0) cm diameter. The first attempt involved the 9.0 cm
ones and volumes of (30, 20, 10 and 5) mL of the
solution were casted into. The casted films were kept
standing at lab temperature (20–24 °C) for three days to
evaporate the excess solvent.
Casted solutions of 20 mL and 30 mL are also
suitable for making collagen films; nonetheless, those
volumes need more time for optimal solvent evaporation.
If 10 mL is casted, its films get their shapes in two days,
but the solvent still remains (eye observation and sensory
evaluation). The extra day is for completing the solvent
evaporation.
A total of 5 mL was discarded as a pouring volume,
because it scarcely took the shape of the 9.0 cm Petri
dishes. Then, it was decided that 10 mL should be the
starting volume for the culture dishes of this diameter.
The vacuum and the temperature may be utilised for
film drying, thereby reducing the evaporation time. How-
ever, fast solvent evaporation and vacuum influence the
surface smoothness (Figures 4b and 4c).
With reference to temperature, it may not exceed
30–35°C, which is on average the denaturation tempera-
ture of collagen acid solutions, and refers to the collapse
of the collagen triple helix to a random coil configu-
ration.24–27 As the vapour pressure rises, the surface
roughness increases, and to render smooth films, it is
extremely important to focus on the employed solvent
and its evaporation. Solvents with low evaporation rates
tend to produce better quality films than fast evaporating
ones.28,29
3 SURFACE MORPHOLOGY
The specimens made from 10 mL had the best film
appearance (dry, smooth, resilient and the solvent
seemed to be thoroughly evaporated). For example, the
digital camera images depict a surface that apparently
has neither any eye-visible damage nor scratches (Figure
5). This assessment is reinforced by the SEM images,
where the surfaces are relatively smooth, except for some
J. LÓPEZ GARCÍA et al.: DIFFERENT SOURCE ATELOCOLLAGEN THIN FILMS: PREPARATION, PROCESS ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 473–479 475
Figure 4: a) Atelocollagen film obtained from a solution of 0.1 % of the mass fraction atelocollagen from bovine tendons that was poured
immediately after mixing; b) atelocollagen films cast and dried at 30 °C of 0.1 % atelocollagen from bovine tendons; c) by using a desiccator of
0.1 % atelocollagen from bovine splits 1
Slika 4: a) Atelokolagenske plasti, pridobljene iz 0,1 % masnega dele`a raztopine atelokolagena iz goveje kite, ki je bil nanesen takoj po me-
{anju; b) ulita in su{ena atelokolagenska plast pri 30 °C iz 0,1 % atelokolagena iz goveje kite; c) z uporabo eksikatorja iz 0,1 % atelokolagena iz
govejega dela 1
Figure 3: Mixing progress: emulsion 0.1 % of atelocollagen from
bovine tendons. SEM images taken at different mixing times.
Slika 3: Proces me{anja: emulzija 0,1 % atelokolagena iz goveje kite.
SEM-posnetki po razli~nih ~asih me{anja.
wavy areas, which may result either from small collagen
flocks or a product of solvent evaporation (Figure 6).
Scanning electron microscopy (SEM) was performed
in a VEGA II LMU microscope (Tescan s. r. o, Czech
Republic) operated in high vacuum/secondary electron
imaging mode at an accelerating voltage of 5 kV. The
specimens were coated with a thin layer of gold/palla-
dium alloy and tilted at 30° to attain a better observation
of the surface topography. The images were taken at a
magnification of 2000-times.
4 THIN-FILM NORMALISATION
To normalise the procedure for multiple diameters,
the height of served solutions in the Petri dishes was
ascertained by means of an equation that signifies the volume of a cylinder v = r2h, where v is the volume, r
is the radius and h is the height. The idea was to keep the
same height no matter the diameter (twice the radius) of
the culture dishes, i.e., same height, distinct radii and
thus, different volumes. The height was calculated by
using 10 cm3 as an initial volume. As was mentioned,
10 mL (equivalent to 10 cm3) was the selected volume
for the 9.0 cm diameter. The obtained volumes from the
previously calculated height were 4.5 mL and 2.0 mL for
the 6.0 cm and 4.0 cm diameters, respectively.
In order to get the same film thickness, the height has
to be settled as a constant. The model may be applied to
any Petri dish, as it is described in Figure 7a or nume-
rous forms that have a known area. Figure 7b shows the
model that may be used for rectangular sheets. In any
rectangular cuboid form, the volume is defined as v =
lwh, where l is the length, w is the width and h is the
height. If the height (film thickness) remains steady and
the length and width (area) are known, the volume may
be ascertained.
This trial was conducted by following two funda-
mental concepts: repeatability and reproducibility. The
former is the effort to keep constant conditions by using
the same instruments in a short period of time; and the
latter is the one with different instruments and other
periods of time.
J. LÓPEZ GARCÍA et al.: DIFFERENT SOURCE ATELOCOLLAGEN THIN FILMS: PREPARATION, PROCESS ...
476 Materiali in tehnologije / Materials and technology 47 (2013) 4, 473–479
Figure 6: Surface morphology of the atelocollagen thin films from: a)
bovine splits 1; b) bovine splits 2; c) bovine tendons; d) bovine tendons
Slika 6: Morfologija povr{ine atelokolagenske plasti iz: a) govejega
dela 1; b) govejega dela 2; c) goveje kite; d) goveje kite
Figure 7: a) Model for Petri dishes; b) model for rectangular cuboids
sheets
Slika 7: a) Model za petrijevke; b) model za pravokotne kubi~ne liste
Figure 5: Glimpse at the thin films of 0.1 % of the mass fraction of
atelocollagen from bovine splits 2 by a standard camera with the
corresponding brightness and contrast adjustments
Slika 5: Pogled na tanko plast 0,1 % masnega dele`a atelokolagena iz
govejega dela 2 s standardno kamero s primerno nastavljeno svetlostjo
in kontrastom
5 CELL RESPONSE
A human immortalised non-tumorigenic keratinocyte
cell line was supplied by Cell Lines Service (Catalogue
No. 300493, Germany). Dulbecco’s modified eagle
medium-high glucose, supplemented with 10 % foetal
bovine serum and Penicillin/Streptomycin, 100 U/mL
(100 μg/mL) respectively (PAA Laboratories GmbH,
Austria) was used as a culture medium.30 HaCaT cells in
the exponential growth phase were seeded onto the thin
films at a concentration of cells 1 × 105 mL–1 and incu-
bated at 37 °C with 5 % CO2 in humidified air.
The cell proliferation was determined after 4 d in the
culture with the MTT cell proliferation assay kit (Invitro-
gen Corporation, USA). The formazan concentration was
measured in a Sunrise microplate absorbance reader
(Tecan, Switzerland) at 570 nm. The photomicrographs
were taken by using an inverted phase-contrast micro-
scope Olympus CKX41 (Olympus, Germany) with an
optical zoom of 40-times.
The HaCaT keratinocyte cells behaviour on the
prepared films was evaluated with the MTT assay and
the results are given in Figure 8. It was found that after
four days of cultivation, in terms of cost-effectiveness
the films made from 0.1 % were the most suitable.
Regardless of the atelocollagen matrix, the cell prolife-
ration on the films prepared at this concentration exhibits
the highest rate. In contrast, higher concentrations of
collagen (0.5 % and 0.25 %) seem to be detrimental to
the keratinocytes cells.
As far as surface morphology is concerned, the films
that were dried at 30 °C, (with faster solvent evaporation
and rougher surfaces than the lab temperature ones) also
had thriving cell proliferation rates (Figure 8g). Conse-
quently, film drying at this temperature may be used
without affecting atelocollagen-scaffold quality at all. It
was previously demonstrated that the keratinocyte cells
adhere and proliferate satisfactorily on both smooth and
rough surfaces, and also the pivotal role of chemistry and
topography as cell regulatory factors.31–33
This outcome was also qualitatively appraised by the
photomicrographs in Figure 9, which show the cell
aggregates that are adhered on the film surfaces. It may
be seen that the film obtained from 0.5 % has a vast
amount of collagen flocks, which presumably hinder the
HaCaT cell growth.
Nevertheless, the mechanisms of keratinocytes adhe-
sion and proliferation are still unclear, even though it is
well known that hydrophilic surfaces (like these ones)
are favourable for HaCaT cell growth.34–36 For this
reason, a protocol for continuous and effective substrate
J. LÓPEZ GARCÍA et al.: DIFFERENT SOURCE ATELOCOLLAGEN THIN FILMS: PREPARATION, PROCESS ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 473–479 477
Figure 8: Comparison of HaCaT cell growth measured by MTT assay
at 570 nm on: a) 0.5 %; b) 0.25 %; c) 0.05 %; d), e) and f) 0.1 % of the
mass fraction of atelocollagen from bovine tendons, atelocollagen
from bovine splits 1 and atelocollagen from bovine splits 2
respectively; g) 0.1 % film, dried at 30 °C. Error bars signify standard
deviations.
Slika 8: Primerjava rasti celic HaCaT, izmerjene z MTT-preizkusom
pri 570 nm: a) 0,5 %; b) 0,25 %; c) 0,05 %; d), e) in f) 0,1 % masnih
dele`ev atelokolagena iz goveje kite, atelokolagena iz govejega dela 1
in atelokolagena iz govejega dela 2; g) 0,1-odstotna plast, su{ena pri
30 °C. Odmiki na grafu so prikaz standardne deviacije.
Figure 9: Photomicrographs of human skin HaCaT keratinocytes in
culture upon collagen compared with control: a) 0.5 %; b) 0.25 %; c)
0.05 %; d), e) and f) 0.1 % of the mass fraction of atelocollagen from
bovine tendons, atelocollagen from bovine splits 1 and atelocollagen
from bovine splits 2; g) 0.1 % film, dried at 30 °C
Slika 9: Slike keratinocitov iz ~love{ke ko`e HaCaT v kulturi s kola-
genom v primerjavi s kontrolo: a) 0,5 %; b) 0,25 %; c) 0,05 %; d), e)
in f) 0,1 % masnih dele`ev atelokolagena iz govejega kite, atelo-
kolagena iz govejega dela 1 in atelokolagena iz govejega dela 2; g)
0,1-odstotna plast, su{ena pri 30 °C
preparation is of paramount importance. Atelocollagen is
certainly a potential scaffold for cell growth purposes,
which also has the advantage of being eliminated by
degradation processes similar to the metabolism of
endogenous collagen.37
6 CONCLUSIONS
In this contribution, the adopted technique for
atelocollagen thin-film preparation was confirmed to be
a cost-efficient procedure, independent of the equipment
or the time. It increases in importance due to the difficul-
ties that collagen has for its manipulation. The materials
referred to herein as šatelocollagen films’ were
succinctly characterised by microscopic analysis. In
addition, HaCaT keratinocytes cell adhesion was
successfully accomplished. This effort may be adjusted
to diverse atelocollagen matrices. Hence, the present
approach strengthens the knowledge in the use of
atelocollagen as a prospective scaffold, and indeed in the
development of suitable materials for tissue-engineering
and wound-healing applications.
Acknowledgements
The authors would like to express their gratitude to
the Ministry of Education, Youth and Sport of the Czech
Republic (CZ.1.05/2.1.00/03.0111). The Slovenia
Ministry of Higher Education, Science, and Technology
(Program P2-0082-2) and Ad Futura L7-4009 are also
gratefully acknowledged for the financial support of this
research.
7 REFERENCES
1 M. D. Shoulders, R. T. Raines, Annu. Rev. Biochem., 78 (2009)
929–958
2 D. J. Prockop, K. I. Kivirikko, Annu. Rev. Biochem., 64 (1995)
403–34
3 H. W. T. Matthew, Polymers for Tissue Engineering Scaffolds, In: S.
Dumitriu, editor, Polymeric biomaterials, CRC Press, Boca Raton
2001, 167–180
4 N. C. Avery, T. J. Sims, C. Karkup, A. J. Bailey, Meat Sci., 42 (1996)
355–369
5 R. Berisio, L. Vitagliano, L. Mazzarella, A. Zagari, Protein Sci., 11
(2001), 262–270
6 A. Rich, F. H. C., Nature, 176 (1955), 915–916
7 K. Okuyama, Connect. Tissue Res., 49 (2008), 299–310
8 R. Parenteau-Bareil, R. Gauvin, F. Berthod, Materials, 3 (2010),
1863–1887
9 N. Nambu, S. Kishimoto, S. Nakamura, H. Mizuno, S. Yanagi-
bayashi, N. Yamamoto, R. Azuma, S. Nakamura, T. Kizosawa, M.
Ishihara, Y. Kanatani, Accelerated wound healing in healing-
impaired db/db mice by autologous adipose tissue-derived stromal
cells combined with atelocollagen matrix, Ann. Plas. Surg., 62
(2009), 317–321
10 J. L. Garcia, J. Pacherník, M. Lehocký, I. Junkar, P. Humpolí~ek, P.
Sáha, Enhanced keratinocyte cell attachment to atelocollagen thin
films through air and nitrogen plasma treatment, Prog. Colloid
Polym. Sci., 138 (2011), 89–94
11 I. Banerjee, D. Mishra, T. Das, S. Maiti, T. K. Maiti, Caprine (Goat)
collagen: A potential biomaterial for skin tissue engineering, J
Biomater. Sci. Polym. Ed., 23 (2012), 355–373
12 A. Vesel, K. Eler{i~, M. Mozeti~, Immobilization of protein strepta-
vidin to the surface of PMMA polymer, Vacuum, 86 (2012), 773–775
13 H. Fasl, L. F. Zemljic, W. Goessler, K. Stana - Kleinschek, V. Ri-
bitsch, Investigation into amphiphilic chitosan: properties and availa-
bility of original and newly introduced functional groups, Macro-
molecular Chemistry and Physics, 15 (2012), 1582–1589
14 F. Silver, A. Garg, Collagen: Characterization, processing and
medical applications, In: A. Domb, J. Kost, D. Wiserman (editors),
Handbook of biodegradable polymers, Part of the book series, Drug
targeting and delivery, Harwood Academic Publishers, Amsterdam
1997, 319–346
15 S. Inaba, S. Nagahara, N. Makita, Y. Tarumi, T. Ishimoto, S. Matsuo,
K. Kadomatsu, Y. Takei, Atelocollagen-mediated systemic delivery
prevents immunostimulatory adverse effects of siRNA in mammals,
Mol Ther, 20 (2012), 356–366
16 F. Takeshita, Y. Minakuchi, S. Nagahara, K. Honma, H. Sasaki, K.
Hirai, T. Teratani, N. Namatame, Y. Yamamoto, K. Hanai, T. Kato,
A. Sano, T. Ochiya, I. M. Verma, Efficient delivery of small inter-
fering RNA to bone-metastatic tumors by using atelocollagen in
vivo, Proc. Nat. Acad. Sci. USA, 102 (2005), 12177–12182
17 A. Sano, M. Maeda, S. Nagahara, T. Ochiya, K. Honma, H. Itoh, T.
Miyata, K. Fujioka, Atelocollagen for protein and gene delivery,
Adv. Drug Delivery Rev., 55 (2003), 1651–1677
18 Y. Tanaka, H. Yamaoka, S. Nishizawa, S. Nagata, T. Ogasawara, Y.
Asawa, Y. Fujihara, T. Takato, K. Hoshi, The optimization of porous
polymeric scaffolds for chondrocyte/atelocollagen based tissue-engi-
neered cartilage, Biomaterials, 31 (2010), 4506–4516
19 H. Yamaoka, Y. Tanaka, S. Nishizawa, Y. Asawa, T. Takato, K.
Hoshi, The application of atelocollagen gel in combination with
porous scaffolds for cartilage tissue engineering and its suitable
conditions, J Biomed. Mater. Res., Part A, 93 (2010), 123–132
20 A. Bernal, R. Balková, I. Kuøítka, P. Sáha, Preparation and charac-
terisation of a new double-sided bio-artificial material prepared by
casting of poly(vinyl alcohol) on collagen, Polym Bull., 70 (2013) 2,
431–453
21 J. Einbinder, M. Schubert, Binding of mucopolysaccharides and dyes
by collagen, J. Biol. Chem., 188 (1951), 335–341
22 H. Sage, R. G. Woodbury, P. Bornstein, Structural studies on human
type IV collagen, J. Biol. Chem., 254 (1979), 9893–9900
23 O. Harada, K. Kadota, T. Yamamoto, Collagen-based new biome-
dical films: Synthesis, property and cell adhesion, J. Appl. Polym.
Sci., 81 (2000), 2433–2438
24 F. Vojdani, Solubility, In: G. M. Hall (editor), Methods of testing
protein functionality, Champan and Hall, London 1996, 11–60
25 H. Tabarestani, Y. Maghsoudlou, A. Motamedzadegan, A. R. Sade-
ghi, H. Rostamzad, Study on some properties of acid-soluble colla-
gens isolated from fish skin and bones of rainbow trout (Onchorhyn-
chus mykiss), Int. Food Res. J., 19 (2012), 251–257
26 Y. Nomura, S. Toki, Y. Ishii, K. Shirai, The physicochemical pro-
perty of shark type I collagen gel and membrane, J. Agric. Food
Chem., 48 (2000), 2028–2032
27 A. A. S. Machado, V. C. A. Martins, A. M. G. Plepis, Thermal and
rheological behavior of collagen chitosan blends, J. Thermal Anal.
Cal., 67 (2002), 491–498
28 Z. Zhang, G. Li, B. Shi, Physicochemical properties of collagen,
gelatin and collagen hydrolysate derived from bovine limed split
wastes, J. Soc. Leath. Technol. Chem., 90 (2005), 23–28
29 G. Lai, Y. Li, G. Li, Effect of concentration and temperature on the
rheological behavior of collagen solution, Int. J. Biol. Macromol., 42
(2008), 285–2891
30 P. Müller-Buschbaum, J. S. Gutmann, M. Wolkenhauer, J. Kraus, M.
Stamm, D. Smilgies, W. Petry, Solvent-Induced Surface Morphology
of Thin Polymer Films, Macromolecules, 34 (2001), 1369–1375
J. LÓPEZ GARCÍA et al.: DIFFERENT SOURCE ATELOCOLLAGEN THIN FILMS: PREPARATION, PROCESS ...
478 Materiali in tehnologije / Materials and technology 47 (2013) 4, 473–479
31 T. To, H. Wang, A. B. Djuri{ic, M. H. Xiea, W. K. Chan, M. H. Xie,
W. K. Chan, Z. Xie, C. Wud, S. Y. Tonge, Solvent dependence of the
evolution of the surface morphology of thin asymmetric diblock
copolymer films, Thin Solid Films, 467 (2004), 59–65
32 P. Boukamp, R. T. Petrussevska, D. Breitkreutz, J. Hornung, A.
Markham, Normal keratinization in a spontaneously immortalized
aneuploid keratinocyte cell line, J. Cell Biol., 106 (1998), 761–771
33 R. G. Flemming, C. J. Murphy, G. A. Abrams, S. L. Goodman, P. F.
Nealey, Effects if synthetic micro- and nano-structured surfaces on
cell behavior, Biomaterials, 20 (1999), 573–588
34 J. L. Garcia, A. Asadinezhad, J. Pacherník, M. Lehocký, I. Junkar, P.
Humpolí~ek, P. Sáha, P. Valá{ek, Cell proliferation of HaCaT kera-
tinocytes on collagen films modified by argon plasma treatment,
Molecules, 15 (2010), 2845–2856
35 N. Roy, N. Saha, T. Kitano, M. Lehocký, E. Vitková, P. Sáha, Signi-
ficant characteristics of medical-grade polymer sheets and their
efficiency in protecting hydrogel wound dressings: A soft polymeric
biomaterial, Int. J. Polym. Mater., 61 (2012), 72–88
36 S. L. Ishaug-Riley, L. E. Okun, G. Prado, M. A. Applegate, A. Rat-
cliffe, Human articular chondrocyte adhesion and proliferation on
synthetic biodegradable polymer films, Biomaterials, 20 (1999),
2245–2256
37 M. Lehocký, P. St’ahel, M. Koutný, J. ^ech, J. Institoris, A. Mrá~ek,
Adhesion of Rhodococcus sp S3E2 and Rhodococcus sp S3E3 to
plasma prepared teflon-like and organosilicon surfaces, J. Mater.
Process. Technol., 209 (2009), 2871–2875
J. LÓPEZ GARCÍA et al.: DIFFERENT SOURCE ATELOCOLLAGEN THIN FILMS: PREPARATION, PROCESS ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 473–479 479

S. A. AL-JLIL, M. S. LATIF: EVALUATION OF EQUILIBRIUM ISOTHERM MODELS ...
EVALUATION OF EQUILIBRIUM ISOTHERM MODELS
FOR THE ADSORPTION OF Cu AND Ni FROM
WASTEWATER ON BENTONITE CLAY
OCENA MODELOV RAVNOTE@NIH IZOTERM ZA ADSORPCIJO
Cu IN Ni IZ ODPADNIH VOD NA BENTONITNO GLINO
Saad A. Al-Jlil, Mohammad Sajid Latif
National Center for Water Technology (NCWT), King Abdulaziz City for Science and Technology (KACST), P. O. Box 6086, 11442 Riyadh,
Kingdom of Saudi Arabia
saljlil@kacst.edu.sa
Prejem rokopisa – received: 2012-12-03; sprejem za objavo – accepted for publication: 2012-12-14
A study was carried out to evaluate equilibrium isotherm models for the adsorption of Cu and Ni on bentonite clay from
wastewater. The adsorption of Cu and Ni increased with an increase in the pH and was high at the pH values of around 7. The
adsorption capacity of bentonite clay increased with an increase in the temperature. The maximum adsorption capacity of
bentonite clay was 11.12 mg g–1 and 6.78 mg g–1 for Cu and Ni, respectively, at 20 °C. Among the three equilibrium isotherm
models used, the Langmuir-Freundlich isotherm model described the experimental data well at different temperatures.
Keywords: equilibrium isotherm models, adsorption, copper, nickel, bentonite clay, temperature, wastewaters
Izvr{ena je bila {tudija ocene modelov ravnote`nih izoterm pri adsorpciji Cu in Ni iz odpadnih vod na bentonitno glino.
Adsorpcija bakra in niklja nara{~a z nara{~anjem pH in je velika pri pH vrednostih okrog 7. Adsorpcijska zmogljivost
bentonitne gline se pove~uje z nara{~anjem temperature. Najve~ja zmogljivost adsorpcije bentonitne gline je bila 11,12 mg g–1
in 6,78 mg g–1 za Cu in Ni pri 20 °C. Med tremi modeli ravnote`nih izoterm pa dobro opisuje eksperimentalne podatke pri
razli~nih temperaturah izotermni model Langmuir-Freundlich.
Klju~ne besede: modeli ravnote`nih izoterm, adsorpcija, baker, nikelj, bentonitna glina, temperatura, odpadne vode
1 INTRODUCTION
Wastewater is an important resource of water for
augmenting the existing inadequate fresh-water supplies
for multiple uses other than drinking purposes. Among
the different biological, organic and inorganic pollutants
in wastewater, the presence of trace elements and
heavy-metal ions above the permissible limits pose
environmental hazards. According to FAO (1985), the
maximum permissible limits of Cu and Ni in drinking
water are 1 mg L–1 and 0.015 mg L–1, respectively.1
Previous studies showed that sewage water in Riyadh
contains an appreciable amount of heavy metals.2
Frequently, industrial-waste effluents are discharged into
the drainage channels.3 The sources of heavy metals in
wastewaters, especially the industrial effluents, are metal
plating and ceramics photography,4 household chemicals
in sewage water,5 and corrosion of the pipes of
drinking-water-supply networks as well as leaching of
the chemicals from polyvinyl chloride (PVC) pipes.6,7 To
minimize the environmental hazards, it is important to
remove the heavy-metal ions from the wastewaters and
maintain their concentration within the recommended
limits before their land disposal as required by the
National Regulatory Authority. Currently, among the
various technologies for wastewater treatment, the use of
bentonite clay is a very common and effective method
for the removal of heavy-metal ions, especially Cu and
Ni, from wastewaters.
The main objective of this study was to test and
evaluate the local, natural, Saudi bentonite clay for the
adsorption of copper and nickel ions from wastewaters
using three equilibrium isotherm models at different
temperatures.
2 METHODOLOGY
2.1 Experimental part
a) Material
Adsorbent: The adsorbent used in this research was
the local natural bentonite clay.
Adsorbates: Copper-and-nickel-ion solution prepared
from the copper sulfate and nickel nitrate purified and
supplied by S. Define-Chem. limited (Laboratory Rasa-
yan) was used in the experiment.
b) Procedures
Preparation of copper-and-nickel-ion samples
Stock solutions of Cu and Ni metal ions with a
concentration of 1000 mg L–1 were obtained. A mixture
of Cu and Ni ions was prepared by mixing both metal
ions having the same concentration and then diluted to
obtain the concentrations (50–1000 × 10–6) required for
the use in equilibrium experiments. After that, the initial
concentrations of the metal-ion samples and the final
concentration of the resultant solution (after the com-
pletion of the experiment) were diluted and analyzed
Materiali in tehnologije / Materials and technology 47 (2013) 4, 481–486 481
UDK 544.723:666.322 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)481(2013)
with the atomic-absorption spectroscopy (model AAna-
lyst 700, PerKin Elmer, an atomic absorption spectro-
meter).
c) Equilibrium experiments
The equilibrium isotherm was determined by placing
a constant mass of clay (1 g) with a solute solution 50
mL in glass bottles in a constantly agitating shaker. In
each isotherm run, the solute-solution concentrations
ranged from 50–1000 mg L–1 and the temperatures
ranged from 0–20 °C during the study.
The equilibrium experiments were run up to the state
of equilibrium. After that the samples were filtered using
filter papers, then diluted and the absorbance was
measured using atomic absorption spectroscopy. The
concentration of Cu and Ni ions was calculated from the
absorbance using the calibration curve. The amount of
metal ions adsorbed on the bentonite clay (adsorbent)
was calculated with the mass-balance equation as
follows:
q
V C C
Me
e=
−( )0
(1)
where M is the adsorbent mass (g), V is the solution
volume (L), qe is the adsorbed metal-ion concentration
(mg/g), Co is the initial concentration of metal ions
(mg L–1) and Ce is the metal-ion concentration in a bulk
solution at equilibrium (mg L–1).
2.2 Multicomponent Equilibrium Isotherm Models
Three equilibrium isotherm models were applied to
interpret the experimental data for bentonite clay as an
adsorbent at different temperatures to obtain the opti-
mum values of the equilibrium parameters. These models
are:
a) Extended Langmuir isotherm model
The extended Langmuir isotherm model was applied
to determine the adsorption of copper and nickel ions on
bentonite clay from wastewater. The extended Langmuir
isotherm model can be written as follows:8
q
k C
b C b Ce1
e
e e
=
+ +
⎛
⎝
⎜
⎞
⎠
⎟1 1
1 1 2 21
(2)
q
k C
b C b Ce2
e
e e
=
+ +
⎛
⎝
⎜
⎞
⎠
⎟2 2
1 1 2 21
(3)
The extended Langmuir parameters K1, b1, K2 and b2
can be obtained by using the non-linear regression tech-
nique with equations 2 and 3.
b) Langmuir-Freundlich isotherm model
A combination of Langmuir and Freundlich isotherm
models makes a new model called the Langmuir-Freund-
lich isotherm model.9,10 The Langmuir-Freundlich iso-
therm parameters K1, b1, K2 and b2, n1 and n2 can be
obtained by using the non-linear regression technique
with equations 4 and 5. This form can be written as
follows:
q
k C
b C b C
n
n ne1
e
e e
=
+ +
⎛
⎝
⎜
⎞
⎠
⎟1 1
1
1 1
1
2 2
21
(4)
q
k C
b C b C
n
n ne2
e
e e
=
+ +
⎛
⎝
⎜
⎞
⎠
⎟1 2
2
1 1
1
2 2
21
(5)
c) Multicomponent isotherm model
The required five parameters of the multicomponent
isotherm model are as follows:9,11
q
k C
b C b Cn ne1
e
e e
=
+ +
⎛
⎝
⎜
⎞
⎠
⎟1 1
1 1
1
2 2
21
(6)
q
k C
b C b Cn ne2
e
e e
=
+ +
⎛
⎝
⎜
⎞
⎠
⎟2 2
1 1
1
2 2
21
(7)
The equilibrium constants K1, b1, K2 and b2, n1 and n2
can be obtained with the non-linear regression technique
with equations 6 and 7.
3 RESULTS AND DISCUSSION
3.1 Characterization of Bentonite Clay
Saudi bentonite clay was analyzed with XRF (model
JSX-3201, JEOL, an element analyzer). The chemical
analysis is presented in Table 1. Physical parameters
such as the BET surface area, the pore volume and the
average pore width of bentonite clay were determined
with a surface-area analyzer (model ASAP 2020, Micro-
meritics). The particle density and porosity of solid
materials were measured in the Micromeritics Material
Analysis Laboratory (Norcross, Georgia, U.S.A.) using
the gas pycnometer method (an Accupc 1330 pycno-
meter). The results are shown in Table 2. The XRD
S. A. AL-JLIL, M. S. LATIF: EVALUATION OF EQUILIBRIUM ISOTHERM MODELS ...
482 Materiali in tehnologije / Materials and technology 47 (2013) 4, 481–486
Table 1: Chemical analysis of the Saudi bentonite clay in mass frac-
tions (w/%)
Tabela 1: Kemijska analiza saudijske bentonitne gline v masnih
dele`ih (w/%)
Compound w/% (in clay)
SiO2 55.0 ± 3.0
Al2O3 22.0 ± 2.0
TiO2 1.5 ± 0.25
Fe2O3 5.67 ± 0.5
MgO 2.30 ± 0.45
CaO <2.00
Na2O <2.00
K2O <1.00
P2O5 <0.20
S O-3 0.002
Cl- 0.2
Cr2O3 0.02
Mn2O3 0.03
Loss on ignition 9.80
(model D8AD VANCE, BRUKER) analysis of clay
showed that the bentonite clay contained 80 % of mont-
morillonite, 10 % of kaolinite and 10 % of illite and
quartz.
Table 2: Characteristics of bentonite clay
Tabela 2: Zna~ilnosti bentonitne gline
Characteristic Value
BET surface area 62.5671 m2/g
Pore volume ( p/po = 0.97) 0.098005 cm3/g
Average pore width 6.2656 nm
Average pore diameter 9.5650 nm
Porosity (%) 17
Solid density 2.6253 g/cm3
3.2 pH Effects
The effect of a changing pH from 1–10 on the
adsorption of copper and nickel ions on bentonite clay
was studied. The adsorption of Cu and Ni on bentonite
clay increased with an increase in the pH and its
maximum value was at around 7. At a low pH, the
positive charges increased12 and the adsorption of metal
ions decreased due to an increase in the positive charges
on the clay (Figure 1).
The main components of the Saudi bentonite clay are
SiO2 and Al2O3 (Table 1). The adsorption properties of
clay are influenced by the pH of the solution and the
Si/Al ratio. The effect of pH can be illustrated as
follows.12
where M is the Al or Si atoms in the clay. The iso-
electric point (PI) is the pH, at which there is no net
charge present on the surface of a clay particle. The
value of the isoelectric point for silica is 2 and that of
alumina is 9.12 However, when the Si/Al ratio is bet-
ween 1 and 5, the clay will adsorb both the cation and
anion ions.12
Overall, the adsorption of the Cu and Ni ions with the
negative charge on the clay is due to the attraction of the
electrostatic force between the negative and positive
charges of heavy metals. The negative charges on the
clay surface are largely due to the presence of silica. The
optimum pH value of 7 was used to analyze the effects of
other variables such as temperature.
3.2.1 Equilibrium experiments
A combination of copper and nickel as a multicom-
ponent system was investigated. The equilibrium results
are presented in Figures 2 and 3. The results indicate
that the adsorption capacity of clay was higher for Cu
ions than for Ni ions. The maximum adsorption capacity
of clay was 11.12 mg g–1 at 20 °C for Cu and 6.78 mg/g
at 20 °C for Ni in the batch solution containing both the
S. A. AL-JLIL, M. S. LATIF: EVALUATION OF EQUILIBRIUM ISOTHERM MODELS ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 481–486 483
Figure 3: Equilibrium isotherm for the adsorption of Ni from waste-
water on bentonite clay at different temperatures
Slika 3: Ravnote`na izoterma za adsorpcijo Ni iz odpadne vode na
bentonitno glino pri razli~nih temperaturah
Figure 1: Effect of pH on the adsorption of copper and nickel on ben-
tonite clay
Slika 1: Vpliv pH na adsorpcijo bakra in niklja na bentonitno glino
Figure 2: Equilibrium isotherm for the adsorption of Cu from waste-
water on bentonite clay at different temperatures
Slika 2: Ravnote`na izoterma za adsorpcijo bakra iz odpadne vode na
bentonitno glino pri razli~nih temperaturah
Cu and Ni ions. This may be attributed to the com-
petition between the Cu and Ni ions on the active site on
the clay. Other investigators obtained similar results for a
multicomponent adsorption on the orthophosphate-
modified kaolinite clay.13 The results infer that the
selectivity of bentonite clay for the adsorption of various
ions is Cu > Ni ion. This can be concluded from the data
presented in Table 3.
Table 3: Copper and nickel properties9,10
Tabela 3: Lastnosti bakra in niklja9,10
Property Copper ion Nickel ion
Ionic radius (nm) 0.072 0.069
Atomic mass 63.4 57.8
Coordination number 2, 4 4, 5
Electron configuration [Ar] 3d9 [Ar] 3d8
Electro negativity of the atom 1.90 1.91
This variability for more Cu ion adsorption than Ni
on clay is clear from Table 3, which shows an unpaired
electron for a Cu ion in addition to the Cu ion being
paramagnetic.14 Consequently, the Cu ion can be attrac-
ted by a magnetic field resultant from the clay adsor-
bent15 because the Ni ion is stable due to the absence of
an unpaired electron.
3.3 Equilibrium Experimental Results
a) Extended Langmuir isotherm model
The extended Langmuir parameters K1, b1, K2 and b2
were obtained with the non-linear regression technique
from equations 2 and 3. The equilibrium parameters K1,
b1, K2 and b2 were calculated with the non-linear regres-
sion technique and presented in Tables 4 and 5.
b) Langmuir-Freundlich isotherm model
The parameters of the Langmuir-Freundlich isotherm
model such as K1, b1, K2 and b2, n1 and n2 were obtained
applying equations 4 and 5. The equilibrium parameters
are presented in Tables 6 and 7.
S. A. AL-JLIL, M. S. LATIF: EVALUATION OF EQUILIBRIUM ISOTHERM MODELS ...
484 Materiali in tehnologije / Materials and technology 47 (2013) 4, 481–486
Table 4: Extended Langmuir constants for the Cu-ion adsorption on
clay from a mixture of copper and nickel solution at different
temperatures
Tabela 4: Raz{irjene Langmuirove konstante za adsorpcijo Cu-ionov
na glino iz me{anice raztopine bakra in niklja pri razli~nih tempera-
turah
Temperature
T/oC
K1/
(L/g)
B1/
(L/mg)
AARD/
% R
2 X2
20 0.106 0.007 18.58 0.941 2.41
40 0.1306 0.0089 16.42 0.959 2.06
60 0.167 0.0114 16.88 0.94 2.346
80 0.176 0.012 15.91 0.95 3.41
Table 5: Extended Langmuir constants for the Ni-ion adsorption on
clay from a mixture of copper and nickel solution at different tempera-
tures
Tabela 5: Raz{irjene Langmuirove konstante za adsorpcijo Ni-ionov
na glino iz me{anice raztopine bakra in niklja pri razli~nih tempera-
turah
Temperature
T/oC
K2 /
(L/g)
b2 /
(L/mg)
AARD/
% R
2 X2
20 0.557 0.0002 22.25 0.672 29.23
40 0.068 0.0002 23.07 0.755 25.92
60 0.0843 0.0001 21.76 0.808 35.15
80 0.0865 0.0001 22.40 0.829 37.32
Table 6: Langmuir-Freundlich constants for the Cu-ion adsorption on
clay in a mixture of multicomponent from copper and nickel at
different temperatures
Tabela 6: Langmuir-Freundlich konstante za adsorpcijo Cu-ionov na
glini iz multikomponentne me{anice za baker in nikelj pri razli~nih
temperaturah
Temperature
T/oC
K1/
(L/g)
b1/
(L/mg) n1
AARD/
% R
2 X2
20 0.662 0.0028 0.482 13.51 0.959 0.526
40 0.9778 0.0093 0.4233 11.31 0.964 0.535
60 0.567 0.0046 0.5799 13.57 0.935 1.16
80 0.4623 0.0066 0.7434 13.002 0.969 0.932
Table 7: Langmuir-Freundlich constants for the Ni-ion adsorption on
clay in a mixture of multicomponent from copper and nickel at
different temperatures
Tabela 7: Langmuir-Freundlich konstante za adsorpcijo Ni-ionov na
glino v multikomponentni me{anici iz bakra in niklja pri razli~nih
temperaturah
Temperature
T/oC
K2
/(L/g)
b2/
(L/mg) n2
AARD/
% R
2 X2
20 0.722 0.0135 0.375 12.533 0.892 1.5
40 0.879 0.0005 0.355 16.34 0.901 1.23
60 1.101 0.0300 0.3720 20.2412 0.885 1.386
80 0.1928 0.0108 0.788 17.087 0.874 16.08
Table 8: Multicomponent five-parameter isotherm constants for the
Cu-ion adsorption on clay in a mixture of copper and nickel at diffe-
rent temperatures
Tabela 8: Ve~komponentna petparametri~na izotermna konstanta za
adsorpcijo Cu-ionov na glini iz me{anice bakra in niklja pri razli~nih
temperaturah
Temperature
T/oC
K1/
(L/g)
b1/
(L/mg) n1
AARD/
% R
2 X2
20 1.714 18.098 0.0664 20.254 0.918 3.935
40 1.072 5.599 0.1479 16.71 0.947 2.142
60 1.343 4.761 0.0119 15.49 0.938 1.263
80 1.489 1.279 0.0088 9.057 0.958 0.6468
Table 9: Multicomponent five-parameter isotherm constants for the
Ni-ion adsorption on clay in a mixture of copper and nickel at diffe-
rent temperatures
Tabela 9: Ve~komponentna petparametri~na izotermna konstanta za
adsorpcijo Ni-ionov na glini iz me{anice bakra in niklja pri razli~nih
temperaturah
Temperature
(oC)
K2/
(L/g)
b2/
(L/mg) n2
AARD/
% R
2 X2
20 0.9066 0.0048 1.456 26.598 0.614 38.312
40 0.5629 0.0042 1.397 21.79 0.769 24.854
60 0.6829 0.1277 0.909 20.309 0.8252 23.44
80 0.7455 0.4703 0.7206 20.232 0.873 9.521
c) Multicomponent five-parameter isotherm model
The equilibrium constants for the model such as K1,
b1, K2 and b2, n1 and n2 were obtained applying equations
6 and 7 as shown in Tables 8 and 9.
3.4 Estimation of the Best Fit
The theoretical results and the experimental equili-
brium data were compared using the average absolute
relative deviation percent (AARD/%), the chi-square
method (X2) and the adjusted coefficient of determination
(R2). The AARD was calculated to determine the best fit
between the theoretical and experimental results from the
applied equilibrium isotherm models as summarized
below.
AARD/%=
100
1N
q q
q
i i
ii
N
exp
exp
−
=
∑ calc (8)
where N is the number of data points, qicalc is the cal-
culated amount of the metal-ion adsorption on clay and
qiexp is the experimental amount of the metal-ion
adsorption on clay for a given data point i. Also, the
chi-square method (X2) was applied to evaluate the rela-
tionship between the theoretical and the experimental
data from the equilibrium experiments. The chi-square
(X2) formula is:
(X2) =
( )
exp
q q
q
i i
ii
N −
=
∑ calc
calc
2
1
(9)
where N is the number of data points, qiexp is the
experimental amount of the metal-ion adsorption on
clay and qicalc is the calculated amount of the metal-ion
adsorption on clay for a given data point i. In the case of
the chi-square method (X2), the X2 has a small value, the
result of the model is close to the result of the equili-
brium experiment and vice versa. The adjusted coeffi-
cient of determination, R2, is normally used to evaluate
the best fit.
Significant differences were observed between the
extended Langmuir isotherm model, the Langmuir-
Freundlich isotherm model and the multicomponent
five-parameter isotherm model. The Langmuir-Freund-
lich isotherm model provided the best correlation with
the equilibrium data.
3.5 Numerical Solution of the Non-Linear Isotherm
Models
A non-linear, least-square, data-fitting algorithm
using the fminsearch function from MATLAB was used
to find the optimum values of the isotherm models. Then
the average absolute relative deviation percent
(AARD/%), the adjusted coefficient of determination, R2,
and the chi-square method (X2) were calculated for the
experimental and theoretical values. The program was
written using MATLAB and the flow chart is shown
below:
4 CONCLUSIONS
The study showed that pH is a significant factor in
the adsorption processes as it causes electrostatic
changes in the solutions. The optimum pH value for the
Cu and Ni adsorption is around 7. The maximum
adsorption capacity of clay increased with an increase in
the temperature, which was 11.12 mg g–1 for copper and
6.78 mg g–1 for nickel at 20 °C in the experimental
solutions. Among the various isotherm models, the
Langmuir-Freundlich isotherm model described the
experimental data very well.
Acknowledgements
The authors would like to thank King Abdulaziz City
for Science and Technology (KACST) for the support
and encouragement to carry out the study.
5 REFERENCES
1 M. J. Zamzow, J. E. Murphy, Removal of metal cations from water
using zeolites, Separation Science and Technology, 27 (1992) 14,
1969–1984
2 A. S. Mashhady, Heavy metals extractable from a calcareous soil
treated with sewage sludge, Environmental pollution (series B), 8
(1984), 51
3 A. M. Al-Rehaili, Municipal wastewater treatment and reuse in Saudi
Arabia, The Arabian Journal for Science and Engineering, 22 (1997),
143
S. A. AL-JLIL, M. S. LATIF: EVALUATION OF EQUILIBRIUM ISOTHERM MODELS ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 481–486 485
4 U. Forstner, G. T. Wittmann, Metal pollution in the aquatic environ-
ment, Springer verlag, New York 1979
5 D. Jenkins, Effect of reformulation of household powder laundry
detergents on their contribution to heavy metals levels in wastewater,
Water Environment Research, 70 (1998) 5, 980
6 A. I. Alabdula’aly, Trace metals in Riyadh public water supplies, The
Arabian Journal for Science and Engineering, 22 (1997), 165
7 M. Sadiq, G. Hussain, Drinking water quality in Saudi Arabia- an
overview, The Arabian Journal for Science and Engineering, 22
(1997), 153
8 G. Mckay, Use of adsorbents for the removal of pollutants from
wastewater, CRC, Press, Inc., 1996
9 W. Fritz, E. U. Schuender, Simulation adsorption equilibria of
organic solutes in dilute aqueous solution on activated carbon,
Chemical Engineering Series, 29 (1974), 1279
10 A. I. Liapis, D. W. T. Rippin, A general model for the simulation of
multi-component adsorption from a finite batch, Chemical Engi-
neering science, 32 (1977), 619
11 C. M. Yon, P. H. Turnock, Multicomponent adsorption equilibria on
molecular sieves, American Institute Chemical Engineers, Sympo-
sium, 67 (1971), 75
12 H. A. Elliott, C. P. Huang, Adsorption characteristics of some Cu(II)
complexes on aluminosilicates, Water research, 15 (1981), 849–855
13 K. O. Adebowale, I. E. Unuabonah, B. I. Olu-Owolabi, The effect of
some operating variables on the adsorption of lead and cadmium ions
on kaolinite clay, Journal of Hazard Material, 134 (2006), 130
14 K. F. Purcell, J. C. Kotz, An Introduction to Inorganic Chemistry,
Saunders College Publishing, Philadelphia 1980
15 K. H. Chong, B. Volesky, Metal biosorption equilibria in a ternary
system, Biotechnology and Bioengineering, 49 (1996), 629
486 Materiali in tehnologije / Materials and technology 47 (2013) 4, 481–486
S. A. AL-JLIL, M. S. LATIF: EVALUATION OF EQUILIBRIUM ISOTHERM MODELS ...
T. KOKALJ et al.: MODELLING AND OPTIMIZATION OF A LASER DROPLET-FORMATION PROCESS
MODELLING AND OPTIMIZATION OF A LASER
DROPLET-FORMATION PROCESS
MODELIRANJE IN OPTIMIZACIJA LASERSKEGA TVORJENJA
KAPLJICE
Tadej Kokalj1, Igor Grabec2, Edvard Govekar2
1Institute of Metals and Technology, Lepi pot 11, 1000 Ljubljana, Slovenia
2Faculty of Mechanical Engineering, University of Ljubljana, A{ker~eva 6, 1000 Ljubljana, Slovenia
tadej.kokalj@imt.si
Prejem rokopisa – received: 2013-04-12; sprejem za objavo – accepted for publication: 2013-05-06
The laser droplet-formation process (LDFP) is a part of the novel joining technology for forming high-temperature joints. The
advantages of this technology are: good heat-input control, good control of the added material and limited local heating. A
molten metal droplet is used as the basic unit for the filling material, and a determination of the process parameters for the
formation of droplets with desired properties is crucial. A physical and numerical model of the process was built to allow a
theoretical determination of the process parameters. The numerical model enables a simulation of the process at different sets of
parameters and a genetic algorithm optimization was implemented to find the best set. To verify this general procedure for
determination and optimization of the parameters, we applied it in a specific case of a nickel wire. On the basis of the numerical
model of the process, laser pulses for pendant-droplet formation and droplet detachment were determined and applied in
experiments. With numerically determined laser pulses pendant-droplet formation and detachment were accomplished. In most
cases, the process showed a very high repeatability. The encouraging experimental verification shows that a numerical approach
is a very helpful tool for a determination of the process parameters and a better understanding of the process. It significantly
reduces the number of necessary experiments when the laser droplet-formation setup or the target droplet properties are
changed.
Keywords: droplet formation, laser, modelling, optimization, experimental verification
Proces laserskega tvorjenja kapljice (LTK) je del nove tehnologije spajanja za tvorjenje visokotemperaturnih spojev. Prednosti te
tehnologije so: dobra kontrola vnesene toplote in dodanega materiala ter omejeno lokalno segrevanje podlage. Staljena kovinska
kapljica se uporabi kot dodajni material, parametri za tvorjenje kapljice s primernimi lastnostmi pa so bistveni za uspe{nost
procesa. Zgradili smo fizikalni in numeri~ni model procesa, ki omogo~a teoreti~no dolo~itev optimalnih procesnih parametrov.
Numeri~ni model nam omogo~a simulacijo procesa pri razli~nih naborih parametrov, optimizacijo z genetskimi algoritmi pa
smo uporabili za iskanje najbolj{ega nabora. Za eksperimentalno potrditev tega splo{nega na~ina smo metodo uporabili v
konkretnem primeru nikljeve `ice. Na osnovi teoreti~nega modela smo dolo~ili laserski blisk za tvorjenje in odlet kapljic ter jih
uporabili v poskusih. Z numeri~no dolo~enimi laserskimi bliski smo uspe{no tvorili in lo~ili kapljico od `ice. V ve~ini primerov
je proces izredno dobro ponovljiv. Uspe{na eksperimentalna potrditev potrjuje pomembnost teoreti~nega modela za dolo~itev
procesnih parametrov in bolj{e razumevanje procesa. Numeri~ni model mo~no zmanj{a potrebno {tevilo eksperimentov za
dolo~itev procesnih parametrov ob spremembi vrste `ice ali spremembi ciljnih lastnosti tvorjenih kapljic.
Klju~ne besede: tvorjenje kapljice, laser, modeliranje, optimizacija, eksperimentalna verifikacija
1 INTRODUCTION
A constant technological advance is increasing the
demand for joining new and dissimilar materials, for
which conventional technologies are neither appropriate
nor convenient. The differences in the thermophysical
properties of base materials and the trend of miniaturi-
sation require a better control of the heat input and the
amount of a filler material. In order to improve this con-
trol, droplet-joining technologies are investigated where
a small, molten metal droplet is used as the basic unit of
a filling material. The droplet is deposited on a joining
spot creating a material-to-material joint or a bridge over
a gap in the substrate.1,2
Several droplet-formation technologies have been
proposed for the applications in the joining technologies.
Individual droplets are generated by different "drop-on-
demand" generators.3–5 In the case of low-melting mate-
rials (300 °C to 500 °C), the melted material is provided
from a heated reservoir. A droplet is then generated with
a short pressure pulse.5,6 Produced joints are not suscep-
tible to high temperatures and usually contain lead that is
environmentally disputable. The generation of droplets
from a high-melting material (typically over 1000 °C), is
usually accomplished with local heating of a wire. The
first systems for droplet generation used a pulsed electric
arc between the electrodes and the tip of the wire.7–9 A
molten metal droplet was then detached from the wire by
the electromagnetic force, as a consequence of an
increase in the electric current at the end of the pulse10,11,
which could be accompanied by mechanical oscillations
of the wire.12 The diameter of a droplet in this techno-
logy is limited down to 1 mm.13 A system for the arc-
droplet formation does not allow for a precise energy
control. Besides, an electric arc induces the formation of
hot plasma, which strongly heats the surroundings. The
system has to be intensely cooled and it is not appro-
priate in the cases when the heat susceptibility of the
substrate is limited.
Materiali in tehnologije / Materials and technology 47 (2013) 4, 487–495 487
UDK 621.791.724:004.942 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 47(4)487(2013)
To avoid the listed deficiencies of the formation of
high-temperature joints, a new laser droplet-formation
technology was proposed.14,15 In this novel technology a
laser beam is used for droplet formation and detachment
from the feed wire (Figure 1). The advantage of this
technology is a good heat-input control and limited local
heating enabling a formation of smaller droplets.15–17
The laser droplet-formation process (LDFP) was
proposed as part of a new droplet-joining technology in15
and was first characterized with an IR camera and
acoustic emission signals in14. A detailed characte-
rization of the process on the basis of acoustic emission,
where different phases of the process such as heating,
melting, droplet formation and droplet detachment are
observed, is given in18. A possibility of an indirect
process characterization by means of reflected laser light
is explained in16. The cooling of a droplet after its
detachment and examples of blind joints are discussed
in17 and an application for joining the parts of different
materials and geometrical properties are elaborated in19.
Later a sequential formation of droplets was presented
and the chaotic regimes of sequential-droplet formations
were elaborated and discussed in20. Lately, some promis-
ing studies were done on the application of droplet
welding of coated steel sheets.21
Determining the process parameters, especially the
time course of the laser power, proves to be of the central
importance for the repeatable formation of droplets with
desired properties. The first experimental investigations
showed that it is difficult to achieve a repeatability of the
process and to avoid undesired radial scatter and
splashing of the melt by the heuristically determined
laser pulses. At the beginning, a laser pulse of a constant
power was used with a steel wire, however, the time
course of the laser power and velocity of the wire move-
ment were later identified to be the most important
process parameters. A simple theoretical analysis of the
process22 as well as the experimental results showed that
it is convenient to divide LDFP into two phases: the
pendant droplet formation and the detachment phase.
With this approach the process was significantly im-
proved. However, in order to further improve the control
of droplet properties, like the size, temperature homoge-
neity and heat content, and to avoid undesired splashes, a
numerical model of the process was built.
The numerical model enables the simulation of the
process at different sets of process parameters. With the
use of the numerical model and genetic-algorithm opti-
mization method, the process parameters can be opti-
mized in accordance with the selected optimization
objectives. These optimization objectives incorporate the
desired droplet properties corresponding to the purpose
of an application. The theoretical approach showed to be
very useful not only to get a better insight into the
process, but especially as a general tool to significantly
decrease the number of the experiments needed to find
new process parameters when some of the parameters
are changed (i.e., wire material or diameter, laser optics,
wire velocity, etc.).
The article starts with a description of the physical
and numerical model of the process. An application of
genetic algorithms for determining the pendant-droplet-
formation pulse is described together with the applied
objective functions. At the end of the theoretical part of
the paper the procedure for determining the droplet-
detachment pulse is explained on the basis of a sim-
plified model of the keyhole phenomenon. In the second
part of the paper we focus on the experimental verifi-
cation of theoretical results. The laser droplet-formation
process is explained for a specific case of a nickel wire.
The results of the theoretical treatment with determined
laser pulses and the corresponding simulations are
presented in parallel with the experimental results for
both pendant-droplet formation and detachment of a
droplet. In the last section we discuss the results and
underline the significance of this research for a further
development of the process.
2 THEORETICAL MODEL AND NUMERICAL
MODEL
With the aim of theoretically determining the laser
pulse, a numerical model of the process was built. The
basic equation of this model is the heat equation in cylin-
drical coordinates (r,, z):
T. KOKALJ et al.: MODELLING AND OPTIMIZATION OF A LASER DROPLET-FORMATION PROCESS
488 Materiali in tehnologije / Materials and technology 47 (2013) 4, 487–495
Figure 1: In the laser droplet-formation process (LDFP) a thin metal
wire is fed to the focus of laser beams. A hot molten droplet is formed
at the tip of the wire, detached by a sudden increase in the laser power
and deposited on the substrate.
Slika 1: Pri laserskem tvorjenju kapljice (LTK) dodajamo tanko
kovinsko `ico v gori{~e laserskih `arkov. Vro~o staljeno kapljico, ki
nastane na koncu `ice, lo~imo s kratkotrajnim pove~anjem laserske
mo~i. Kapljica pade na podlago, kjer tvori spoj.
∂
∂
∂
∂
∂
∂
T r z t
t
D T
r r
r
T r z t
r
r
( , , , )
( )
( , , , ) 
= ⎡⎣⎢
⎛
⎝
⎜ ⎞
⎠
⎟ +
+
1
1
2
2 2
2
∂
∂
∂
∂
T r z t T r z t
z
( , , , ) ( , , , )


 +
(1)
where T is the temperature and D(T) = k(T)/
(T)cp(T) is
the temperature-dependant heat-diffusion coefficient. In
this equation the temperature-dependent gradient of the
heat-conduction coefficient k was neglected.
Phase transitions were treated as jumps in specific
heat:
cpt(T) = cp(T) + L(T – Tp) (2)
where cpt is the total specific heat, including the latent
heat L, cp is the specific heat at the instant temperature,
 is the Dirac delta function and Tp is the temperature of
the phase transition. For the numerical approximation of
, a Gaussian function was used:
( ) exp
( )
T T
T
T T
Tp
p
− ≈ −
−⎛
⎝
⎜
⎞
⎠
⎟1
2 2
2
2πΔ Δ
(3)
Here T is the adaptable width of the Gaussian
function.
The temperature of the wire at the beginning of the
process is the same as the environmental temperature T0,
therefore we formulate the initial condition as:
T r T( , )

0 0= (4)
With respect to the boundary conditions we consider
the thermal radiation from the hot surface as:
j r t Tr ( , )

sur =−
4 (5)
Since the radiation flux from the wire is much larger
than the radiation flux from the environment to the wire,
the latter is neglected in this formulation.
The energy flux of the laser beams to the wire is a
function of the spatial part j rl ( )

, the temporal part g(t)
and it depends on the material absorptivity A:
j r t Aj r g tf l( , ) ( ) ( )
 
sur = (6)
The boundary condition on the surface of the wire is
finally formulated as:
k T
T
n
j jf r( )
∂
∂
= − (7)
where k is the thermal conductivity and T/n is the
temperature gradient in the direction perpendicular to
the surface. The sum of thermal fluxes is on the right-
hand side of the equation.
The heat equation including the described initial and
boundary conditions can be solved numerically using an
explicit finite-difference scheme in cylindrical coordi-
nates.23 With the described numerical model the time
development of the temperature field of the wire at
various sets of process parameters can be simulated.
Based on the simulations a proper set of process para-
meters can be determined and optimized. This work is
focused on determining the laser pulse.
3 LASER-PULSE DETERMINATION
The laser pulse for pendant-droplet formation and the
pulse for droplet detachment were determined separately.
In the first part of the process the heat input is essential
and in the second part a sufficient force is needed to
detach the droplet. Although both parts of the process are
coupled, we treated them separately for the sake of
simplicity. The pendant-droplet-formation pulse was
parameterized with seven parameters (Figure 2) and
these were subject to optimization. The droplet-detach-
ment pulse was determined on the basis of a simplified
keyhole-effect model. We need to stress that only the
parameterized laser pulse was optimized in this inve-
stigation, while the other parameters were pre-selected
by selecting the wire, limiting our experimental setup
and the desired size of the droplet and were not subject
to optimization. However, the described optimization
procedure can also be used to optimize other parameters.
3.1 Laser pulse for pendant-droplet formation
In the pendant-droplet-formation phase of the pro-
cess, a well-melted pendant droplet has to be produced at
the tip of the wire. Besides, the splashing of the melt and
the radial scatter of deposited droplets have to be
reduced. Since the droplet properties (i.e., heat content
and temperature homogeneity) are importantly influ-
enced in this part of the process, the laser pulse should
be determined in accordance with the desired droplet
properties.
The selection of a proper method for laser-pulse
determination depends on the time necessary for one
T. KOKALJ et al.: MODELLING AND OPTIMIZATION OF A LASER DROPLET-FORMATION PROCESS
Materiali in tehnologije / Materials and technology 47 (2013) 4, 487–495 489
Figure 2: Simple laser pulse for LDFP consists of two parts:
pendant-droplet-formation part and detachment part (top). In the
present work we parameterized the first part of the pulse with seven
parameters in order to further optimize the heat input (bottom).
Slika 2: Enostaven laserski blisk za LTK je sestavljen iz dveh delov:
del za tvorjenje vise~e pretaljene kapljice in del za lo~itev kapljice
(zgoraj). V predstavljenem delu smo prvi del bliska parametrizirali s
sedmimi parametri za nadaljno optimizacijo vnosa toplote (spodaj).
process simulation. The numerical model was imple-
mented in the C++ language. It is basically 3D and can
be used either for a 2D or 3D simulation of the process.
An increased spatial dimensionality of the model
drastically increases the simulation time, therefore, the
dimensionality of the problem is reduced. Since three
laser beams are positioned equidistantly around the
circumference of the wire, and the diameter of the laser
beam is similar to the wire radius, the process can be
considered as quasi-symmetrical. In this ax-symmetrical
case the angular dependence can be abandoned and a 2D
model is sufficient to describe the process.
A short simulation time allows an implementation of
the optimization method that requires many process si-
mulations. We choose the genetic-algorithm method24–26
for the process optimization.
To perform the optimization, we need to determine
the fitness function that includes the objectives of the
optimization encoded in a single mathematical expres-
sion. We determined two different fitness functions. The
first fitness function was assigned as:
J
m
m
m
m1
1= − +t i (8)
where mt/m is the portion of the melted material, and
mi/m is the portion of the vaporized material. The
objective of this function is to get the highest possible
amount of the melted material at the lowest possible
vaporization. By finding the minimum of this function,
we achieve the best trade-off between the amounts of
the melted and vaporized material at the end of a laser
pulse.
The second fitness function is more complex and
composed of two parts:
J
T t T
m
m
T M
m t
m
j v z
j
Dj
N
z
ik k
Dk
M
2
2
1
1
1
2
1
2
1
=
− ⋅
+ ⋅
=
=
∑
∑
( ( ) )
( )
(9)
where Tj(tv) is the temperature of the j-th volume part of
the wire at the time tv and mj is its mass; mD is the mass
of the observed part of the wire and Tz is the desired
final temperature of the wire; mik is the mass of the
vaporized material at the time tk.
The first part of the fitness function is calculated at
the end of the laser pulse tv and should bring the final
temperature of the wire as close to the desired tempe-
rature Tz as possible. The role of the second part of the
fitness function is to prevent vaporization during the
pulse. It is calculated as the sum of the portions of the
vaporized material at the M = 100 intervals from the
beginning, t = 0, to the end of the laser pulse, t = tv.
With these two fitness functions and the selection of
desired temperatures Tz, different laser pulses for pen-
dant-droplet formation were determined. In addition, the
arbitrary-fitness function can be constructed in order to
meet the demand of a specific application. The opti-
mization was performed by combining the executable
C++ file, the returning value of the fitness function at the
selected set of parameters and MATLAB Genetic Algo-
rithm Toolbox, creating a new set of parameters. The
time for the optimization of one laser pulse was around
20 h including several hundred process simulations.
3.2 Laser pulse for droplet detachment
After a successful formation of a pendant droplet at
the tip of the wire an additional force is needed to over-
come the surface tension force20 and detach the droplet
from the wire. This additional force is provided by an
onset of the keyhole. The keyhole is a highly nonlinear
phenomenon and therefore difficult to model, parti-
cularly in the first transient phase when the keyhole is
being formed. Since the numerical model of the transient
phase of the keyhole effect was not found in the litera-
ture, we built a simplified numerical model of this part of
the process. We treated the keyhole as a fast vaporization
of the material and we neglected the self-focusing
dynamics of the keyhole due to the surface tension. In
this way we could approximately simulate the growth of
the keyhole.
It was shown in22 that a droplet can be detached from
the wire when the depth of the keyhole approximately
reaches the axes of the wire. Therefore, the droplet-
detachment pulse was determined in the following way:
First, a constant laser power was selected for the laser
pulse. Second, the time development of the temperature
field was calculated for the selected laser power and the
temperature on the axis of the wire was checked on each
time step. Third, the calculation was stopped when the
temperature on the axis of the wire reached the boiling
temperature, that is when the keyhole reached the axis of
the wire.
With this procedure the necessary time to form a
keyhole and detach the droplet from the wire, or better,
its upper limit, was determined since the keyhole growth,
considering the self-focusing, was even faster. It is
important to minimize the detachment-pulse time in
order to reduce the unnecessary heating of the wire with
the high-power laser beam, resulting in undesired effects
like vaporization and splashes of the melt.
4 EXPERIMENTS AND RESULTS
The theoretical model, pulse determination and
optimization procedure described above are general and
independent of the wire, laser and wire-feed unit.
However, in order to evaluate the adequacy of theoretical
work, we determined the laser pulses and performed
experiments for a specific case of a nickel wire of a
diameter 0.7 mm using the experimental setup described
below. The results of numerical calculations and
experimental outcomes are described in the following
section.
T. KOKALJ et al.: MODELLING AND OPTIMIZATION OF A LASER DROPLET-FORMATION PROCESS
490 Materiali in tehnologije / Materials and technology 47 (2013) 4, 487–495
4.1 Experimental system
The main parts of the experimental system are: a
laser, a wire-feed unit, an opto-mechanical positioning
system and a measuring computer.
The heart of this system is a Nd:YAG pulse laser
which operates at the wavelength of  = 1064 nm. The
maximum frequency of the laser-pulse generation is  =
300 Hz and the average power is P = 0.25 kW. The
minimum power of the laser pulse is Pmin = 0.48 kW and
the maximum laser power is Pmax = 8 kW. The minimum
and maximum durations of the laser pulse are tmin = 0.3
ms and tmax = 20 ms. The time course of the laser power
of a pulse can be set with a time step of 0.1 ms and a
power step of 40 W.
The wire-feed unit enables the maximum accele-
ration of the wire, amax = 20 m/s2, and the maximum
velocity is vmax = 0.3 m/s. The step size of the wire move
is zmin= 3.1 μm.
For the investigation of LDFP the primary laser beam
is split into three laser beams of equal power. These laser
beams are led by means of optic fibres to exit the optics
of the laser. This optics is equidistantly positioned
around the wire in the plane perpendicular to the wire.
The focal length of the exit lens is f = 100 mm and the
diameter of the laser beam in focus is df = 0.4 mm. For
an easier positioning of the wire and monitoring of the
process, the laser optics is equipped with the CCD came-
ras connected to the screens. Laser optics can be pre-
cisely positioned in accordance with the wire using a
system of micrometer positioning stages.
4.2 Pendant-droplet formation
The laser droplet-formation process was separated
into two phases. The first phase is the formation of a
pendant droplet at the tip of the wire and the second
phase is the detachment of the droplet from the wire.
Therefore, we divided the total laser-pulse duration of 20
ms into a 12 ms part for the formation of the pendant
droplet and the remaining 8 ms of the laser pulse for the
droplet detachment.
First, the laser pulses for pendant-droplet formation
were theoretically determined as described in section
3.1. Examples of the laser pulses are shown in Figure 3.
T. KOKALJ et al.: MODELLING AND OPTIMIZATION OF A LASER DROPLET-FORMATION PROCESS
Materiali in tehnologije / Materials and technology 47 (2013) 4, 487–495 491
Figure 3: Examples of laser pulses (black line): a) laser pulse deter-
mined with the fitness function J1 and laser pulses determined with the
fitness function J2 with the selected final temperatures: b) Tz1 = 2000
K and c) Tz2 = 3000 K. Grey line denotes the velocity profile of the
wire movement.
Slika 3: Primeri laserskih bliskov (~rna ~rta): a) laserski blisk,
dolo~en s kriterijsko funkcijo J1 in laserska bliska, dolo~ena s krite-
rijsko funkcijo J2 in razli~nimi ciljanimi kon~nimi temperaturami: b)
Tz1 = 2000 K in c) Tz2 = 3000 K. Siva ~rta prikazuje hitrostni profil
podajanja `ice.
Figure 4: Time courses of the shares of the material in solid, liquid
and vapour phases during the laser pulses. Graphs correspond to the
laser pulses shown in Figure 3.
Slika 4: ^asovni potek dele`a materiala v trdnem, teko~em in pli-
nastem stanju med trajanjem laserskega bliska. Slike po vrsti ustrezajo
laserskim bliskom s slike 3.
The first pulse is determined with the fitness function
J1 that was used to assure the best trade-off between the
shares of the melted and vaporized material. With the
second J2 fitness function we aimed at bringing the
uniform temperature of the droplet as near to the desired
temperature as possible. Besides, we wanted to decrease
the vaporization of the material during the pendant-
droplet formation, which was achieved with the second
term in the fitness function. The first selected tempe-
rature is just above the melting temperature of nickel and
the second is near the boiling temperature of nickel. The
above-mentioned figure shows how the pulses differ, not
just in the energy, but also in the time course of the laser
power. At the beginning they differ a lot, but at the end
they all have similar shapes. The laser power increases in
the interval from 8 ms to 10 ms because the wire is
moving with the largest velocity. At the end, when the
velocity is decreased, the laser power is correspondingly
reduced.
From the theoretical model we can also predict the
share of the material in each phase during the laser-pulse
duration, as shown in Figure 4.
As can be seen from the above figure, at the beginn-
ing, all material is in the solid phase. After a certain
period the melting sets in. By the end of the pulse most
of the material is melted and some of the material is
vaporized. In the model vaporization is treated as the
overheated liquid material that turns into a liquid if it
cools down.
The evolution of the temperature field of a cross-sec-
tion of the wire is shown in Figure 5.
The horizontal line denotes the position of laser
beams and the black curves divide the material in diffe-
rent phases. The temperature fields can be compared to
the experimental results. Three typical final states of the
pendant droplet after solidification are shown in Figure
6.
The dashed cross in the image denotes the final posi-
tion of the laser beam and the non-dashed cross was used
for an easier positioning of the wire.
The first temperature field of the wire predicts the
melted wire, but, as we can see in the photo, the tip of
the wire is not completely melted. There is a knob on the
wire tip that corresponds to the area of the lowest tem-
perature predicted by our model. The core of the wire
was probably not completely melted at the wire tip and
the consequence is an irregular shape of the droplets.
As shown in the second figure of the temperature
field, the predicted temperature of the wire tip is just
above the melting temperature of nickel (1727 K). In the
image of the experimental outcome we can see that no
droplet was formed at the tip of the wire. The only
noticeable effect of laser heating is a deformation of the
wire at the position where the highest wire temperature
is predicted by the model. This deformation is denoted
with a white arrow in the image of the experimental out-
come. The reason for a disagreement of the model pre-
diction with the experimental results is the presumption
on the uniform heating of the wire circumference, due to
the simplification of the model.
T. KOKALJ et al.: MODELLING AND OPTIMIZATION OF A LASER DROPLET-FORMATION PROCESS
492 Materiali in tehnologije / Materials and technology 47 (2013) 4, 487–495
Figure 6: Typical examples of experimental outcomes of pendant-
droplet formation. Droplets correspond to the laser pulses shown in
Figure 3 denoted with the same letters.
Slika 6: Tipi~ni primeri eksperimentalnih tvorjenj vise~e kapljice.
Kapljice ustrezajo laserskim bliskom s slike 3 in so ozna~ene z ustrez-
nimi ~rkami.
Figure 5: Evolution of the temperature field of the wire. Images
correspond to the laser pulses shown in Figure 3.
Slika 5: Razvoj temperaturnega polja `ice. Slike ustrezajo laserskim
bliskom s slike 3.
In the third case the model predicts a very hot tip of
the wire near the boiling temperature. The experimental
outcomes yield a repeatable formation of the melted
droplets at the tip of the wire.
4.3 Droplet detachment
After the formation of a pendant droplet, we have to
detach it from the wire. An additional high-intensity
laser pulse is needed to assure a droplet detachment. This
additional laser pulse is determined as described in
section 3.2. The laser power is selected in advance and
the duration of a laser pulse is determined in line with
the requirement that the boiling temperature reaches the
axis of the wire.
The best laser pulse for pendant-droplet formation
was selected and a short time of 2 ms for the minimum
laser power of Pmin = 480 W was added to that pulse.
This pause was heuristically introduced to allow a drop-
let relaxation, since the droplet oscillates after it is
formed at the tip of the wire. Three examples of the
calculated laser pulses are shown in Figure 7.
The required time for droplet detachment decreases
with an increased laser power as expected. The depen-
dence of the necessary pulse duration on the applied
laser power is shown in Figure 8.
The time development of the temperature field of the
wire during a droplet-detachment pulse is shown in
Figure 9.
The detachment pulse is stopped when the boiling
temperature reaches the axis of the wire. The horizontal
line denotes the position of the laser beams. The results
of the model show that more material is melted when a
higher laser power is used in spite of a shorter pulse
time.
After the theoretical determination of detachment
pulses, the experiments were performed for the eva-
luation of theoretical results. A laser pulse for a droplet
detachment cannot be tested alone. Therefore, different
detachment pulses were added after a relaxation period 2
ms that followed the same pendant-droplet-formation
pulse.
With each laser pulse 10 droplets were produced and
deposited on the substrate placed horizontally 3.5 mm
below the focus of the laser beams. Examples of droplets
are shown in Figure 10.
Radial scatter and a number of splashes on the
substrate depend on the applied laser pulse. Droplets are
systematically shifted to the left. The images of typical
droplets in Figure 10 show that the droplets differ also in
the roughness of the contact region and the shape of a
droplet on the substrate.
T. KOKALJ et al.: MODELLING AND OPTIMIZATION OF A LASER DROPLET-FORMATION PROCESS
Materiali in tehnologije / Materials and technology 47 (2013) 4, 487–495 493
Figure 9: Time development of the temperature field during a
droplet-detachment pulse
Slika 9: ^asovni razvoj temperaturnega polja kapljice med bliskom za
lo~itev kapljice
Figure 7: Three examples of laser pulses with the same pendant-drop-
let-formation parts and different detachment parts. Droplet-detach-
ment pulses are of: a) 2000 W, b) 4000 W and c) 6000 W.
Slika 7: Trije primeri laserskih bliskov z enakim za~etnim delom za
tvorjenje vise~e kapljice in razli~nimi bliski za lo~itev kapljice. Bliski
za lo~itev kapljice imajo mo~i: a) 2000 W, b) 4000 W in c) 6000 W.
Figure 8: Dependence of detachment-pulse duration on the selected
laser power
Slika 8: Odvisnost trajanja bliska za lo~itev kapljice od izbrane mo~i
laserja
For the purpose of evaluation and comparison a more
detailed quantitative analysis of the experimental results
was performed. The basic criterion for the successfulness
of the process is the detachment of a droplet. Undesired
splashes on the substrate were also counted. The repeat-
ability of droplet deposition and droplet size was charac-
terized with the size and standard deviation of these
parameters. The results of the analysis are presented in
Table 1.
We introduced variability of the process, V:
V
d d
d r= +
 
(10)
where the standard deviation of the droplet diameter is
  d dx dy= +
2 2 , the standard deviation of the droplet
position from the expected position is   r rx ry= +
2 2
and d is the average droplet diameter for a set. The ratio
between the number of splashes on the substrate Nizb
and the number of detached droplets in a set Nkap was
calculated.
At the highest power of a detachment pulse, a droplet
was always detached, but accompanied by numerous
splashes. At a smaller laser power, droplets sometimes
stay undetached. At a laser power of 2 kW droplets
always stayed attached to the wire. At a laser power of 4
kW two droplets stayed undetached and at 5 kW one
droplet stayed undetached. At a laser power of 6 kW all
the droplets detached from the wire. Also, the variability
of the process decreases with the laser power. The
process variabilities at the laser powers of 4 kW, 5 kW
and 6 kW are very similar.
5 DISCUSSION AND CONCLUSION
A laser droplet-formation process is discussed in the
article. For the first time a more detailed physical and
numerical model of the process is presented. After
building an experimental system and conducting quite an
extensive experimental work we find out that the process
is very complex in a sense of a high number of
influencing parameters and instable in a sense of the
sensitivity to small parameter change. To get a deeper
insight into the process we decided to build a more
detailed theoretical and numerical model of the process.
Instead of using the model just for the analysis of the
process we decided to determine certain selected process
parameters on the basis of a numerical optimization of
the process. In order to model the keyhole growth we
proposed a very simple model that adequately serves our
purpose.
The results of experimental work show that some
theoretically determined laser pulses yield very good
results. In some cases we notice a weaker agreement
between the theoretical and experimental results, which
is the consequence of a relatively simplified theoretical
model. However, based on the numerical simulation of
the process the experimental results can still be easily
understood and interpreted. The theoretical model and
optimization procedure are not meant to replace the
experimental work completely, but can significantly
decrease the extent of the required work. In this manner
we should understand the importance of the presented
work as being complementary to the experimental work.
It allows a faster search for proper parameters and an
easier interpretation of experimental results, especially
when an important part of the process is changed like the
wire material, laser optics, wire-feed system, etc.
Though we give a quantitative statistical description of
experimental results for the purpose of comparing
experimental outcomes, the suitability of the process for
a certain application should be determined by the final
user in accordance with the specific demands of the
application.
T. KOKALJ et al.: MODELLING AND OPTIMIZATION OF A LASER DROPLET-FORMATION PROCESS
494 Materiali in tehnologije / Materials and technology 47 (2013) 4, 487–495
Figure 10: Magnifications of typical droplets from the sets produced
with different detachment pulses
Slika 10: Pove~ava tipi~nih kapljic iz zaporedij, tvorjenih z razli~nimi
bliski za lo~itev kapljice
Table 1: Quantitative analysis of the experimental outcome for different detachment pulses
Tabela 1: Kvantitativna analiza eksperimentalnih rezultatov za razli~ne lo~itvene bliske
P d/mm d/d rx/mm ry/mm r /mm r/d v Nizb/Nkap
8 kW 1.237 0.155 1.024 1.585 1.895 0.355 0.510 100/10 = 10
6 kW 1.223 0.169 –0.008 1.199 1.213 0.155 0.324 48/10 = 4.8
4 kW 1.149 0.098 0.679 1.276 1.458 0.255 0.323 29/8 = 3.6
5 kW 1.111 0.163 0.332 0.843 0.946 0.181 0.344 45/9 = 5
6 REFERENCES
1 D. Attinger, D. Poulikakos, Melting and resolidification of a substra-
te caused by molten microdroplet impact, Journal of Heat transfer,
123 (2001) 12, 1110–1122
2 A. M. Ahmed, R. H. Rangel, Metal droplet deposition on non-flat
surfaces: effect of surface morphology, International Journal of Heat
and Mass Transfer, 45 (2002), 1077–1091
3 S. Cheng, S. Chandra, A pneumatic droplet-on-demand generator,
Experiments in fluids, 34 (2003), 755–762
4 J. Brünahl, A. M. Grishin, Piezoelectric shear mode drop-on-demand
inkjet actuator, Sensors and Actuators A, 101 (2002), 371–382
5 H. Sohn, D. Y. Yang, Drop on demand deposition of superheated
metal droplets for selective infiltration manufacturing, Materials
Science and Engineering A, 392 (2005) 1–2, 415–421
6 S. X. Cheng, T. Li, S. Chandra, Producing metal droplets with a
pneumatic droplet-on-demand generator, Journal of Materials Pro-
cessing Technology, 195 (2005), 295–302
7 J. Haidar, An analysis of the formation of metal droplets in arc weld-
ing, Journal of Physics D: Applied physics, 31 (1998), 1233–1244
8 E. L. Dreizin, Droplet welding: A new technology for welding elec-
trical contacts, Welding Journal, 76 (1997), 67–73
9 J. J. C. Buelens, Metal droplet deposition: review of innovative join-
ing technique, Science and Technology of Welding and Joining, 6
(1997) 2, 239–243
10 F. Wang, W. K. Hou, S. J. Hu, E. Kannatey-Asibu, W. W. Schultz, P.
C. Wang, Modelling and analysis of metal transfer in gas metal arc
welding, Journal of Physics D: Applied physics, 36 (2003),
1143–1152
11 W. W. Appelman, Miniaturisation of droplet deposition, Technical
report, Philips Centre for Manufacturing Technology, Eindhoven
1995
12 Y. Wu, R. Kovacevic, Mechanically assisted droplet transfer process
in gas metal arc welding, Proceedings of the Institution of Mecha-
nical Engineers, Part B, Journal of engineering manufacture, 216
(2002), 555–564
13 P. P. Conway, E. K. Y. Fu, K. Williams, Precision high temperature
lead-free solder interconnections by means of high-energy droplet
deposition techniques, CIRP Annals, 51 (2002) 1, 177
14 E. Govekar, B. Jahrsdoerfer, J. Klemen~i~, P. Mu`i~, I. Grabec,
Characterization of laser droplet formation process by IR camera
images and AE signals, Dynamic Days, Dresden, 2001
15 W. Hovig, B. Jahrsdorfer, Laser droplet weld - ein innovatives
fügeverfahren, Laser in Elektronikproduktion & Feinwerktechnik,
LEF 2001, 2001
16 J. Klemen~i~, Investigation of laser droplet formation from thin wire,
Master thesis, FS, UL, Ljubljana, 2003
17 B. Jahrsdorfer, G. Eßer, M. Geiger, E. Govekar, Laser droplet weld –
an innovative joining technology opens new application possibilities,
Photon Processing in Microelectronics and Photonics II, Proceedings
of SPIE, 2003, 4977
18 E. Govekar, J. Klemen~i~, T. Kokalj, B. Jahrsdorfer, P. Mu`i~, I. Gra-
bec, Characterization of laser droplet formation process by acoustic
emission, Ultrasonics, 42 (2004), 99–103
19 B. Jahrsdorfer, M. Schmidt, M. Geiger, Laser droplet welding and its
potential for joining dissimilar materials, Laser Assisted Net Shape
Engineering 4, Proceedings of LANE 2004, 2004
20 B. Krese, M. Perc, E. Govekar, The dynamics of laser droplet gene-
ration, Chaos, 20 (2010), 013129
21 A. Jeri~, I. Grabec, E. Govekar, Laser droplet welding of zinc coated
steel sheets, Science and Technology of Welding and Joining, 14
(2009) 4, 362–368
22 T. Kokalj, J. Klemen~i~, P. Mu`i~, I. Grabec, E. Govekar, Analysis of
the laser droplet formation process, Journal of Manufacturing
Science and Engineering, 128 (2006) 1, 307–314
23 J. W. Thomas, Numerical Partial Differential Equations: Finite
Difference Methods, Springer-Verlag, New York 1995
24 D. Whitley, A genetic algorithm tutorial, Statistics and Computing, 4
(1994), 65–85
25 K. F. Man, K. S. Tang, S. Kwong, Genetic algorithms: Concepts and
applications, IEEE Transactions on Industrial Electronics, 5 (1996)
43, 519–534
26 A. E. Eiben, J. E. Smith, Introduction to Evolutionary Computing,
Springer, Berlin 2003
T. KOKALJ et al.: MODELLING AND OPTIMIZATION OF A LASER DROPLET-FORMATION PROCESS
Materiali in tehnologije / Materials and technology 47 (2013) 4, 487–495 495

F. KAVICKA et al.: COMPARISON OF THE TEMPERATURE FIELDS OF CONTINUOUSLY CAST STEEL SLABS ...
COMPARISON OF THE TEMPERATURE FIELDS OF
CONTINUOUSLY CAST STEEL SLABS WITH
DIFFERENT CHEMICAL COMPOSITIONS
PRIMERJAVA TEMPERATURNEGA POLJA KONTINUIRNO ULITIH
JEKLENIH SLABOV RAZLI^NIH KEMIJSKIH SESTAV
Frantisek Kavicka1, Karel Stransky1, Bohumil Sekanina1, Josef Stetina1,
Milos Masarik2, Tomas Mauder1, Vasilij Gontarev3
1Brno University of Technology, Technicka 2, 616 69 Brno, Czech Republic
2Evraz Vitkovice Steel, a.s. Ostrava, Czech Republic
3University of Ljubljana, A{ker~eva 12, 1000 Ljubljana, Slovenia
kavicka@fme.vutbr.cz
Prejem rokopisa – received: 2012-08-21; sprejem za objavo – accepted for publication: 2012-12-19
The numerical model made by the authors was used for simulating the transient temperature fields of the continuously cast steel
slabs with two different chemical compositions. The model solves the Fourier-Kirchhoff equation of the temperature fields of
the slab-crystallizer system and the slab-ambient system with the following main thermophysical parameters: thermal
conductivity, specific-heat capacity, density and enthalpy. When both melts follow each other closely, the critical state of the
so-called breakout occurs at a certain point in the secondary cooling zone of a caster. It is probably a combination of surface
defects. However, different chemical compositions of the two steels and their mixture are apparently decisive. Therefore, the
temperature model simulated the temperature history of every point of a cross-section of a slab during its movement through the
caster from the level of the melt in the crystallizer to the cutting torch for both melts and their mixture. The calculation of the
temperature field of a slab is mainly focused on the part of the slab before the breakout and its surroundings. The results for the
temperature field are used to set up a model of chemical heterogeneity of the steel supported by the material investigation of the
samples taken from the breakout.
Keywords: continuously cast slab, temperature field, chemical composition, heterogeneity, numerical model, breakout
Uporabljen je bil originalen numeri~ni model za simulacijo prehodnega temperaturnega polja kontinuirno ulitih jeklenih slabov
dveh razli~nih kemijskih sestav. Model re{uje Fourier-Kirchhoffovo ena~bo za sistema slab-kristalizator oz. slab-okolje
z naslednjimi glavnimi termo-fizikalnimi parametri: toplotna prevodnost, specifi~na toplotna kapaciteta, gostota in entalpija. Ko
se talina v formi zme{a s prej{njo, pride do kriti~nega stanja, t. i. prodora v dolo~eni to~ki livne sekundarne hladilne cone. To
nastane verjetno zaradi kombinacije povr{inskih napak. Tako je razli~na kemijska sestava dveh jekel in njihovo me{anje o~itno
odlo~ilno pri tem pojavu. Temperaturni model je simuliral temperaturni potek v vsaki to~ki pre~nega prereza slaba med
njegovim gibanjem skozi celotni livni stroj od nivoja taline v kristalizatorju do rezalnega plamena za obe talini in njune
me{anice. Izra~un temperaturnega polja slaba je bil osredinjen predvsem na del slaba pred prodorom in njegovo okolico. Model
kemijske heterogenosti jekla se lahko uveljavi z rezultati temperaturnega polja in z raziskavami vzorcev, ki so bili odvzeti iz
obmo~ja prodora.
Klju~ne besede: kontinuirno ulit slab, temperaturno polje, kemijska sestava, heterogenost, numeri~ni model, prodor
1 INTRODUCTION
The authors continue the study of the causes for the
breakouts of continuously cast steel slabs of 250 mm ×
1530 mm covered in the previous publications.1–3 This is
a defect that cannot be repaired. The break was detected
in the unbending point of a slab, at a distance of 14.15 m
away from the level of the melt inside the mould, where
a breakout occurred between the 7th and 8th cooling
segments of the secondary cooling zone. The difference
in the heights between the level inside the mould and the
breakout point was 8605 m. This tear in the shell
occurred on the small radius of the caster. Here, the slab
is beginning to straighten out and the breakout of the
steel can occur in the points of the increased local
chemical and temperature heterogeneities of the steel.
The changes in the chemical composition of the steel
during the actual continuous casting are especially
dangerous. This change in the chemical composition of
the steel of two qualities, A and B, was carried out very
quickly, by changing the tundish. Inside the mould,
steel B mixed with steel A of the previous melt. There-
fore, the results of the calculation of the temperature
fields of the slabs with the chemical compositions A, B
and of the mixed composition A + B will be presented.
The mixed content of each element is considered as the
average value of the composition A and B.
2 3D NUMERICAL MODEL OF THE
TEMPERATURE FIELD OF A STEEL SLAB AND
PREPARATION FOR THE SIMULATION
The optimisation of the production on casters, with
the aim of achieving the maximum savings and maxi-
mum quality of the product is unthinkable without the
knowledge of the course of the solidification and cooling
of the concasting. The solidification and cooling of a
Materiali in tehnologije / Materials and technology 47 (2013) 4, 497–501 497
UDK 519.61/.64:621.74.047 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(4)497(2013)
concast slab is a global problem of the 3D transient heat
and mass transfer. If the heat conduction within the heat
transfer in this system is decisive, the process is de-
scribed with the Fourier-Kirchhoff equation. It describes
the temperature field of the solidifying slab in all three of
its states: at the temperatures above the liquidus (i.e., the
melt), within the interval between the liquidus and
solidus (i.e., in the mushy zone) and at the temperatures
below the solidus (i.e., the solid state). In order to obtain
these results, it is convenient to use the explicit nume-
rical method of finite differences. A numerical simula-
tion of the release of latent heats of the phase or structu-
ral changes is carried out by introducing the enthalpy
function dependent on temperature T. The latent heats
are expressed after an automated generation of the net-
work (pre-processing) ties on the entry of the thermo-
physical material properties of the investigated system.
They are the heat conductivity k, the specific heat
capacity c and the density 
 of the cast metal.
The temperature distribution in the slabs described
with the enthalpy balance equation is as follows:
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
( )
( ) ( ) ( )



 
 

H
t x
uH
y
vH
z
wH
x
k
T
x
+ + + =
= ⎛⎝
⎜ ⎞
⎠
⎟ +
⎛
⎝
⎜
⎞
⎠
⎟ + ⎛⎝
⎜ ⎞
⎠
⎟∂
∂
∂
∂
∂
∂
∂
∂y
k
T
y z
k
T
z
(1)
The simplified equation (1), suitable for an applica-
tion on the radial casters with a great radius, where only
the speed (of the movement of the slab) component w in
the z-direction is considered, is:
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
∂
( )
( )




H
t z
wH
x
k
T
x y
k
T
y
+ =
= ⎛⎝
⎜ ⎞
⎠
⎟ +
⎛
⎝
⎜
⎞
⎠
⎟ +
z
k
T
z
∂
∂
⎛
⎝
⎜ ⎞
⎠
⎟
(2)
The unknown enthalpy of the general nodal point of a
slab in the next time step (t  t): H i j k
t t
, ,
( )+ Δ is expressed
with the explicit formula and it is a function of the
enthalpies of the same node and six adjacent nodes in the
Cartesian coordinate system from the previous time step
t.
The thermophysical properties of the cast materials,
k, c and 
, contained in the equation and dependent on
the chemical compositions are also functions of the tem-
perature.4 The function of enthalpy H is not known as an
analytical function, but as a set of tabular values, which
means that a reverse determination of the temperature is
numerically a highly demanding task. It is also depen-
dent on the composition of the steel and on the rate of
cooling.
Figure 1 shows that the task is symmetrical along the
x-axis; it is, therefore, sufficient to investigate only one
half of the cross-section. The 3D model was first de-
signed as an off-line version and later as an on-line
version so that it could work in real time. After the
correction and testing, it will be possible to implement it
on any caster thanks to the universal nature of the code.
The numerical model takes into account the temperature
field of the entire slab (from the meniscus of the level of
the melt in the mould to the cutting torch) with the
number of the nodes exceeding 106 on the half of the
cross-section of the rectangular profile. The solution is
performed under these boundary conditions:
T T= cast at the meniscus (3a)
− =k
T
n
∂
∂
0 at the plane of symmetry (3b)
− = −k
T
n
h T T
∂
∂
( )surf mould in the mould (3c)
− = − + −k
T
n
h T T se T T
∂
∂
( ) ( )surf amb surf amb
4 4 (3d)
in the secondary and tertiary cooling zones
F. KAVICKA et al.: COMPARISON OF THE TEMPERATURE FIELDS OF CONTINUOUSLY CAST STEEL SLABS ...
498 Materiali in tehnologije / Materials and technology 47 (2013) 4, 497–501
Figure 1: Caster and definition of the coordinate system
Slika 1: Livni stroj in definicija koordinatnega sistema
− =k
T
n
q
∂
∂
beneath the support rollers (3e)
The initial condition for obtaining a solution is the
setting of the initial temperature in individual points of
the network. The suitable value is the highest possible
temperature, i.e., the casting temperature Tcast.
It is, therefore, a solution of a distinctly non-linear
task, since even for the boundary conditions their depen-
dence on the surface temperature of a continuously cast
slab is respected. All the boundary conditions were
derived and were then divided to the areas of the mold
(the primary cooling) and of the secondary and tertiary
cooling. The zero heat flow perpendicular to the symme-
try plane is the boundary condition that is common to all
the zones of the caster.5
In equations (1–3) T is the temperature in K, t is the
time in s, k is the heat conductivity in W m–1 K–1, 
 is the
density in kg m–3, x, y, z are the axes in the given
directions in m, u, v, w are the velocities in the given
directions (the shift rate) in m s–1, H is the specific
enthalpy in J kg–1, Tsurf, Tamb are the surface and ambient
temperatures in K, Tcast, Tmould are the casting and the
mould temperatures in K,  is the Stefan-Boltzmann
constant (5.76 × 10–6 W m–2 K–4),  is the emissivity
(0.8), n is normal to the surface in m, h in W m–2 K–1 is
the heat-transfer coefficient (HTC), q in W m–2 is the
specific heat flow.
The heat-transfer coefficient h is a function of the
local cooling rate and the surface temperature, the
temperature Tamb is the cooling-water temperature in the
secondary-cooling zone and the air temperature, where
only radiation occurs. Based on the results of the
previous investigations, the boundary conditions are set
identically within each zone individually. The definition
of the boundary conditions under the cooling jet is espe-
cially difficult. In order for the model to function
correctly, it is necessary to determine the correct value of
the heat-transfer coefficient h underneath the jet. There-
fore, extensive experiments were conducted on a labora-
tory device simulating the process of solidification
within the secondary-cooling zone. These measurements
were conducted for various operation conditions, i.e., the
pressure of the water and the shift rate. The numerical
model contains a function that calculates the current on
the basis of the entered water pressure, the shift rate, the
positions of the jet and the surface temperature.
The casting speed was 0.0130 m s–1 (steel A) and
0.0126 m s–1 (steel B), the superheat was 23 °C (steel A)
and 18 °C (steel B), the solidus temperature was 1427.0
°C (steel A) and 1480.6 °C (steel B), the liquidus tempe-
rature was 1493.9 °C (steel A) and 1512.3 °C (steel B).
The chemical composition of steel A in mass frac-
tions (%) was: w(C) = 0.416, w(Cr) = 0.95, w(Ni) = 0.03,
w(Mn) = 0.7, w(Mo) = 0.206, w(Si) = 0.28.
The chemical composition of steel B in mass frac-
tions (%) was: w(C) = 0.174, w(Cr) = 0.07, w(Ni) = 0.02,
w(Mn) = 1.46, w(Mo) = 0.005, w(Si) = 0.23.
The thermophysical properties of the steels were cal-
culated using the IDS solidification analysis package of a
prestigious laboratory (the Laboratory of Metallurgy,
Helsinki University of Technology). The IDS package
calculates enthalpy, specific heat, density, thermal con-
ductivity, thermal contraction, etc., from the liquid state
down to the room temperature using graphic or numeric
outputs of the results. The dependences of the main
F. KAVICKA et al.: COMPARISON OF THE TEMPERATURE FIELDS OF CONTINUOUSLY CAST STEEL SLABS ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 497–501 499
Figure 3: Density of steels A and B and its dependence on tempera-
ture
Slika 3: Gostota jekel A in B v odvisnosti od temperature
Figure 2: Heat conductivity of steels A and B and its dependence on
temperature
Slika 2: Toplotna prevodnost jekel A in B v odvisnosti od temperature
Figure 4: Enthalpy of steels A and B and its dependence on tempe-
rature
Slika 4: Entalpija jekel A in B v odvisnosti od temperature
thermophysical properties on the temperature – the heat
conductivity, the density and the enthalpy of both steels
are shown in Figures 2 to 4. The definitions of the boun-
dary conditions for all three variants of the calculation
were identical for all the cooling zones.
3 NUMERICAL RESULTS
The off-line version of the original temperature
model was now used to simulate the temperature fields
of the steel slab of A quality, the steel slab of B quality
and the steel slab of A + B quality (with the average che-
mical compositions). After the computation, it is
possible to obtain the temperatures for each node of the
network and for any time during the process. The course
of the calculated isotherms in the cross-section of the
breakout, for example, for a slab of B quality at a
distance of 14.15 m from the level of the melt in the
mold is in Figure 5.
The courses of the calculated isoliquidi and isosolidi
as the characteristic isotherms in one half of the cross-
section of the breakout for the steel slabs of A, B and A
+ B qualities are in Figure 6. The isoliquidi and isosolidi
in both longitudinal axial sections of the steel slabs of A,
B and A + B qualities are plotted in Figures 7 and 8.
4 CONCLUSIONS
In the secondary-cooling zone, where the slab is
beginning to straighten out, the breakout of the steel can
occur in the points of increased local chemical and tem-
perature heterogeneities of the steel, from the increased
F. KAVICKA et al.: COMPARISON OF THE TEMPERATURE FIELDS OF CONTINUOUSLY CAST STEEL SLABS ...
500 Materiali in tehnologije / Materials and technology 47 (2013) 4, 497–501
Figure 7: Isoliquidi and isosolidi in the horizontal section for the steel
slabs of A, B and A + B qualities
Slika 7: Izolikvidusne in izosolidusne ~rte v horizontalnem prerezu za
jekli A in B ter za A + B
Figure 6: Isoliquidi and isosolidi in the cross-section of the breakout
for the steel slabs of A, B and A + B qualities
Slika 6: Izolikvidusne in izosolidusne ~rte v pre~nem prerezu prodora
za jekli A in B ter za A + B
Figure 5: Isotherms in the cross-section of the breakout for the steel
slab of B quality
Slika 5: Izoterme v pre~nem prerezu prodora jeklenega slaba B
tension as a result of the bending of the slab and also a
high local concentration of non-metal, slag inclusions.
The changes in the chemical composition of the steel
during the fill-up of the tundish with continuous casting
are especially dangerous. The consequence of this opera-
tion, an immediate change in the chemical composition
of the steel, which is not prevented by a breakout system
directly inside the mould, can lead to an immediate inter-
ruption of the concasting and a breakout at a greater
distance from the mould than usual, thus leading to a
significant material loss and downtime.
During the continuous casting of the slab of 250 mm
× 1530 mm and after a quick change of the tundish, a
change in the chemical composition of steel A and B
occurred. Inside the mould, steel B thus mixed with
steel A of the previous melt. After 20 min of casting
steel B with a different chemical composition, the caster
stopped as a result of a breakout at a distance of 14.15 m
from the level of the melt in the mold. Therefore, this
critical state of continuous casting was analyzed via the
temperature model that may be followed by an analysis
of the chemical heterogeneity in the plain of the break-
out. The results of the formation of the temperature
fields of the steel slabs of quality A, B, and A + B (with
the average chemical compositions) were obtained. The
objective of the paper is to analyse the temperature fields
in the plain of the breakout, which may be followed by
an analysis of chemical heterogeneities in this plain. In
the case of a breakout we do not have more detailed
information about the internal relationship between the
dimensional quantities and the essence of the breakout,
and we do not even have a partial mathematical and
physical description of this phenomenon. That is why it
is necessary to use the theory of similarity for a dimen-
sional analysis determining the dimensionless criteria.6
Based on the -theorem, 5 similarity criteria were
derived containing 12 technological, geometrical and
thermo-physical dimensional quantities that characterise
both steel grades A and B and also the process of their
continuous casting. This analysis also provides the cal-
culations of the temperature fields of the slabs to be cast,
in order to determine the quantities needed for the fulfil-
ment of some of these criteria. The analysis performed
by using the similarity criteria clearly demonstrates a
significantly increased tendency for steel B to breakout,
in comparison with steel A.
Acknowledgment
This analysis was conducted using a program devised
within the framework of the GA CR project No.
P107/11/1566.
5 REFERENCES
1 F. Kavicka, K. Stransky, B. Sekanina, J. Stetina, T. Mauder, The
calculation of the temperature field of a concast steel slab tightly
before the breakout, Metal 2011 Conference Proceedings Papers
Symp. A, Metal. Ostrava, Tanger, (2011), 93–98
2 F. Kavicka, K. Stransky, J. Dobrovska, B. Sekanina, A breakout of a
slab after quick changing the tundish and steel quality, Metal 2011
Conference Proceedings Papers Symp. A, Metal. Ostrava, Tanger,
(2011), 99–104
3 C. Ojeda et al., Mathematical Modeling of Thermal-Fluid Flow in
the Meniscus Region During an Oscillation Cycle, AISTech 2006
Proceedings, 1 (2006), 1017–1028
4 S. Louhenkilpi, E. Laitinen, R. Nieminen, Real-Time Simulation of
Heat Transfer in Continuous Casting, Metallurgical Transactions B,
24B (1993) 4, 685–693
5 J. Stetina, A dynamic model of temperature field of concast steel
slab, PhD Thesis, VSB-TU Ostrava, Faculty of Metallurgy and
Materials Engineering, Ostrava, 2007
6 K. Stransky, J. Dobrovska, F. Kavicka, J. Stetina, B. Sekanina, M.
Masarik, Heterogeneity of continuously cast steel slab shortly before
breakout, Metal 2012, Proceedings of Abstracts and CD ROM of the
20th International Metallurgical & Materials Conference Brno,
Czech Republic, (2012)
F. KAVICKA et al.: COMPARISON OF THE TEMPERATURE FIELDS OF CONTINUOUSLY CAST STEEL SLABS ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 497–501 501
Figure 8: Isoliquidi and isosolidi in the vertical-axis section for the
steel slabs of A, B and A + B qualities
Slika 8: Izolikvidusne in izosolidusne ~rte v vertikalnem osnem pre-
rezu za jekli A in B ter za A + B

P. LICHÝ et al.: INVESTIGATION OF THE THERMOMECHANICAL PROPERTIES AND MICROSTRUCTURE ...
INVESTIGATION OF THE THERMOMECHANICAL
PROPERTIES AND MICROSTRUCTURE OF SPECIAL
MAGNESIUM ALLOYS
PREISKAVA TERMOMEHANSKIH LASTNOSTI IN
MIKROSTRUKTURE POSEBNIH MAGNEZIJEVIH ZLITIN
Petr Lichý, Michal Cagala, Jaroslav Beòo
V[B-Technical University of Ostrava, Faculty of Metallurgy and Materials Engineering, Department of Metallurgy and Foundry Engineering,
17. listopadu 15/2172, 708 33 Ostrava, Czech Republic
petr.lichy@vsb.cz
prejem rokopisa – received: 2012-08-31; sprejem za objavo – accepted for publication: 2013-01-03
Thanks to their high specific strength together with a low density, magnesium-based alloys have an extensive potential for the
use of their castings in automotive industry. The castings based on these materials show relatively good mechanical properties
that decrease significantly with an increasing temperature of their thermal exposure. The aim of this work is to study the
properties of two alloys commonly used (Mg-Al-Zn, Mn type). From the studied materials the test castings were made and these
were formed into the test bars used for a tensile test. This test was carried out within the temperature range of 20–300 °C.
Further on, the tests of the thermomechanical properties were complemented with a microstructure analysis with the aim of
checking metallurgical interventions (the effect of inoculation).
Keywords: castings, magnesium alloys, thermomechanical properties, microstructure
Zaradi velike specifi~ne trdnosti in majhne gostote imajo magnezijeve zlitine velik potencial za uporabo v obliki ulitkov v
avtomobilski industriji. Ulitki iz teh materialov ka`ejo relativno dobre mehanske lastnosti, ki pa se mo~no zmanj{ajo, ~e so
izpostavljeni povi{anim temperaturam. Namen tega dela je {tudij teh lastnosti pri dveh pogosto uporabljanih vrstah zlitine
(Mg-Al-Zn, Mn). Iz preiskovanih materialov so bili izdelani preizkusni ulitki, iz katerih so bile pripravljene palice za natezne
preizkuse. Ti so se izvajali v temperaturnem obmo~ju 20–300 °C. Ugotovitve termomehanskih lastnosti so bile dopolnjene z
analizo mikrostrukture, namenjene za preverjanje metalur{kih ukrepov (u~inek inokulacije).
Klju~ne besede: ulitki, magnezijeve zlitine, termomehanske lastnosti, mikrostruktura
1 INTRODUCTION
Magnesium alloys have always been very prospective
materials, in particular for automobile and aircraft
industries where the most important requirement is for a
low weight or sufficiently high specific strength, i.e., the
strength-characteristics-to-low-specific-weight ratio. In
spite of the above mentioned reasons, a more extensive
use of these alloys in the above industrial branches is
limited by their low resistance to corrosion and a
decrease in their mechanical properties under higher
temperatures. Standard alloys are not stable under high
temperatures and therefore not suitable for the applica-
tions, in which they would be stressed by such tempera-
tures. In addition, a more advanced casting technology
must be taken into account as magnesium and its alloys
are highly reactive metals. But these difficulties can
already be partly eliminated by choosing a suitable
preparation in a liquid metal, moulding mixture or a
protective atmosphere in a melting device.
The use of magnesium alloys in a car structure can
considerably help in reducing the car weight without
deteriorating its safety.1
2 DIVISION OF MAGNESIUM ALLOYS
ACCORDING THEIR USES
The basis for the foundry magnesium alloys are
binary alloys enhanced with additional alloying elements
in order to improve technological properties, mechanical
properties or to increase the corrosion resistance. There
are the basic systems – Mg-Al, Mg-Zn and Mg-Mn. The
most widespread magnesium alloys for manufacturing
castings are the Mg-Al-(Zn, Mn)-type alloys. The advan-
tages of these materials are as follows:
• They are well castable and have a low thaw point,
which improves some other foundry characteristics;
• A proper alloying-element selection can eliminate the
occurrence of casting defects – microshrinks and
thermal cracks.
The AZ91 alloy is perhaps the most frequently used
from this group.2 A rather fast decrease in the strength
when stressing them with a growing temperature is a
disadvantage of these alloys. The alloys of the Mg-Al-Sr
and Mg-Al-RE types could compensate for the low
resistance to high temperatures with better microstruc-
tural stabilities and also fairly good strength properties3
under these conditions. These parameters are decisive
when using these materials for the parts that are exposed
to a considerable thermal stress, e.g., engine blocks. In
Materiali in tehnologije / Materials and technology 47 (2013) 4, 503–506 503
UDK 669.721.5:621.74:620.1 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(4)503(2013)
addition, another important requirement is the corrosion
resistance (internal and external) because the majority of
drive units are cooled with a liquid and the used mate-
rials have to be able to resist the corrosive effects of the
cooling medium for a long time. The alloys retaining a
sufficient strength under the temperatures of up to 250 °C
are of the Mg-Y-RE, Mg-Sc and Mg-Gd types without an
aluminium addition. These materials meet all the above
mentioned requirements for industrial use, but a proble-
matic issue is their economical availability. The require-
ment for a high resistance to creep is met by the mag-
nesium alloys with thorium (up to the temperatures of
370 °C). Then, the Mg-Li type alloys belong to the
lightest structural materials; however, it is difficult to
manufacture and treat them.
3 DESCRIPTION OF THE USED ALLOYS
The main alloying element of magnesium alloys is
aluminium. For the experimental evaluation both
commonly used alloys (AZ91, AZ91Be, AM60) and the
AMZ40 alloy were chosen as they had not yet been
industrially processed in the Czech Republic. All the
compared materials were supplied by a Czech manufac-
turer and Table 1 shows their chemical compositions
taken from the supplier’s certificates.
4 PREPARATION OF THE TEST SAMPLES
Castings were gravity cast in an iron mould
(Figure 1) that was preheated before the casting to
achieve a sufficient running property of the metal. The
samples for the tensile test (Figure 2), the metallo-
graphic analysis and the measurements of the other
mechanical properties were subsequently made from
these castings. The material was melted in a steel cru-
cible in an electric resistance furnace. To ensure the
metal’s protection during the melting, a covering pre-
paration with a commercial product named EMGESAL
was used to prevent excessive oxidation or burning on
the melt surface.
During the measurement the temperatures of the
metal mould and the cast alloy were observed for process
checking and description.
To ensure the improved mechanical properties of the
large-sized castings poured in the sand moulds, the melt
was treated in foundries with a preparation based on
hexachloroethane supplied under the commercial name
of MIKROSAL MG T 200. After introducing the nuclei,
the casting structure became finer and a fine-grained
structure of the material with improved mechanical
properties was formed. In our case this preparation was
also used for the melt.
5 MICROSTRUCTURES OF THE CAST
SAMPLES
The microstructures of the used alloys were evaluated
in different parts of the casting from which the test bars
were prepared. The differences between the structures of
the observed materials could only be found in the central
parts of the castings where the heat removal during the
solidification and cooling was not so intensive. Figures 3
and 4 show the microstructures of the AZ91 alloy cast in
a preheated metal mould. In the AZ91 alloy without an
inoculant addition (Figure 3) high shares of the Mg17Al12
P. LICHÝ et al.: INVESTIGATION OF THE THERMOMECHANICAL PROPERTIES AND MICROSTRUCTURE ...
504 Materiali in tehnologije / Materials and technology 47 (2013) 4, 503–506
Table 1: Chemical compositions of the used magnesium alloys in mass fractions, w/%
Tabela 1: Kemijska sestava uporabljenih magnezijevih zlitin v masnih dele`ih, w/%
Alloy
Element
Zn Al Si Cu Mn Fe Ni Ca Be residue
AZ91D 0.56 8.80 0.06 0.004 0.20 0.004 0.001 0.000 0.000 7 <0.01
AZ91Be 0.62 8.22 0.03 0.001 0.15 0.006 0.000 0.000 0.009 0 <0.01
AMZ40 0.14 3.76 0.02 0.001 0.34 0.003 0.000 0.000 0.001 1 <0.01
AM60 0.07 5.78 0.03 0.001 0.33 0.003 0.001 0.000 0.000 9 <0.01
Figure 2: Scheme of a test bar destined for the general tensile test
(dimensions are in millimeters)
Slika 2: Shematski prikaz preizkusne palice za natezni preskus
(dimenzije so v milimetrih)
Figure 1: Metal mould for casting the test bars according to the ^SN
42 0334 standard
Slika 1: Kovinska kokila za ulivanje preizkusnih palic, skladna s
standardom ^SN 42 0334
tabular precipitate and +	 eutectic can be observed. In
the alloy with an inoculant addition (Figure 4) the pre-
cipitate share was considerably higher. The situation was
similar in the case of AZ91Be, an alloy with a beryllium
addition.
The figures of the AM60 alloy microstructure evi-
dently show lower proportions of the Mg17Al12 precipi-
tate and the eutectic phase in the non-inoculated material
structure (Figure 5) as well as in the inoculated one
compared to the AZ91 alloys, due to a lower aluminium
content in the alloy. There are apparent material grain
boundaries in the microstructure figures and the inocu-
lant-addition effect is clearly evident here (Figure 6). In
the case of the AMZ40 alloy the difference between the
P. LICHÝ et al.: INVESTIGATION OF THE THERMOMECHANICAL PROPERTIES AND MICROSTRUCTURE ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 503–506 505
Figure 7: Non-inoculated material AMZ40
Slika 7: Material AMZ40 brez inokulacije
Figure 4: Inoculated material AZ91
Slika 4: Inokuliran material AZ91
Figure 6: Inoculated material AM60
Slika 6: Inokuliran material AM60
Figure 3: Non-inoculated material AZ91
Slika 3: Material AZ91 brez inokulacije
Figure 5: Non-inoculated material AM60
Slika 5: Material AM60 brez inokulacije
Figure 8: Inoculated material AMZ40
Slika 8: Inokuliran material AMZ40
non-inoculated material (Figure 7) and the inoculated
one (Figure 8) is the least significant of all the studied
alloys. This material structure with the minimum
proportion of the Mg17Al12 precipitate is very close to the
AM60 alloy due to their similar chemical compositions.
6 MEASUREMENT OF THE MECHANICAL
PROPERTIES AT ELEVATED TEMPERATURES
The tensile test was carried out on a TSM 20 ten-
sion-testing machine. The elevated-temperature test was
performed as a general one. When measuring mechanical
properties, we carried out the above mentioned general
test at room temperature and at elevated temperatures,
beginning at 100 °C, at a gradient of 50 °C, up to 300 °C.
The final temperature for the observed alloys varied and
it was selected according to the possibilities, with respect
to a limited amount of specimens. When testing at ele-
vated temperatures, the holding time at the temperature
for each specimen was 5 min for the temperature
equalization on the surface and inside the specimen.
7 ACHIEVED RESULTS
The influence of the test temperature on the resulting
values of the tensile strength of the materials was
evaluated with the tensile test. The graph in Figure 9
shows the results for the metallurgically untreated alloys
and for the alloys influenced by the inoculating agent.
The objectives of the tensile test were to compare the
particular alloy types, observe the inoculation effect and
the influence of the elevated temperature on mechanical
properties (the tensile strength). The highest tensile
strength values under room temperature were achieved
for the AZ91Be alloy (above 200 MPa). As a matter of
fact, the AZ91, AM60 and AMZ40 alloys showed the
same tensile strengths. A considerable decrease in the
tensile strengths of the AZ91, AZ91Be and AMZ40
alloys was observed with the growing test temperature.
Below the temperature of 300 °C the lowest strengths of
about 80 MPa were observed. The inoculation effect was
more significant only for the AZ91 and AM60 alloys,
where it was also reflected in the microstructures of
these alloys. During the tests under room temperature a
growth by 8 % or 10 %, respectively, was observed. With
the growing test temperature (above 150 °C) the tensile
strengths considerably decreased. The steepest drop was
observed for the AZ91Be alloy where the tensile strength
below the temperature of 200 °C was only about 120 MPa.
In the case of the AMZ40 alloy there was no inoculation
effect, which was also confirmed with the analysis of the
given-alloy microstructure.
8 CONCLUSION
The presented work aimed at studying the thermo-
mechanical properties of magnesium alloys. The highest
tensile strength was achieved for the AZ91Be alloy at
room temperature. The inoculation effect was more
significant only at low temperatures and for the AZ91
and AM60 alloys, where a grain refinement was evident
in the microstructures of the observed samples. In the
other materials no structure refinement was observed,
perhaps due to the extensive heat removal from the cast-
ings.
Acknowledgement
This work was performed within the frame of the
research project TA02011333 (Technology Agency of
the CR).
9 REFERENCES
1 M. J. F. Gándara, Mater. Tehnol., 45 (2011) 6, 633–637
2 B. L. Mordike, T. Ebert, Materials Science A, 302 (2001), 37–45
3 S. Roskosz, J. Adamiec, M. Blotnicki, Archives of Foundry Engi-
neering, 7 (2007) 1, 143–146
P. LICHÝ et al.: INVESTIGATION OF THE THERMOMECHANICAL PROPERTIES AND MICROSTRUCTURE ...
506 Materiali in tehnologije / Materials and technology 47 (2013) 4, 503–506
Figure 9: Thermal dependence of the tensile strength
Slika 9: Temperaturna odvisnost natezne trdnosti
L. KLIMES et al.: IMPACT OF CASTING SPEED ON THE TEMPERATURE FIELD ...
IMPACT OF CASTING SPEED ON THE TEMPERATURE
FIELD OF CONTINUOUSLY CAST STEEL BILLETS
VPLIV LIVNE HITROSTI NA TEMPERATURNO POLJE
KONTINUIRNO ULITIH GREDIC IZ JEKLA
Lubomir Klimes1, Josef Stetina1, Pavol Bucek2
1Brno University of Technology, Faculty of Mechanical Engineering, Technicka 2896/2, 616 69 Brno, Czech Republic
2Zeleziarne Podbrezova, Research and Development Center, Kolkaren 35, 976 81 Podbrezova, Slovakia
klimes@fme.vutbr.cz
Prejem rokopisa – received: 2012-08-31; sprejem za objavo – accepted for publication: 2013-01-04
In continuous casting, the casting speed is one of the most important parameters that influence the entire process of steel
production and its productivity. From the physical point of view, the casting speed affects the temperature field formation along
the cast steel blank, e.g., the surface and corner temperatures as well as the solidification of the steel, e.g., the shell thickness,
the isosolidus and isoliquidus curves, the metallurgical length and the width of the mushy zone. In order to achieve a particular
steel structure, the temperature ranges in particular positions (e.g., due to the straightening), and to minimize the occurrence of
defects, it is very important to pay attention to a proper determination of the casting speed. The aim of this paper is to
investigate the impact of the casting speed on the temperature field formation of continuously cast steel billets, and on the
aforementioned parameters. Three steel grades with various chemical compositions and the originally implemented numerical
model of the transient temperature field of continuously cast billets in Zeleziarne Podbrezova in Slovakia were utilized for the
analysis. The results proved the significant influence of the casting speed on the temperature field formation and on related
parameters. The conclusions may be utilized, e.g., for the determination of the casting speed and for the setup of the caster.
Keywords: continuous casting, casting speed, dynamic solidification model, temperature field
Pri kontinuirnem ulivanju je hitrost ulivanja eden od odlo~ilnih parametrov, ki vplivajo na celoten proces proizvodnje jekla in na
storilnost. Iz fizikalnega vidika hitrost ulivanja vpliva na nastanek temperaturnega polja vzdol` jeklene gredice, kot na primer na
temperaturo robov in povr{ine ulitega bloka, kot tudi na strjevanje jekla, kot na primer debelina strjene skorje, krivulje
izosolidusa in izolikvidusa, metalur{ka dol`ina in {irina ka{astega podro~ja. Da bi dosegli `eleno strukturo jekla, temperaturno
podro~je v dolo~enih polo`ajih (npr. zaradi ravnanja) in da bi zmanj{ali nastanek po{kodb, je zelo pomembno, da smo pozorni
na pravilno dolo~itev hitrosti ulivanja. Namen prispevka je raziskati vpliv hitrosti ulivanja na oblikovanje temperaturnega polja
pri kontinuirno ulitih gredicah in na `e navedene parametre. Za analize so bila uporabljena tri razli~na jekla z razli~nimi
kemijskimi sestavami in originalen numeri~ni model prehodnega temperaturnega polja gredic v @elezarni Podbrezova na
Slova{kem. Rezultati potrjujejo pomemben vpliv hitrosti ulivanja na nastanek temperaturnega polja in s tem povezanih
parametrov. Ugotovitve se lahko uporabijo na primer za dolo~itev hitrosti ulivanja in za postavitev livne naprave.
Klju~ne besede: kontinuirno ulivanje, hitrost ulivanja, dinami~ni model strjevanja, temperaturno polje
1 INTRODUCTION
The quality and productivity of continuously cast
steel blanks depend on many technological parameters
and caster options.1 The casting speed, which is directly
related to the productivity, is one of the most important
parameters for continuous steel casting. Its appropriate
setting according to the operating conditions, the steel
grade being cast and the caster parameters is crucial in
order to produce steel blanks with the desired quality and
structure, and also to minimize the occurrence of defects.
In addition, the produced steel and its properties can be
further optimized.2 The importance of the proper setting
of the casting speed is obvious when taking into account
the parameters on which the casting speed may have an
influence: mainly the entire temperature distribution
(e.g., the surface and corner temperatures that are very
often required to fit a certain range of temperatures, e.g.,
due to the straightening) along the cast blank, the
isosolidus and isoliquidus curves that characterize the
solidification process, the metallurgical length, or, e.g.,
the shell thickness along the blank.3
Nowadays, dynamic solidification models are com-
monly used in steelworks for the control and monitoring
of casting.4,5 The on-line models for the casting control
are fast and rather simple, whereas the off-line models
enable an analysis of multiple and detailed factors, but
require a long time to run.6 Numerical models of the
temperature field of continuous casting often utilize the
finite-difference method,7 the finite-volume method8, or
the finite-element method.9 Recently, new computational
approaches are also used, due to their specific advanta-
ges, e.g., the meshless finite-point method,10 the front-
tracking boundary-element method,11 or the very fast
moving-slice-based models.12
The numerical solidification models allow a precise
determination of the entire temperature distribution
along the cast blank, the prediction of surface defects
and the improvement of steel quality.13 The numerical
model in the off-line mode working as a simulator also
makes it possible to perform a case study and an analysis
of the operating parameters without any affect on the real
caster and the steel production in a steelworks.6 Hence, it
Materiali in tehnologije / Materials and technology 47 (2013) 4, 507–513 507
UDK 621.74.047:536.421.4 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(4)507(2013)
is a good idea to utilize the dynamic solidification
models in order to determine a suitable casting speed for
continuously cast blanks.
The purpose of this paper is to investigate the
influence of the casting speed on the temperature field
formation and the related parameters of continuously
cast billets. For the analysis, three steel grades and the
200 mm × 200 mm billets, which are cast in Zeleziarne
Podbrezova in Slovakia, are considered. The analysis
was carried out by means of the originally implemented
dynamic solidification model of continuously cast
billets.
2 DYNAMIC SOLIDIFICATION MODEL
The study of the influence of the casting speed on the
thermal behavior of continuously cast steel billets was
performed with the use of the originally developed
numerical model of the transient temperature field.14,15
The model, which is fully 3D, allows a calculation of the
transient temperature distribution along the entire,
continuously cast billets from the pouring level inside
the mould, through the secondary and tertiary cooling
zones, to the torch where the semi-infinite billets are cut
to desired lengths suitable for the next processing step
(Figure 1). Since the temperature field is symmetric with
respect to the longitudinal vertical plane (Figure 1), only
one half of the billet is considered.
The transient heat transfer and the solidification of
the billet is governed by the Fourier-Kirchhoff equation:
∂
∂
∂
∂
H
t
k T v
H
zz
= ∇⋅ ∇ +( ) (1)
where H is the volume enthalpy (J m–3), T is the tem-
perature (K), t is time (s), k is the thermal conductivity
(W m–1 K–1), vz is the casting speed (m s–1) and z is the
spatial coordinate along the cast billet (m). The mass
transfer and the fluid flow of the melt inside the billet
are neglected. The thermodynamic function of the volu-
me enthalpy16 in Eq. (1), which is defined as follows:
H T c L
f
d
T
( ) = −
⎛
⎝
⎜ ⎞
⎠
⎟∫ 
 
  f
s∂
∂0
(2)
where 
 is the density (kg m–3), c is the specific heat
(J kg–1 K–1), Lf is the latent heat (J kg–1) and fs is the
solid fraction (1), is used due to the release of the latent
heat of structural and mainly phase changes, which the
steel undergoes during its solidification.17
The model is completed with the initial (3) and
boundary conditions (4a)–(4d):
T Ti j k, ,
0 = 0 (3)
T Ti j
t
, , 0 = cast in the meniscus plane (4a)
− =k
T
n
∂
∂
0 in the plane of symmetry (4b)
− =k
T
n
q
∂
∂
 in the mould and beneath the rollers (4c)
− = − + −∞ ∞k
T
n
h T T T T
∂
∂
( ) ( ) 4 4
in the secondary and tertiary cooling zones (4d)
where T0 is the initial temperature (K), Tcast is the
casting temperature (K), q is the heat flux (W m–2), h is
the heat transfer coefficient (W m–2 K–1), T is the
ambient temperature (K),  = 5.67 · 10–8 W m–2 K–4 is
the Stefan-Boltzmann constant and  is the emissivity
(1).
In order to accurately and reliably predict the tem-
perature field of cast blanks by means of solidification
models, it is also worth pointing out the variability of the
thermophysical properties (H, k, 
, c: see Eq. (1) and
Eq. (2)) according to the chemical composition of the
steel being cast and their dependency on the temperature.
Due to the mentioned reasons, the developed solidifi-
cation model utilizes the results of the solidification
analysis package IDS.18 It enables the calculation of the
temperature dependency of the aforementioned thermo-
physical properties according to the chemical compo-
sition of the steel. The precise chemical composition can
be obtained, e.g., from the chemical analysis of a sample
from the tundish.
The dynamic solidification model of continuously
cast billets was created with the use of the control-
volume method.19 This approach is based on meshing the
entire 3D billet into control volumes, and then the energy
balance is established for each control volume of the
billet. With the use of the explicit discretization in time,19
the energy balance of the general control volume (i, j, k)
(Figures 1 and 2) in the Cartesian coordinates reads:
 ( )
(
, , , ,
, ,
Q v x y H H
x y z
t
H
i
i
z i j k
t
i j k
t
i j
=
−∑ + − =
=
1
6
1Δ Δ
Δ Δ Δ
Δ k
t t
i j k
tH+ −Δ , , )
(5)
where Q denotes the heat transfer rate (W), x, y, z
are the spatial dimensions (m) of the control volume and
L. KLIMES et al.: IMPACT OF CASTING SPEED ON THE TEMPERATURE FIELD ...
508 Materiali in tehnologije / Materials and technology 47 (2013) 4, 507–513
Figure 1: The billet caster and the mesh definition
Slika 1: Naprava za ulivanje gredic in opredelitev mre`e
t is the time step (s). The first term on the left-hand
side in Eq. (5) represents the sum of the conduction heat
transfer rates given by Fourier’s law19 to the control
volume from the neighboring six control volumes
(Figure 2). For instance, the heat transfer rate to the
control volume in the positive direction of the x-axis is
then:

, , , ,
Q k y z
T T
xx
i j k
t
i j k
t
+ −= −
−
Δ Δ
Δ
1
(6)
The second term on the left-hand side in Eq. (5) takes
into account the energy incoming to the control volume
due to the movement of the billet through the caster with
the casting speed vz. The term on the right-hand side in
Eq. (5) represents the change of the internal energy of
the control volume during the time step t, including the
release of the latent heat (comprised in the enthalpy) due
to the solidification.16,17
In the case of the explicit time discretization,19 the
unknown enthalpy of the general control volume (i, j, k)
in time t + t can be then directly calculated from Eq. (5)
with the use of known values of the temperature and
enthalpy in time t:
H H
t
x y z
Q v
t
z
Hi j k
t t
i j k
t
i
i
z i j k, , , , , ,
 (+
=
−= + +∑Δ
Δ
Δ Δ Δ
Δ
Δ1
6
1
t
i j k
tH− , , )
(7)
In the case of the boundary-control volume, the
particular heat transfer rates in Eq. (7) are replaced by
the corresponding heat transfer rates determined from
the boundary conditions (4a)–(4d). The desired tempe-
rature in time t + t is then recalculated from the
enthalpy given by Eq. (7) with the use of the tempe-
rature-enthalpy relationship provided by the solidifi-
cation package IDS according to the particular chemical
composition of the cast steel.
In the curved region (the radial part) of the caster, the
Cartesian coordinate system is transformed to the cylin-
drical system (the Cartesian coordinates y and z are
transformed to the radius coordinate and to the angle
coordinate, respectively) and the equations (5)–(7) are
appropriately modified.
The model used in the presented analysis utilizes
1136520 control volumes (21 in axis x, 41 in axis y, and
1320 in axis z) with the size of the general control volu-
me of 5 mm × 5 mm × 15 mm. Due to the unconditional
stability19 of the explicit time discretization and the
numerical stability also for higher casting speeds, the
time step was set to 0.35 s. With the described configura-
tion of the mesh and the time step, the numerical model
is still able to run faster than real time. For example, on
the HP Z800 workstation the computing time reaches
90 % of the real time.20
The heat transfer coefficients in the boundary con-
dition (4d) for the cooling nozzles in the secondary
cooling were determined experimentally in the labora-
tory on the hot models.21 Figure 3 presents the experi-
mentally measured heat-transfer coefficient for the
cooling nozzle TG1 used in the secondary cooling. In
order to quantify the heat withdrawal in the mould and
the secondary cooling zone,20 the heat flux in the mould
given by Eq. (4c) was equivalently expressed to the heat
transfer coefficient and plotted in Figure 4.
As can be seen from Eq. (1), Eq. (5), and Eq. (7), the
casting speed vz straightforwardly influences the entire
transient temperature distribution of cast billets, and
therefore the appropriate setting of the casting speed is a
crucial issue in continuous casting.
The developed dynamic solidification model was
experimentally verified on the billet caster with the use
L. KLIMES et al.: IMPACT OF CASTING SPEED ON THE TEMPERATURE FIELD ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 507–513 509
Figure 2: The energy balance of the general control volume in Carte-
sian coordinates
Slika 2: Energijska bilanca splo{nega kontrolnega volumna v karte-
zi~nih koordinatah
Figure 3: The heat transfer coefficient for the cooling nozzle TG1 and
the water flow 0.64 l/min in the centerline of the nozzle: a) perpen-
dicular to the casting direction, b) along the casting direction
Slika 3: Koeficient prenosa toplote za hladilno {obo TG1 in pretok
vode 0,64 l/min v simetrali {ob: a) pravokotno na smer ulivanja, b)
vzdol` smeri ulivanja
of two pyrometers placed behind the mould in the secon-
dary cooling zone (1500 mm from the pour level) and in
the tertiary cooling zone (15420 mm from the pour
level). Both the pyrometers were positioned in the
centerline of the small radius surface of the cast billets.
The maximum temperature difference between the
measured temperatures by the pyrometers and the
calculated temperatures by the solidification model was
40 °C.20
3 ANALYSIS SETUP AND PARAMETERS
The study was carried out for the 200 mm × 200 mm
steel billets that are continuously cast in Zeleziarne Pod-
brezova in Slovakia. The radial billet caster has the tubal
mould, it comprises the secondary cooling with three
independent cooling zones (denoted by S-0, S-1, and S-2
in figures) including 96 water cooling nozzles of three
types and two straightening mills (denoted by SM-1,
SM-2 in figures).
The following three steel grades are considered (only
the main elements are listed in mass fractions w): un-
alloyed fine-grained steel for constructions and welding
S355J2G3 (0.187 % C, 1.17 % Mn, 0.22 % Si, 0.016 % P,
and 0.012 % S), carbon steel for refining C60 (0.617 % C,
0.73 % Mn, 0.29 % Si, 0.014 % P, 0.013 % S, 0.03 % Cr,
0.05 % Ni, and 0.18 % Cu), and heat-resistant Cr-Mo
steel for use at higher temperatures 13CrMo4-5 (0.144 % C,
0.53 % Mn, 0.24 % Si, 0.007 % P, 0.006 % S, 0.98 % Cr,
0.46 % Mo, 0.12 % Cu, and 0.019 % Al). The range of
the casting speed was considered to be between
0.7 m/min and 1.1 m/min; all three mentioned steel
grades are commonly cast with a casting speed of
0.9-1.0 m/min. In figures, only the casting speeds
0.7 m/min, 0.9 m/min, and 1.1 m/min are pictured for
reasons of clarity.
For the analysis the casting temperature was set iden-
tically to the real casting process: 1553 °C for the grade
S355J2G3, 1515 °C for the grade C60, and 1547 °C for
the grade 13CrMo4-5. The analysis was carried out for
the identical setting of the caster, mainly the water flow
rates through the cooling nozzles in the secondary
cooling and through the mould for all the considered
steel grades.
4 RESULTS AND DISCUSSION
4.1 Surface and corner temperatures
The analysis confirmed a significant dependency of
the temperature field of the cast billets on the casting
speed. For all steel grades, a very similar thermal beha-
vior was observed (Figure 5 for the steel grade
S355J2G3): the higher casting speed, the higher both the
corner temperature and the surface temperature beneath
the nozzles. Moreover, with an increasing casting speed
the rise of both the surface and corner temperatures
decreases (Figure 5). As already mentioned, although all
three investigated steel grades have various chemical
L. KLIMES et al.: IMPACT OF CASTING SPEED ON THE TEMPERATURE FIELD ...
510 Materiali in tehnologije / Materials and technology 47 (2013) 4, 507–513
Figure 4: The heat transfer coefficient in the mould and in the
secondary cooling zone: a) small radius surface, b) large radius
surface of cast billet
Slika 4: Koeficient prenosa toplote v kokili in v sekundarni coni
hlajenja: a) povr{ina z majhno ukrivljenostjo, b) povr{ina z veliko
ukrivljenostjo lite gredice
Figure 5: The impact of the casting speed on the surface and corner
temperatures for the steel grade S355J2G3
Slika 5: Vpliv hitrosti ulivanja na temperaturo povr{ine in robov pri
jeklu S355J2G3
compositions and each of them is intended for a different
usage, their relative courses of the surface and the corner
temperatures along the caster with a casting speed in the
range between 0.7 m/min and 1.1 m/min are almost
identical. The only significant difference is the tempera-
ture shift of both the surface and the corner temperature
curves due to the various casting temperatures.
4.2 Isosolidus and isoliquidus curves, metallurgical
length, mushy zone
The thermal behavior inside the cast billets, the
solidification process, and the influence of the casting
speed on them can be investigated in the longitudinal
axial cross-section of the billet with the use of the iso-
solidus and isoliquidus curves. Moreover, the isosolidus
curve is directly related to the metallurgical length,
which is an important parameter of continuous casting.
Figure 6 shows the influence of the casting speed on the
solidification process of steel in the core of the billets:
the isosolidus and isoliquidus curves, and the mushy
zone where both the liquid and solid phases coexist.
As can be observed from Figure 6, the profile of the
isosolidus and isoliquidus curves, the metallurgical
length as well as the width of the mushy zone differs for
each of the three investigated steel grades. In general, the
higher casting speed makes both the isosolidus and
isoliquidus longer (a positive shift in the length) and
lengthens the metallurgical length as well (Figure 6).
Moreover, the higher casting speed also enlarges the
mushy zone. It implies that the positive shift of the iso-
solidus curve due to the increasing casting speed is larger
than the shift of the isoliquidus curve.
For instance, consider the steel grade C60 (see the
middle graph in Figure 6): for the casting speed of
0.7 m/min the metallurgical length is 10.1 m and the
width of the mushy zone (in the axis of the billet) is
4.5 m. For the casting speed of 0.9 m/min, the metallur-
gical length increases to 13.9 m and the width of the
mushy zone to 6.1 m. However, in the case of the casting
speed of 1.1 m/min, the metallurgical length reaches
17.6 m (e.g., in the positions of both the straightening
mills SM-1 and SM-2 the billet still contains the liquid
phase) and the mushy zones distends to 7.7 m.
4.3 Shell thickness
The casting speed also has an impact on the rate of
the shell thickness growth along the caster. A similar
shell thickness behavior was observed for all three
investigated steel grades. In general, the higher casting
speed, the slower rate of the shell thickness growth
(Figure 7). The shell inside the mould grows almost
linearly, and behind the mould the rate of the shell
thickness growth declines due to the abrupt reduction of
the heat withdrawal. In the secondary cooling the rate of
the shell thickness growth slowly increases owing to the
L. KLIMES et al.: IMPACT OF CASTING SPEED ON THE TEMPERATURE FIELD ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 507–513 511
Figure 7: The influence of the casting speed on the shell thickness for the steel grade C60
Slika 7: Vpliv hitrosti ulivanja na debelino skorje pri jeklu C60
Figure 6: The influence of the casting speed on the isosolidus and
isoliquidus curves, and the mushy zone
Slika 6: Vpliv hitrosti ulivanja na izosolidusne, izolikvidusne krivulje
in na ka{asto podro~je
heat withdrawal of the cooling nozzles and also of the
free convection and radiation heat transfer mechanisms.
For instance, Figure 7 for the steel grade C60. In the
length of 10.1 m where the billet cast with the casting
speed of 0.7 m/min is completely solidified (i.e., the
shell thickness is 100 mm), the shell thickness of the
billet cast with the casting speed of 0.9 m/min is only
50 mm and the shell thickness of the billet cast with the
casting speed of 1.1 m/min is even only 35 mm. The
billet cast with the casting speed of 0.9 m/min is com-
pletely solidified in the length of 13.8 m, where the shell
thickness of the billet cast with the casting speed of
1.1 m/min is only 55 mm.
4.4 Local period of solidification
The solidification process of the steel billets can also
be investigated by the local period of solidification,
which represents the local width of the mushy zone
expressed in time related to the casting speed. Thus, for
the given point (x, y) in the cross-section of the billet
(Figure 1), the local period of solidification lps (x, y)
expresses the time (s) needed for the solidification
(Figure 6), i.e., for cooling down from the liquidus tem-
perature (where the first grains of solid phase appear) to
the solidus temperature (where the last melt is solidified)
along the billet assuming its movement through the
caster with the casting velocity:
lps x y
L x y L x y
v z
( , )
( , ) ( , )
=
−solidus liquidus
(8)
where Lsolidus(x, y) (m) and Lliquidus(x, y) (m) are the
lengths of the isosolidus and isoliquidus curves in the
cross-section point (x, y) (Figure 6 plotted for y = 100
mm), and vz is the casting velocity (m s–1). The local
periods of solidification in the horizontal centerline of
the cross section (i.e., for the cross section points (x, y)
for which y = 100 mm) of cast billets are shown in
Figure 8.
In general, a higher casting speed makes the local
period of solidification longer (Figure 8). However, the
profiles of the local periods of solidification are different
for particular steel grades. This behavior is mainly
caused by the different steel structure attained during the
solidification of the particular steel grade directly related
to the chemical composition and by the setup of cool-
ing.20 It was shown that a modification of the chemical
composition in the range of the chemical composition
given by norms and standards can change the profile of
the local period of solidification.22
In the case of the grade S355J2G3, the slowest
solidification is in the core of the billet. However, the
slowest solidification of the grade C60 is not in the core,
but about 30 mm out of the core (width = ±30 mm in
Figure 8). The described behavior is even more evident
in the case of the grade 13CrMo4-5, where the slowest
solidification is in approximately the same position
(width = ±30 mm) as in the case of C60, but it needs an
extra 80 s to be solidified than the core of the billet
(width = 0 mm).
5 CONCLUSION
The paper presents the results of an analysis aimed at
the impact of the casting speed on the temperature field
of continuously cast steel billets and related parameters.
The analysis was carried out with the use of the origi-
nally implemented dynamic solidification and three
various steel grades were taken into account: common
unalloyed steel for constructions, carbon steel for
refining, and anticorrosive Cr-Mo steel for use at higher
temperatures. The casting speed in the range between 0.7
m/min and 1.1 m/min was considered.
The performed analysis proved that the casting speed
can significantly influence the temperature field and the
solidification during the continuous steel casting. In
general, the higher casting speed increases the tempera-
ture of the cast billets due to a reduced heat withdrawal.
Consequently, the isosolidus and isoliquidus curves
become longer and it also causes an increase of the
metallurgical length. The higher casting speed also
causes a drop in the shell growth as well as making the
mushy zone wider, which implies a longer local period
of solidification.
The study confirmed that the dynamic solidification
models make it possible to perform an analysis of the
caster and billets responses to various operating condi-
tions and situations. The results can then be used by ope-
rators to set-up the caster.
L. KLIMES et al.: IMPACT OF CASTING SPEED ON THE TEMPERATURE FIELD ...
512 Materiali in tehnologije / Materials and technology 47 (2013) 4, 507–513
Figure 8: The influence of the casting speed on the local period of
solidification in the axial cross-section
Slika 8: Vpliv hitrosti ulivanja na lokalni ~as strjevanja v osnem
prerezu
Acknowledgement
The presented research was supported by the project
GACR P107/11/1566 of the Czech Science Foundation
and by the BUT project FSI-J-12-22. The main author,
the holder of Brno PhD Talent Financial Aid sponsored
by Brno City Municipality, also gratefully acknowledges
that financial support.
6 REFERENCES
1 A. Cramb Ed., Making, shaping and treating of steel: casting, 11th
edition, Assn of Iron & Steel Engineers, 2003
2 P. Popela, Optimizing mechanical properties of iron during melting
process, ZAAM: Zeitschrift fur Angewandte Mathematik und
Mechanik, 77 (1997) 2, S649–S650
3 M. Sadat, A. H. Gheysari, S. Sadat, The effects of casting speed on
steel continuous casting process, Heat and Mass Transfer, 47 (2011)
12, 1601–1609
4 Y. Meng, B. G. Thomas, Heat-Transfer and Solidification Model of
Continuous Slab Casting: CON1D, Metallurgical and materials
transactions B, 34 (2003) 5, 685–705
5 J. Stetina, L. Klimes, T. Mauder, F. Kavicka, Final-structure pre-
diction of continuously cast billets, Mater. Tehnol., 46 (2012) 2,
155–160
6 B. G. Thomas, Modeling of the continuous casting of steel-past, pre-
sent, and future, Metallurgical and materials transactions B, 33
(2002) 6, 795–812
7 B. Petrus, K. Zheng, X. Zhou, B. G. Thomas, J. Bentsman, Real-
time, model-based spray-cooling control system for steel continuous
casting, Metallurgical and materials transactions B, 42 (2011) 1,
87–103
8 J. C. Ma, Z. Xie, Y. Ci, G. L. Jia, Simulation and application of
dynamic heat transfer model for improvement of continuous casting
process, Materials science and technology, 25 (2009) 5, 636–639
9 M. Gonzalez, M. B. Goldschmit, A. P. Assanelli, E. F. Berdaguer, E.
N. Dvorkin, Modeling of the solidification process in a continuous
casting installation for steel slabs, Metallurgical and materials tran-
sactions B, 34 (2003) 4, 455–473
10 L. Zhang, Y. M. Rong, H. F. Shen, T. Y. Huang, Solidification
modeling in continuous casting by finite point method, Journal of
materials processing technology, 192–193 (2007), 511–517
11 A. Fic, A. J. Nowak, R. Bialecki, Heat transfer analysis of the con-
tinuous casting process by the front tracking BEM, Engineering
analysis with boundary elements, 24 (2000) 3, 215–223
12 B. Sarler, R. Vertnik, A. Z. Lorbiecka, I. Vusanovic, B. Sencic, A
multiscale slice model for continuous casting of steel, International
conference on modeling of casting, welding and advanced solidifi-
cation processes, IOP Conference series – Materials science and
engineering, 33 (2012), 012021
13 C. A. Santos, J. A. Spim, A. Garcia, Mathematical modeling and
optimization strategies (genetic algorithm and knowledge base)
applied to the continuous casting of steel, Engineering applications
of artificial intelligence, 16 (2003) 5–6, 511–527
14 J. Stetina, F. Kavicka, T. Mauder, L. Klimes, Transient simulation
temperature field for continuous casting steel billet and slab,
Proceedings of METEC InSteelCon 2011, 13–23
15 J. Stetina, F. Kavicka, The influence of the chemical composition of
steels on the numerical simulation of a continuously cast slab, Mater.
Tehnol., 45 (2011) 4, 363–368
16 C. R. Swaminathan, V. R. Voller, A general enthalpy method for
modeling solidification processes, Metallurgical transactions B, 23
(1992) 5, 651–664
17 D. M. Stefanescu, Science and Engineering of Casting Solidification,
Second edition, Springer, 2009
18 J. Miettinen, S. Louhenkilpi, H. Kytonen, J. Laine, IDS: Thermo-
dynamic-kinetic-empirical tool for modelling of solidification,
microstructure and material properties, Mathematics and computers
in simulation, 80 (2010) 7, 1536–1150
19 F. P. Incropera, D. P. DeWitt, T. L. Bergman, A. S. Lavine, Principles
of heat and mass transfer, Seventh edition, Wiley, 2012
20 J. Stetina, L. Klimes, Analysis of continuous casting of steel billets
with the use of the transient 3D numerical model of temperature
field, Technical report for Zeleziarne Podbrezova, Slovakia, Brno
University of Technology, 2011
21 M. Raudensky, J. Horsky, Secondary cooling in continuous casting
and Leidenfrost temperature effects, Ironmaking and steelmaking, 32
(2005) 2, 159–164
22 L. Klimes, J. Stetina, L. Parilak, P. Bucek, Influence of Chemical
Composition of Cast Steel on the Temperature Field of Continuously
Cast Billets, Proceedings of Metal 2012, Ostrava: Tanger, 2012,
34–39
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Materiali in tehnologije / Materials and technology 47 (2013) 4, 507–513 513

M. DUCHEK et al.: DEVELOPMENT OF THE PRODUCTION OF ULTRAFINE-GRAINED TITANIUM ...
DEVELOPMENT OF THE PRODUCTION OF
ULTRAFINE-GRAINED TITANIUM WITH THE
CONFORM EQUIPMENT
RAZVOJ PROIZVODNJE ULTRADROBNOZRNATEGA TITANA Z
OPREMO CONFORM
Michal Duchek, Tomas Kubina, Josef Hodek, Jaromir Dlouhy
COMTES FHT a.s., Prùmyslová 995, 334 41 Dobrany, Czech Republic
michal.duchek@comtesfht.cz
Prejem rokopisa – received: 2012-09-10; sprejem za objavo – accepted for publication: 2012-12-18
Ultrafine-grained titanium belongs to the materials that possess a high potential for medical applications and, thanks to its
mechanical properties, it is already frequently employed in this field. The know-how for manufacturing ultrafine-grained
titanium represents one of the most complex issues in the titanium processing. The CONFORM process is one of the techniques
available for this purpose. The principle of the CONFORM technique is the continuous ECAP (Equal Channel Angular
Pressing) process. At present, it is primarily used for processing the materials with a low flow stress and a good formability.
This paper gives a description of this forming technique and its applications in the production of titanium semi-products.
Grade-4 titanium was used for the experimental programme. The factors explored in the process trials included the impacts of
the shape of the exit chamber and of the temperatures at the key locations within the CONFORM process on the resulting
microstructure of the titanium product. Numerical simulations of the process were performed in the DEFORM 3D software
using various friction conditions. The actual flow of the material was compared with the results of the simulation.
Keywords: titanium, CONFORM, ECAP, ultrafine grain
Ultradrobnozrnat titan spada k materialom, ki imajo velike mo`nosti za uporabo v medicini in se `e uporablja na tem podro~ju
po zaslugi svojih mehanskih lastnosti. Tehnologija izdelave ultradrobnozrnatega titana je najbolj kompleksen izziv pri njegovi
proizvodnji. Postopek CONFORM je ena od razpolo`ljivih tehnik za ta namen. Osnova tehnike CONFORM je kontinuiren
postopek ECAP (stiskanje skozi pravokoten kanal). Sedaj se uporablja za predelavo materialov z nizko napetostjo te~enja in
dobro preoblikovalnostjo. ^lanek opisuje to tehniko preoblikovanja in njeno uporabo pri proizvodnji titanovih polproizvodov.
Titan ~istosti grade-4 je bil uporabljen za preizkuse. V preizkusnih procesih so bili uporabljeni nekateri dejavniki na klju~nih
mestih pri CONFORM-u, vklju~no z u~inkom oblike izhodnega kanala in temperature, na mikrostrukturo proizvoda iz titana.
Numeri~ne simulacije procesa so bile izvr{ene s programsko opremo DEFORM 3D z uporabo razli~nih razmer pri trenju.
Resni~ni tok materiala je bil primerjan z rezultati simulacije.
Klju~ne besede: titan, CONFORM, ECAP, ultradrobna zrna
1 INTRODUCTION
Numerous SPD processes (Severe Plastic Deforma-
tion) were developed in the recent 15 years. These
processes are used for achieving a grain refinement in
materials, typically to a grain size between 100 nm and
400 nm. Their efficiency in processing a large volume of
a material is, however, still insufficient for industrial-
scale applications. This drawback was eliminated with
the CONFORM method. The method has been known
for a long time. Today, it is used for the continuous
industrial-scale production of sections, mostly from
aluminium. Essentially, it is an upgraded ECAP (Equal
Channel Angular Pressing) method, where the feedstock
is forced through a die by a friction force exerted by a
roll that turns ECAP into a continuous process.
This ECAP–CONFORM process was reported for the
first time in1. In this study, a commercial-grade alumi-
nium wire was extruded using one to four passes. The
preparation of ultrafine-grained titanium using this
method was reported in2. In this study, a 7.2 mm ×
7.2 mm square wire with an accumulated strain of e = 6
was subsequently redrawn into a diameter wire 3 mm
(resulting in the total strain of 7.9).
In the present paper, the production of commercial-
grade-titanium round bars using a single pass through the
ECAP-CONFORM equipment is described.
2 EXPERIMENTAL WORK
The feedstock consisted of the CP Ti grade 2 bar with
a diameter of 10 mm. Their chemical composition is
listed in Table 1. It was measured using a Bruker Q4
Tasman optical emission spectrometer and a Bruker G8
Galileo gas analyzer.
Table 1: Chemical composition of the feedstock in mass fractions (w/%)
Tabela 1: Kemijska sestava surovca v masnih dele`ih (w/%)
Fe O C H N Ti
0.046 0.12 0.023 0.0026 0.0076 99.822
Titanium bars were converted into the bars of the
same diameter as that of the feedstock using a CON-
FORM 315i machine.
Materiali in tehnologije / Materials and technology 47 (2013) 4, 515–518 515
UDK 669.295:621.777:519.876.5 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(4)515(2013)
The resulting microstructure was examined using a
Nikon Eclipse MA 200 optical microscope. The speci-
mens for the observation in the optical microscope were
prepared by metallographic grinding and polishing and
then etched using a Kroll agent. The specimens for an
EBSD analysis were prepared using a JEOL SM-09010
ion polisher. The EBSD analysis was conducted using a
JEOL 7400 electron microscope with a HKL-NORD-
LYSS EBSD camera. The data were processed using the
Chanel5 software.
The mechanical properties at room temperature were
measured using the cylindrical tension-test specimens
with a gauge length of 25 mm and a diameter of 5 mm.
In addition, impact-toughness tests were conducted using
3 mm × 4 mm section specimens.
The coefficient of the friction between CP-Ti and
Inconel (the material of the die chamber) was measured
by means of a tribological pin-on-disc test.
The process of continuous extrusion of titanium was
modelled using the DEFORM 3D software based on the
FEM method. The strain and temperature fields that
varied with time were analysed with mathematical tools.
The heat-transfer coefficient between the titanium feed-
stock and the wheel was measured to be 1000 W/(m2 K).
The input material data were obtained from the mate-
rial-data calculation in the JMatPRO software.
3 RESULTS AND DISCUSSION
In the trials of extruding a Ti wire with the CON-
FORM equipment, the key parameters varied: the speed
of the wheel, die-chamber temperatures, the cooling
downstream of the die chamber and others. The tempe-
rature vs. the time record from the experiment is shown
in Figure 1. The decisive parameter for the entire pro-
cess is the die-chamber temperature that varied between
the initial 500 °C and the final value of 350 °C. Numbers
1–4 in the figure denote the time instants, at which the
samples were taken for an EBSD analysis. The point, at
which the Ti wire was cooled right after exiting the die
chamber, is identified as well.
The microstructure of specimen 1 consists of unde-
formed equiaxed grains. The grain-size pattern is
bimodal and the average grain size is 1.9 μm. Neither the
small nor the large grains show signs of deformation.
The microstructure of specimen 2 is identical to that
of specimen 1. The chamber temperature of 450 °C is
sufficient for the microstructure to recover/recrystallize.
No effects of cooling were observed. The recovery/recry-
stallization and potential grain growth were finished
before the specimen was cooled.
Specimen 3 (the die-chamber temperature of 400 °C)
contains a small amount (10–15 %) of distorted unrecry-
stallized grains. The specimen was cooled with water
upon exiting the chamber. The deformed grains with a
size of no more than 5 μm × 10 μm are divided by low-
angle boundaries into subgrains. Their average size is
1.9 μm.
Specimen 4 was formed at the die-chamber tempe-
rature of 350 °C and then cooled with water upon exiting
the chamber. Its microstructure consists of slightly
elongated, deformed grains that, however, lack the
above-mentioned substructure (Figure 2). The EBSD
analysis focused on the centre of the circular cross-
section of the extruded product. The analysed surface is
on the plane that is parallel to the bending plane/flow
plane in the CONFORM chamber. On the EBSD maps
shown, the axis of the extruded section is vertical. The
grain size in this specimen was 1.4 μm.
The textures in all the specimens are aligned on the
basal plane. The texture of specimen 4 is documented in
Figure 3.
M. DUCHEK et al.: DEVELOPMENT OF THE PRODUCTION OF ULTRAFINE-GRAINED TITANIUM ...
516 Materiali in tehnologije / Materials and technology 47 (2013) 4, 515–518
Figure 1: Temperatures at the key locations during the ECAP-CON-
FORM processing of a titanium bar
Slika 1: Temperature na klju~nih mestih med ECAP-CONFORM
stiskanjem titana
Figure 2: EBSD micrograph of specimen no. 4
Slika 2: EBSD-posnetek vzorca {t. 4
The mechanical properties of the feedstock and of the
Ti wire upon a single pass at the die-chamber tempera-
ture of 350 °C are listed in Table 2. As expected, the
yield stress and ultimate strength of the product are
higher than those of the feedstock. On the other hand, its
contraction and elongation as well as the impact
toughness upon the first pass are lower than those of the
feedstock.
Figure 4 shows the macrostructure of the product in
the die chamber after the processing was interrupted. The
dead zones in the transitional region are clearly visible.
The traces of deformation can be seen at some points.
Figure 5a shows a micrograph of the undeformed
material. The average grain size in this case is 25 μm.
The microstructure above the abutment is shown in Fig-
M. DUCHEK et al.: DEVELOPMENT OF THE PRODUCTION OF ULTRAFINE-GRAINED TITANIUM ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 515–518 517
Figure 5: Microstructures in the selected locations within the die
chamber. The notation corresponds to that in Figure 4.
Slika 5: Mikrostrukture na izbranih lokacijah v komori. Oznake ustre-
zajo tistim na sliki 4.
Figure 4: Macrostructure of CP-Ti inside the die chamber
Slika 4: Makrostruktura v notranjosti CP-Ti komore
Figure 6: Distribution of the absolute material-flow velocity in
mm s–1. Variant: a) with the friction coefficient of 0.2; b) friction
coefficient of 0.7.
Slika 6: Razporeditev absolutne hitrosti toka materiala v mm s–1.
Varianta: a) koeficient trenja 0,2, b) koeficient trenja 0,7.
Figure 3: Areal densities of the points
Slika 3: Podro~ja gostote to~k
Table 2: Mechanical properties and the average grain size in various states of CP-Ti
Tabela 2: Mehanske lastnosti in povpre~na velikost zrn pri razli~nih stopnjah CP-Ti
Condition
Rp0.2 Rm Ag A5 Z KCV d
MPa MPa % % % J cm–2 μm
Feedstock 354.3 470.4 9.3 32.3 64.2 64.2 23.5
Upon pass 1 620.1 693.5 12.0 26.3 55.7 27.5 1.4
ure 5b. In this location, the grains are elongated in the
direction at an angle of 45°. The elongation is not uni-
form; instead, it shows a wave-like pattern. A better
alignment of the grains can be seen in Figure 5c, which
shows a micrograph of the bottom part of the die-exit
area. The dead zones exhibit the finest grains without
any traces of the flow lines (Figure 5d).
3.1 Comparison between mathematical simulation and
physical experiment
The coefficient of friction between CP-Ti and the tool
material was measured using a non-standard, tribological
test arrangement: μ = 0.62. The result is not conclusive,
as the conditions of the test cannot perfectly match those
of the real-world titanium-forming process.
Two FEM simulations were performed with two
different friction-coefficient values (0.2 and 0.7), while
keeping the other initial and boundary conditions iden-
tical.
A sample result of the calculation of the absolute
flow velocity is shown in Figure 6 for both values of the
friction coefficient. Upon comparing the results with an
actual specimen (Figure 4), it becomes apparent that the
value of μ = 0.7 is closer to the actual value. A com-
parison between the distribution of the zones with the
minimum movement of the material calculated by the
numerical simulations and that found in the specimen
upon interrupting the forming process suggests the same
result: the value of μ = 0.7 is more appropriate. The ob-
tained value is very far from the values of the coefficient
of friction during the wire drawing with lubrication.3
4 CONCLUSIONS
The process trials were performed, whereby the
impact of the settings of the CONFORM equipment on
the microstructure of the resulting Ti wire was explored.
The strongest influence is that of the die-chamber tempe-
rature. The smallest average grain size found with the
EBSD analysis is 1.4 μm. The limit temperature for
achieving this value in an uninterrupted process was
350 °C. The Ti wire produced under these conditions
showed a strength of 693 MPa and an impact toughness
of 27.5 J cm–2.
The main goal of this study was to describe the
continuous extrusion of titanium by means of a mathe-
matical model. The model was constructed in the FEM
software DEFORM-3D. The outcome of the simulations
is in good agreement with the experimental results. The
die chamber of the Conform™ machine was adapted for
titanium forming.
The coefficient of friction between titanium and the
die-chamber temperature was 0.7. In the further optimi-
sation efforts, the die chamber will be adjusted in order
to suppress the dead zones where no movement of the
material occurs. This will be first verified with a mathe-
matical simulation and then with additional process trials
including multiple passes in order to obtain a nanostruc-
tured titanium bar.
Acknowledgments
This study was undertaken as part of the TIP
FR-TI1/415 Research and development of nanostruc-
tured materials for medical applications project funded
by the Ministry of Industry and Trade of the Czech
Republic.
5 REFERENCES
1 G. J. Raab, R. Z. Valiev, T. C. Lowe, Y. T. Zhu, Continuous pro-
cessing of ultrafine grained Al by ECAP-Conform, Materials Science
and Engineering: A, 382 (2004) 1–2, 30–34
2 G. Raab et al., Long-Length Ultrafine-Grained Titanium Rods Pro-
duced by ECAP-Conform, Materials Science Forum, 584–586
(2008), 80–85
3 B. Plá{ek, V. Tittel, Lime-based lubricants improve wire drawing,
Wireworld International, 29 (1987) 2, 33–34
M. DUCHEK et al.: DEVELOPMENT OF THE PRODUCTION OF ULTRAFINE-GRAINED TITANIUM ...
518 Materiali in tehnologije / Materials and technology 47 (2013) 4, 515–518
T. KUBINA et al.: CONVERSION OF THE Ms70 ALPHA-BRASS TORSION-TEST DATA ...
CONVERSION OF THE Ms70 ALPHA-BRASS
TORSION-TEST DATA INTO HOT-FORMING
PROCESSING MAPS
PRETVORBA PODATKOV TORZIJSKEGA PREIZKUSA ALFA
MEDI Ms70 V PROCESNE ZEMLJEVIDE VRO^E PREDELAVE
Tomá{ Kubina1, Josef Boøuta2, Rudolf Pernis3
1COMTES FHT, Prumyslova 995, 334 41 Dobrany, Czech Republic
2Material & Metallurgical Research Ltd., Pohranicni 693/31, 706 02 Ostrava-Vitkovice, Czech Republic
3Faculty of Special Technology, Alexander Dubcek University in Trencin, Trencin, Slovak Republic
tomas.kubina@comtesfht.cz
Prejem rokopisa – received: 2012-09-16; sprejem za objavo – accepted for publication: 2012-12-03
The present paper gives a summary of the plastic behaviour of the Ms70 deep-drawing cartridge brass under the hot-forming
conditions explored by using a SETARAM-Vítkovice torsion plastometer. The specimens were cut from the hot-extruded and,
subsequently, cold-drawn brass bars.
Formability tests were performed on the diameter 6 mm specimens at (650, 700, 750, 800 and 850) °C and at the speeds of (16,
80, 400 and 800) r/min. These testing conditions are equivalent to the conditions of the ordinary industrial hot extrusion of
brass: the single most problematic operation in the manufacturing process in terms of an occurrence of serious defects.
The results of the torsion tests were converted into the conventional format: the maximum flow-stress levels vs. the peak-strain
intensity. The points thus obtained can be used for finding the activation energy for forming Q, which is useful for determining
the thermally compensated strain rate with the aid of the Zener-Hollomon Z parameter.
The power-dissipation maps with the axes showing the deformation temperature and logarithmic strain rate were computed
using a dynamic material model. The calculation of the dissipation efficiency was verified by comparing it with the known data
on the 30 % zinc brass. The maps constructed in this manner allow the data from the torsion and compression plastometers to be
compared. The paper also includes a discussion of the changes in the processing maps at higher strains showing the optimum
regions for the hot extrusion of brass.
Keywords: flow stress, brass, hot formability, strain rate, torsion test, power-dissipation maps
V ~lanku je povzetek plasti~nega vedenja globokovle~ne medi Ms70 pri vro~i predelavi z uporabo torzijskega plastometra
SETARAM-Vítkovice. Vzorci so bili izrezani iz vro~e iztiskane in nato hladno vle~ene palice iz medi.
Preizkusi preoblikovalnosti so bili izvr{eni na vzorcih premera 6 mm pri temperaturah (650, 700, 750, 800 in 850) °C in
hitrostih (16, 80, 400 in 800) r/min. Te razmere pri preizkusih so enakovredne tistim pri navadnem industrijskem iztiskanju
medi; najbolj problemati~ni operaciji pri procesu izdelave zaradi pojava huj{ih po{kodb.
Rezultati torzijskih preizkusov so bili pretvorjeni v navadno obliko: maksimalne natezne napetosti proti najve~jemu raztezku.
Tako dobljene to~ke se lahko uporabi za izra~un aktivacijske energije preoblikovanja Q, ki je uporabna za dolo~anje toplotno
kompenzirane hitrosti trganja z uporabo Zener-Hollomon Z-parametra.
Zemljevidi spreminjanja mo~i z osmi, ki prikazujejo temperaturo deformacije in logaritem hitrosti preoblikovanja, so bili
izra~unani z dinami~nim modelom materiala. Izra~un spreminjanja u~inkovitosti je bil preverjen s primerjavo poznanih
podatkov za med s 30 % cinka. Tako konstruirani zemljevidi omogo~ajo primerjavo podatkov iz torzije in plastometra. ^lanek
vklju~uje tudi razpravo sprememb v zemljevidih procesiranja pri vi{jih napetostih in prikazuje optimalna podro~ja za vro~o
ekstruzijo medi.
Klju~ne besede: napetost te~enja, med, vro~a preoblikovalnost, hitrost preoblikovanja, torzijski preizkus, mape spreminjanja
mo~i
1 INTRODUCTION
The Ms70 brass belongs to the group of alpha brasses
with deep-drawing properties that are used for making
cartridge shells in a cold process. The conventional
method of the production of brass cartridge shells uses
rolled feedstock. The punching of the blanks from a strip
leads to large amounts of waste material that requires
reprocessing in metallurgical plants. Munitions plants
have launched the projects introducing a zero-waste
production of brass cartridge shells from Ms70 bars.
This change from the rolled strip feedstock to hot-
extruded bars required a technology upgrade in the
metallurgical plants.
The hot-extrusion process proved to be the critical
stage in the production of brass bars. As the Ms70 brass
exhibits excellent cold formability, it is difficult to form
it at high temperatures. The flow stress of the Ms70 brass
in the extrusion process is governed primarily by the
forming temperature, the strain rate and the strain. The
most common technique for exploring the flow stress at
elevated temperatures is the tension testing.1,2 However,
its results are not adequate for mapping the forming
processes at high temperatures and high strain rates. As
measuring the flow stress in a production is not
cost-effective, the forming process was simulated using a
torsion plastometer.3
A conversion of the results of the compression
plastometer testing into the processing maps for the
Materiali in tehnologije / Materials and technology 47 (2013) 4, 519–523 519
UDK 669.35'5:620.175:621.777 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(4)519(2013)
30 % zinc brass was reported in4,5. The evaluation was
based on the dynamic material model reported by
Prasad6. The total energy absorbed by the deformed body
P is converted into two forms: G – causing a rise in the
temperature; and J – leading to microstructural changes.
The factor that partitions the energy between G and J is
the strain-rate coefficient m. There is an effect of the
stress as well. The J content can be calculated using the
following equation:
J
m
m
=
+

1
(1)
where  is the flow stress and  denotes the strain rate.
An ideal linear dissipator term is given by m = 1 and
J = Jmax =    . The dissipation efficiency for a non-
linear dissipator may be expressed as a dimensionless
parameter:
= =
+
J
J
m
mmax ( )
2
1
(2)
The variation of  with the temperature and strain
rate represents a dissipation of the power through the
microstructural changes in the workpiece material. The
purpose of this study was to compare and contrast the
energy-dissipation maps constructed using the data from
the compression and torsion plastometers.
2 MATERIALS AND METHODS
The computer that controls the plastometer records
the following quantities: time, torque Mk, axial force F,
temperature T and the number of twists N. The test
specimen for a torsion plastometer is a bar with the
diameter of D (radius R) and a gauge length of L. The
key output of the torsion test for a formability assess-
ment is the number of twists to fracture Nf. As twisting
causes the test bar to shrink in length, the grips are fixed
from the beginning of the twisting cycle. Consequently,
the axial force F begins to act on the bar. The flow stress
is given by the following equation:
 =
⎛
⎝
⎜
⎞
⎠
⎟ +⎛⎝
⎜ ⎞
⎠
⎟
3 3
2 3
2
2
2M
R
F
R
k
π π
(3)
The strain intensity e in the torsion testing was found
from:
e
RN
L
= ⎛⎝
⎜ ⎞
⎠
⎟2
3
2
3
arg sinh
π
(4)
In a hot-torsion test, the stress-strain dependence is
monitored under the prescribed conditions (specimen
geometry, temperature, strain rate).7
The experimental hot-torsion testing was performed
using the Ms70 deep-drawing brass with a chemical
composition given in Table 1. The test specimens were
10 mm diameter bars. Ms70 brass bars were produced in
the following sequence: melting in an electric induction
furnace, semi-continuous casting of round ingots, hot
extrusion and final sizing in a cold-drawing process.
The hot-torsion testing of Ms70 brass bars was
performed according to a 4 × 5 schedule. This means
that four strain-rate values of (0.2, 1, 5 and 10) s–1
corresponding to the twisting speeds of (16, 80, 400 and
800) min–1 were combined with five temperature levels
of T is (650, 700, 750, 800 and 850) °C.
3 DISCUSSION OF RESULTS
The results of the torsion testing at 650 °C are shown
in Figure 1. The tests at the other temperatures were eva-
luated in an identical manner.8
The peak stress values p and the corresponding
strain levels ep were derived from these plots. The peak
stress (the maximum stress) p is governed by two
variables: the temperature and the shear-strain rate. This
dependency can be expressed as the function p = f(t, e),
T. KUBINA et al.: CONVERSION OF THE Ms70 ALPHA-BRASS TORSION-TEST DATA ...
520 Materiali in tehnologije / Materials and technology 47 (2013) 4, 519–523
Table 1: Chemical composition of the Ms70 brass for torsion testing
Tabela 1: Kemijska sestava medi Ms70 za torzijske preizkuse
Element Cu Pb Sn Fe Ni Mn Al Si
Content, w/% 70.39 0.0004 0.0042 0.0232 0.0022 0.0003 0.0012 0.0002
Element As Sb Bi Cr Cd Ag P Zn
Content, w/% 0.0001 0.0031 0.0001 0.0001 0.0001 0.0001 0.0002 balance
Figure 1: Flow stress of the Ms70 brass at 650 °C
Slika 1: Krivulje te~enja medi Ms70 pri 650 °C
which is shown as a 3D-plot in Figure 2. This depen-
dence is in line with the expected deformation behaviour
of the single-phase materials. The peak flow-stress
values decline with the increasing temperature and
decreasing strain rate. The p – e plots reveal that the
Ms70 brass undergoes a recovery during the defor-
mation. The shapes of these curves suggest that a
dynamic recrystallization is the primary mechanism.
Values p and ep were used to calculate the activation
energy of the deformation and were published in an
article.9
The ultimate strain ef (strain to fracture) is another
characteristic of a material’s formability. The values of
the strain to fracture for a torsion testing of the Ms70
brass are summarised in Figure 3. The effects of the
differences between the strain rates are indistinct but
certain trends can be found. Despite the scatter in the
measured values resulting from the sensitivity of the
torsion test to the inhomogeneity of the test bar surface,
it is clear that the optimum forming temperature is
between 800 °C and 850 °C.
Power-dissipation maps were constructed using the
dynamic material model. The power-dissipation coeffi-
cient  was found using the following procedure: The
stress-strain rate function was transformed into the
log-stress vs. log-strain rate relationship (Figure 4).
T. KUBINA et al.: CONVERSION OF THE Ms70 ALPHA-BRASS TORSION-TEST DATA ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 519–523 521
Figure 5: Power-dissipation map for the alpha brass with 30 % Zn
constructed from the data from10 for the strain of 0.5. The numbers
represent dissipation efficiency in fractions (%).
Slika 5: Zemljevid raztrosa mo~i za alfa med s 30 % Zn, konstruiran
iz podatkov iz10 za hitrost trganja 0,5. [tevilke pomenijo u~inkovitost
raztrosa v dele`ih (%).
Figure 3: Relationship between the strain intensity to fracture and
temperature
Slika 3: Odvisnost med intenziteto napetosti do preloma in tempe-
raturo
Figure 4: Conversion of the results of the plastometer testing at
650 °C from the recorded stress values at the strain of e = 0.5 into the
strain-rate coefficient m
Slika 4: Pretvorba rezultatov preizkusov na plastometru pri 650 °C iz
zabele`enih vrednosti napetosti pri hitrosti trganja e = 0,5 v koeficient
hitrosti natezanja m
Figure 2: Representation of the p = f(t, e) function in a 3D plot
Slika 2: Prikaz funkcije p = f(t, e) v 3D-diagramu
Cubic splines were used to find the log-stress values
between the experimentally measured points, from which
the strain-rate coefficient was calculated. Figure 4 illu-
strates three different dependences at the temperature of
650 °C. The stress curve from the torsion plastometer
has a narrow range of strain rates. The other curves are
based on the values for the 30 % Zn brass published in10
and the values for the brass with the mass fraction of Zn
30 % and 0.22 % Zr reported in5. The strain-rate coeffi-
cient m vs. the log-strain-rate relationship is very similar
in all the CuZn30 brasses, regardless of the testing-
machine type. The peak value of the m coefficient is
lower in the brass with a 0.22 % Zr addition.
Figure 5 shows the power-dissipation map for the
0.5 strain in the brass with 30 % zinc. This map was
developed using the values published in10. Upon the
comparison with the map shown in10, certain differences
become apparent, such as that between the high
temperature and low-strain rate areas. In the original
map, dissipation increases in this area. In the map con-
structed upon a re-calculation, a region with a decrease
followed by an increase is found. The maps re-calculated
for the brass with 30 % zinc and 0.22 % Zr do not exhi-
bit any substantial differences (Figure 6). Their charac-
ters are very similar but the absolute values are shifted.
This may be attributed to the algorithm used for con-
structing the 3D map. The new procedure obviates
excessive smoothing of dissipation-coefficient curves at
the individual temperature levels. The older maps were
constructed to verify the entire procedure: to allow a
comparison with the power-dissipation map for the 0.5
strain obtained from the torsion test.
As evidenced by Figure 7, the ranges of the strain
rates and temperatures are narrower than in the studies
reported in5,10. The SETARAM torsion plastometer
T. KUBINA et al.: CONVERSION OF THE Ms70 ALPHA-BRASS TORSION-TEST DATA ...
522 Materiali in tehnologije / Materials and technology 47 (2013) 4, 519–523
Figure 7: Power-dissipation map for the alpha brass with a chemical
composition according to Table 1 constructed from the data measured
with the SETARAM-VÍTKOVICE torsion plastometer for the strain
rate of 0.5. The numbers represent dissipation efficiency in fractions
(%).
Slika 7: Zemljevid raztrosa mo~i za alfa med s kemi~no sestavo iz
tabele 1, konstruiran iz podatkov, izmerjenih s torzijskim plastome-
trom SETARAM- VÍTKOVICE , za hitrost trganja 0,5. [tevilke pome-
nijo u~inkovitost raztrosa v dele`ih (%).
Figure 8: Power-dissipation map for the alpha brass with a chemical
composition according to Table 1 constructed from the data measured
with the SETARAM-VÍTKOVICE torsion plastometer for the strain
rate of 0.9. The numbers represent dissipation efficiency in fractions
(%).
Slika 8: Zemljevid raztrosa mo~i za alfa med s kemi~no sestavo iz
tabele 1, konstruiran iz podatkov izmerjenih s torzijskim plastome-
trom SETARAM- VÍTKOVICE, za hitrost trganja 0,9. [tevilke pome-
nijo u~inkovitost raztrosa v dele`ih (%).
Figure 6: Power-dissipation map for the alpha brass with 30 % Zn and
0.20 % Zr constructed from the data from5 for the strain of 0.5. The
numbers represent dissipation efficiency in fractions (%).
Slika 6: Zemljevid raztrosa mo~i za alfa med s 30 % Zn in 0,20 % Zr,
konstruiran iz podatkov iz5 za hitrost trganja 0,5. [tevilke pomenijo
u~inkovitost raztrosa v dele`ih (%).
cannot cover the range of high strain rates due to the
physical limitations of the torsion testing. The lower
strain rates are achievable with the use of an additional
gearbox. However, they were not explored as the original
experiment was aimed at the dynamic recrystallization as
the governing recovery process. A comparison with the
re-calculated power-dissipation maps in Figures 5 and 6
shows a good agreement with the shape for the brass
with 30 % zinc (Figure 5). However, substantial
deviations can be found at the points representing the
limit strain rates, as the procedure for finding the m
values is very sensitive to the algorithm for calculating
cubic splines and to the accuracy of the flow-stress
values. The strength of the torsion test lies in the use of
the large strains, which are not achievable in the com-
pression plastometers. Consequently, the power-dissi-
pation maps were available for the high strains as well,
such as that in Figure 8 for the 0.9 strain. The map
documents a high efficiency of the dissipation through
the microstructural changes at the lower strain rates for
all the deformation temperatures.
4 CONCLUSION
The torsion test allows the temperature and
strain-rate dependences of the flow stress to be studied
under various deformation conditions. Using the peak
stress values found, i.e., the values representing the flow
stress, the extrusion force can be calculated and the
examples for Ms70 are published in11. The torsion simu-
lation over a wide range of deformation temperatures
suggested that the capacity of the Ms70 brass for defor-
mation may become depleted. A technological testing of
hot extrusion12,13 showed that the formability of brass is
poor, which leads to transverse surface cracking. The
results of the torsion plastometer testing were converted
into the power-dissipation maps for selected strain levels.
The dissipation-efficiency  values are comparable with
the identically processed literature data. However, there
are variances resulting from the difference between the
algorithms for computing 3D maps and from the
sensitivity to the correct data processing of the boundary
strain rates.
Acknowledgment
The results presented in this paper were obtained
within the project West-Bohemian Centre of Materials
and Metallurgy CZ.1.05/2.1.00/03.0077, co-funded by
the European Regional Development Fund.
5 REFERENCES
1 M. Bur{ák, J. Bacsó, Skú{anie, kontrola a hodnotenie kvality mate-
riálov, Emilena, Ko{ice 2008
2 R. Pernis, O. Híre{, J. Kasala, Rozdelenie mechanických vlastností
ahanej mosadznej ty~e po priereze, In: Metal 2009, 18th interna-
tional material and metallurgical conferences, Hradec nad Moravicí,
^R [CD-ROM], Ostrava, Tanger, 2009, 8
3 I. Schindler, J. Boøuta, Utilization Potentialities of the Torsion
Plastometer, Silesian Technical University, Katowice, 1998, 104
4 D. Padmavardhani, Y. Prasad, Effect of zinc content on the pro-
cessing map for hot-working of alpha-brass, Materials science and
engineering A, 157 (1992) 1, 43–51
5 D. Padmavardhani, Y. Prasad, Effect Of Zirconium On The Pro-
cessing Maps For Hot-working Of Alpha And Alpha-beta Brass,
Zeitschrift Fur Metallkunde, 84 (1993) 1, 57–62
6 Y. Prasad et al., Modeling of dynamic material behavior in hot defor-
mation: Forging of Ti-6242, Metallurgical and Materials Transac-
tions A, 15 (1984), 1883–1892
7 J. Boøuta et al., Plastometrický výzkum deforma~ního chování øízenì
tváøených materiálù, Hutnické listy, 61 (2008) 1, 80–87
8 J. Boøuta, T. Kubina, A. Boøuta, Deforma~ní napìtí mosazi Ms70 pøi
zkou{ce krutem za tepla, Kovárenství, (2009) 39, 13–16
9 R. Pernis, J. Kasala, J. Boøuta, High temperature plastic deformation
of CuZn30 brass – calculation of the activation energy, Kovove
Materialy, 48 (2010) 1, 41–46
10 D. Padmavardhani, Y. Prasad, Characterization of hot deformation-
behavior of brasses using processing maps. 1. alpha-brass, Metallur-
gical Transactions A, 22 (1991) 12, 2985–2992
11 R. Pernis, J. Bidulská, Nový prístup pri výpo~te lisovacej sily z
výsledkov krutovej skú{ky, Výrobné in`inierstvo, 7 (2008) 4, 17–19
12 J. Kasala, O. Híre{, R. Pernis, Start-up phase modeling of semi
continuous casting process of brass billets, In: Metal 2009, 18th
international material and metallurgical conferences, Hradec nad
Moravicí, ^R [CD-ROM], Ostrava, Tanger, 2009, 6
13 R. Pernis, O. Híre{, J. Jurenová, Povrchové trhliny pri lisovaní
mosadze CuZn30 za tepla, In: Metal 2007, 16th international material
and metallurgical conferences, Hradec nad Moravicí, ^R
[CD-ROM], Ostrava, Tanger, 2007, 9
T. KUBINA et al.: CONVERSION OF THE Ms70 ALPHA-BRASS TORSION-TEST DATA ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 519–523 523

F. D. ALSEWAILEM, S. A. ALJLIL: RECYCLED POLYMER/CLAY COMPOSITES FOR HEAVY-METALS ADSORPTION
RECYCLED POLYMER/CLAY COMPOSITES FOR
HEAVY-METALS ADSORPTION
RECIKLIRAN KOMPOZIT POLIMER-GLINA ZA ADSORPCIJO
TE@KIH KOVIN
Fares D. Alsewailem1, Saad A. Aljlil2
1Petrochemicals Research Institute, King Abdulaziz City for Science and Technology (Kacst), P.O. Box 6086, 11442 Riyadh, Saudi Arabia
2The National Center for Water Technology, King Abdulaziz City for Science and Technology (Kacst), P.O. Box 6086, 11442 Riyadh, Saudi Arabia
fsewailm@kacst.edu.sa
Prejem rokopisa – received: 2012-09-26; sprejem za objavo – accepted for publication: 2012-11-27
Post-consumer plastics in the form of disposable cutlery, plates, cups, and containers for dairy products that are made of
polystyrene (PS) were used to formulate polymer/clay composite materials for the purpose of using them as an adsorbent for the
removal of heavy metals from waste water. The clay used in this study was a natural clay obtained from the Tabuk region in the
north of Saudi Arabia. The composites were formulated by mixing 5–10 % mass fractions of clay with PS in the molten state
using conventional polymer-processing equipment. Composites of virgin PS and clay were also formulated in order to compare
their ability to adsorb heavy metals with that of the recycled composites. The results showed that composites based on recycled
PS have a good adsorption capability in comparison with those of the virgin composites. For example, the amount of lead
adsorbed per gram of adsorbent was tripled when using composites containing only 5 % of clay and 95 % of recycled PS in
comparison with those with virgin PS.
Keywords: clay, polystyrene, recycled PS, heavy metals
Iz odpadne potro{ni{ke plastike v obliki jedilnega pribora za enkratno uporabo: kro`nikov, skodelic in embala`e za mle~ne
proizvode, iz polistirena (PS), je bil pripravljen kompozitni material polimer-glina kot adsorbent za odstranitev te`kih kovin iz
odpadne vode. Glina, uporabljena v tej {tudiji, je naravna glina, pridobljena iz obmo~ja Tabuk na severu Savdske Arabije.
Kompoziti so bili izdelani z me{anjem masnih dele`ev 5–10 % gline s PS v staljenem stanju s konvencionalno opremo za
predelavo polimerov. Za primerjavo zmo`nosti adsorpcije te`kih kovin so bili izdelani kompoziti z devi{kim PS in glino ter
reciklirani kompoziti. Rezultati so pokazali, da imajo kompoziti na osnovi recikliranega PS ve~jo vpojno zmogljivost v
primerjavi z devi{kim kompozitom. Na primer, koli~ina adsorbiranega svinca na gram adsorbenta se je v primerjavi z devi{kim
PS pri uporabi kompozita s 5 % gline in 95 % recikliranega PS potrojila.
Klju~ne besede: glina, polistiren, reciklirani PS, te`ke kovine
1 INTRODUCTION
Waste water including heavy metals is an inimitable
class of waste water, with all heavy metals being toxic
pollutants. Lead ions in waste water, for example, is a
pollutant and dangerous to human and water environ-
ments. As a result, it is necessary to eliminate heavy
metals from the waste water before discharging it into
the environment. The removal of heavy metals from
aqueous solutions may be achieved by applying several
physical and chemical methods, such as ion exchange,
extraction, flotation, coagulation, electro-deposition and
precipitation. However, most of these methods are
considered non-viable due to the cost factor. Various
materials have been utilized for heavy-metals removal
from waste water include activated carbon, lime and
bio-sorbents.1 For the large-scale treatment of waste
water for heavy-metals removal we need to look for a
low-cost material as an alternative, cost-effective
adsorbent for the removal of heavy metals such as lead
ions from waste water.
Polymeric adsorbents may be used in the adsorption
of heavy metals because of their easy regeneration and
strong mechanical properties in comparison with other
adsorbents such as activated carbon, cellulose and silica
gel.2 For example, CHA-111 AND MCH-111 polymeric
adsorbents were used as perfect adsorbents for adsorp-
tion pollutants such as phenols from waste water.3–5
Despite the fact that polymers are preferably utilized to
adsorb heavy-metal ions, the higher costs of these
materials may hinder their employment as cost-effective
adsorbents for the large-scale production of heavy-metal
free water.
In this work we tried to look for low-cost adsorbents
for the removal of lead from waste water by choosing
two discarded materials that can be obtained with no
cost. These two materials were natural clay and dis-
carded post-consumer polystyrene. In our previous publi-
cation6 we proved that natural clay taken from the north
of Saudi Arabia may be used as a cost-effective
adsorbent for lead ions. In the current work, we report on
the formulation of clay with post-consumer polystyrene
articles that are eventually discarded in landfills as waste
materials. Recycled polystyrene/clay composite material
for the purpose of using it as an adsorbent for heavy-
metals removal from waste water was formulated by the
current research. The clay used in this study was a
natural clay obtained from the Tabuk region in the north
of Saudi Arabia. The composites were formulated by
mixing the mass fractions 5–10 % of clay with poly-
Materiali in tehnologije / Materials and technology 47 (2013) 4, 525–529 525
UDK 678.7:544.723 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(4)525(2013)
styrene (PS) in the molten state using conventional
polymer-processing equipment. Composites of virgin PS
and clay were also formulated in order to compare their
ability to adsorb heavy metals with that of the recycled
composites. The adsorption isotherms of the two poly-
mer composites are also discussed here and the equili-
brium parameters required to design a batch absorber are
also estimated.
2 EXPERIMENTAL
2.1 Materials
The polymers used in this study were virgin poly-
styrene and recycled polystyrene. The virgin polystyrene
was supplied by the SABIC Company in Saudi Arabia,
while the recycled polystyrene was post-consumer
plastics in the form of disposable cutlery, plates, cups,
and containers for dairy products that are made of poly-
styrene. A lead-ion solution was prepared from lead
nitrate purified LR [Pb(NO3)2] and was supplied by
S.define- Limited ( Laboratory Rasayan).
2.2 Preparation of the polymer composites
The thermoplastic polymer is polystyrene (PS).
Polystyrene has a suitable affinity for clay particles and
is readily melt processed using conventional techniques
such as internal mixing and extrusion. The polymer can
be a so-called "virgin" polymer, i.e., a newly formed
polymer, or it can be derived from polymer scrap, such
as recycled polymer, as it permits the reuse of polymer
materials which might otherwise be discarded, necessi-
tating disposal in a landfill or similar.
The clay material was milled to a size below about
100 mesh and then washed with distilled water several
times to remove the impurities. The clay was then dried
in a vacuum oven overnight. The dried clay was then dry
mixed with polymer particles, and an emulsifier, alkyl-
trimethyl-ammonium, was incorporated to ease the dis-
persion of the clay in the polymer matrix. The com-
posites were formulated by mixing 5–10 % of clay with
PS in the molten state using a dynisco mini extruder
(LME) at a temperature above the melting or softening
point of the polystyrene, above about 180 °C. The
extrudate was collected in a water bath, dried and sub-
sequently ground to granules of different sieve sizes,
ranging from 0.5 mm to about 3 mm.
2.3 Scanning Probe Microscopy
The morphology of the neat PS and the prepared
PS/clay composites was visualized using a scanning
probe microscope (SPM); solver next (NT-MDT).
2.4 Adsorption Isotherm
Stock lead solutions with a concentration of 1000
μg/g were diluted to the required concentrations
(50–1300 μg/g), and then used in the equilibrium
experiments. After that, the initial concentration of the
lead ion samples and the final concentration for these
samples were diluted and analyzed by atomic absorption
spectroscopy (model AAnalyst 700, PerkinElmer). The
adsorption isotherm experiments were performed by
placing a constant mass of polymer composite (1 g) in 50
ml of lead-ion solution in glass bottles in a constant
agitation shaker. In each isotherm run, the temperature
used was 20 °C and the particle size was 0.5 mm. The
adsorption process reached equilibrium after 30 min;
however, the equilibrium experiments were conducted
for 3 hours to insure that the equilibrium state was
attained. The samples were then filtered using filter
papers, diluted, and the concentrations were measured.
The adsorbed concentration was calculated from the
mass-balance equation on the batch absorber as follows:
qe = V(Co – Ce)/M (1)
where M indicates the adsorbent mass, V is the solution
volume, qe is the adsorbed lead-ions concentration in
mg/g, Co is the initial concentration of lead ions and Ce
is the lead ions concentration in the bulk solution at
equilibrium. The amount of lead ions adsorbed on the
adsorbent versus the lead ions equilibrium concentration
in the solution can be plotted to obtain the equilibrium
adsorption isotherm curve and the maximum capacity of
the adsorption of lead ions by the polymer composites.
The spent adsorbents, polymer/clay composites, may be
regenerated by treating with strong acids such as
hydrochloric acid (HCl).
3 RESULTS AND DISCUSSION
The amount of lead adsorbed on polystyrene/clay
composites, virgin and recycled, versus the lead
concentration in the solution was plotted to obtain the
equilibrium adsorption isotherm curves for the two
composite systems. This is shown in Figures 1 and 2. As
a first observation we can see from Figures 1 and 2 that
the adsorption capacities for such composite systems are
low. Here we have polystyrene as the major phase that is
a non-polar polymer with no functional groups on the
surface and hence its activity towards lead ions
adsorption may be attributed to other factors other than
its surface functionality. Surface morphologies in the
form of the rough topology of the polystyrene granules
used in the adsorption isotherm experiments may be
responsible for such performance. Another explanation
for the adsorption of lead ions on the polystyrene is that,
the lead ions may be adsorbed by the polystyrene with
van der Waals forces because of the adsorption of the
lead ions on the polymer is of the Langmuir monolayer
type. The same trend was observed by Zhang,7 where he
attributed the adsorption to the monolayer adsorption in
addition to the capillary condensation and micropore
filling.
F. D. ALSEWAILEM, S. A. ALJLIL: RECYCLED POLYMER/CLAY COMPOSITES FOR HEAVY-METALS ADSORPTION
526 Materiali in tehnologije / Materials and technology 47 (2013) 4, 525–529
It is clear from Figures 1 and 2 that composites
based on recycled PS have a good adsorption ability in
comparison with those of the virgin composites. For
example, the amount of lead adsorbed per gram of
adsorbent was tripled when using composites containing
only 5 % of clay and 95 % of recycled PS in comparison
with those containing virgin PS. The adsorption behavior
of the neat recycled and virgin PS is also shown in
Figures 1 and 2 in order to compare their adsorption
ability with those reinforced with clay particles.
Needless to say that the adsorption capacity of neat clay,
that is Tabuk clay in this study, is much higher than that
of the composite system.6
The maximum capacity for the adsorption of lead
ions on the recycled polymer composite was on recycled
polymer-10, as shown in Figure 1. It is clear that
recycled polymer-10 was more efficient in terms of lead
adsorption when compared with that of the other types of
recycled polymers (recycled polymer-5 and recycled
polymer-7). The polar groups (such as carbonyl and
hydroxyl) that may exist on the surfaces of a recycled
polymer from exposure to the sun may enhance the
adsorption of metals ions. The surface of a recycled
polymer material may play an important role in
adsorption of the adsorbate, where it has many different
polar groups, such as carbonyl and hydroxyl.8 Figure 3
shows that the recycled composites have a distinct peak
of IR at a wave number of between 1700 and 1800,
which is an indication of the presence of a carbonyl
group. In addition, the mixed clay plays another role in
enhancing the adsorption of the lead ions on the recycled
polymer composite.
In contrast to what is seen with the recycled
composites, the neat virgin PS behaves better in terms of
adsorption capacity in comparison with that of the virgin
composites, especially at a low clay content, from 5–7
%. This may be attributed to diminishing the role of the
clay particles in the absence of a functional group on the
PS surface. It should also be noted that, as shown in
Figure 2, the maximum capacity of the adsorption of
lead ions on the virgin polymer composite, was less than
that of the recycled composites. Figure 4 shows the
appearance of the neat PS and the composite surfaces
taken by the SPM. It is clear that while the neat PS
surface appears relatively smooth, the clay particles were
not dispersed well within the virgin PS matrix, as shown
in Figure 4b. In contrast, the clay particles were
dispersed evenly in the recycled PS matrix, as shown in
Figure 4c. This may be due to the help of the carbonyl
group on the recycled PS surface.
Two isotherm models – Langmuir and Freundlich –
were applied to describe the experimental data for the
polymer composites used in this study.
The Langmuir model suggests that the adsorption of
heavy-metal ions on the adsorbent surface is a mono-
layer and this is applied to evaluate the maximum
capacity of the adsorption.7 The Langmuir isotherm
model is written as follows:
F. D. ALSEWAILEM, S. A. ALJLIL: RECYCLED POLYMER/CLAY COMPOSITES FOR HEAVY-METALS ADSORPTION
Materiali in tehnologije / Materials and technology 47 (2013) 4, 525–529 527
Figure 3: FTIR comparison of virgin and recycled PS
Slika 3: FTIR-primerjava devi{kega in recikliranega PS
Figure 2: Equilibrium isotherm experiments for the adsorption of lead
ions on the virgin polymer/clay composites with different amounts of
clay in mass fractions: 5 %, 7 % and 10 %
Slika 2: Ravnote`ne izoterme za adsorpcijo ionov svinca v devi{kem
kompozitu polimer-glina z razli~no vsebnostjo gline v masnih dele`ih:
5 %, 7 % in 10 %
Figure 1: Equilibrium isotherm experiments for the adsorption of lead
ions on the recycled polymer/clay composites with different amounts
of clay in mass fractions: 5 %, 7 % and 10 %
Slika 1: Ravnote`ne izoterme za adsorpcijo ionov svinca na reciklira-
nem kompozitu polimer-glina z razli~no vsebnostjo gline v masnih
dele`ih: 5 %, 7 % in 10 %
q
KC
bCe
e
e
=
+1
(2)
where k and b are constants. The mathematical forms of
the Langmuir model in the linear form is given below:
Ce/qe = 1/K + (b/K)Ce (3)
The Langmuir parameters K and b may be obtained
by using the linear regression technique with equation 3.
Figures 5 and 6 show the linear relationship between
Ce/qe and Ce for the polymer composites, recycled and
virgin. The equilibrium parameters K and b are listed in
Tables 1 and 2.
The Freundlich isotherm model may be used to
describe the adsorption isotherm of heterogeneous
adsorption surfaces. The model is given by the following
relation:
qe = Kf Ce
(1/n) (4)
The model, given by eq. 4, may be written in linear
form as follows:
lg qe = lg Kf + (1/n) lg Ce (5)
The equilibrium constants Kf and n may be calculated
using the linear regression technique with equation 5.
F. D. ALSEWAILEM, S. A. ALJLIL: RECYCLED POLYMER/CLAY COMPOSITES FOR HEAVY-METALS ADSORPTION
528 Materiali in tehnologije / Materials and technology 47 (2013) 4, 525–529
Figure 4: SPM graphs of PS/clay composites: a) neat PS; b) virgin
composites; c) recycled composites
Slika 4: SPM-posnetki kompozita polimer-glina: a) gladek PS; b) de-
vi{ki kompozit; c) recikliran kompozit
Table 1: Langmuir constants for the lead-ions adsorption on the
recycled polymer/clay composites
Tabela 1: Langmuirova konstanta za adsorpcijo ionov svinca na reci-
kliranem kompozitu polimer-glina
Material K/(l/g) b/(l/mg) R2
Recycled composite-5 0.1845 0.0316 0.9788
Recycled composite-7 7.880 1.1844 0.9833
Recycled composite-10 0.1629 0.0198 0.9862
Figure 6: Langmuir isotherm fit for the experimental data of the
virgin polymer composites
Slika 6: Ujemanje Langmuirove izoterme z eksperimentalnimi podat-
ki za kompozite iz devi{kega polimera
Figure 5: Langmuir isotherm fit for the experimental data of recycled
polymer composites
Slika 5: Ujemanje Langmuirove izoterme z eksperimentalnimi po-
datki za kompozite iz recikliranega polimera
Table 2: Langmuir constants for the lead-ions adsorption on the virgin
polymer/clay composites
Tabela 2: Langmuirova konstanta za adsorpcijo ionov svinca na de-
vi{kem kompozitu polimer-glina
Material K/(l/g) b/(l/mg) R2
Virgin composite-5 0.005 0.0020 0.9838
Virgin composite-7 0.008 0.0019 0.8828
Virgin composite-10 0.023 0.0060 0.9988
When the value of n is greater than one, this may
indicate that the adsorption of lead ions by the polymer
is favorable.9 Figures 7 and 8 demonstrate the linear
relationship between lg qe and lg Ce for the polymer com-
posite systems employed in this study. The equilibrium
constants, Kf and n, are presented in Tables 3 and 4.
Table 3: Freundlich constants for the lead-ions adsorption on the
recycled polymer/clay composites
Tabela 3: Freundlichova konstanta za adsorpcijo ionov svinca na
recikliranem kompozitu polimer-glina
Material Kf/(l/g) n R2
Recycled composite-5 3.77 16.584 0.1679
Recycled composite-7 5.59 35.587 0.0341
Recycled composite-10 3.39 7.855 0.5564
Table 4: Freundlich constants for the lead-ions adsorption on the
virgin polymer/clay composites
Tabela 4: Freundlichova konstanta za adsorpcijo ionov svinca na
devi{kem kompozitu polimer-glina
Material Kf/(l/g) n R2
Virgin composite-5 0.0245 1.641 0.9712
Virgin composite-7 0.0319 1.528 0.8787
Virgin composite-10 0.5869 3.959 0.9558
The data represented by Figures 5 to 8 and Tables 1
to 4 show that both models describe the experimental
data well, except in the case of the recycled composites
data fitted to Freundlich model, as seen from Table 3,
where the values of R2 are extremely low.
4 CONCLUSIONS
The results from the adsorption isotherm showed that
the recycled polymer composites were more efficient
than the virgin polymer composites. This may be
attributed to the functional groups, such as the carbonyl
that may exist on the surfaces of a recycled polymer, i.e.,
the PS, from a prolonged exposure to sunlight. In
addition, the clay mixed with the recycled polymer plays
another role in enhancing the adsorption of the lead ions
on the recycled polymer composites. The presence of the
carbonyl group on PS surface may contribute to a better
dispersion of the clay particles in the PS matrix. The
results also showed the recycled polymer composites
have an acceptable adsorption capacity for the lead ions
compared to the virgin composites. This finding could be
an advantage when it comes to utilizing plastic waste for
waste water treatment.
The experimental data were fitted using Langmuir
and Freundlich models. Both models used were effective
in describing the experimental data, except for the data
for the recycled composites that was fitted to the Freund-
lich model.
Acknowledgments
The authors thank the King Abdulaziz City for
Science and Technology for providing the support for
this study.
5 REFERENCES
1 V. M. Nurchi, I. Villaescusa, Coordination Chemistry Reviews, 256
(2012), 212
2 A. Li, Q. Zhang, G. Zhang, J. Chen, Z. Fei, F. Liu, Chemosphere, 47
(2002), 981
3 Z. Y. Xu, Q. X. Zhang, C. L. Wu, L. S. Wang, Chemosphere, 35
(1997), 2269
4 Z. Y. Xu, Q. X. Zhang, J. L. Chen, L. S. Wang, G. K. Anderson,
Chemosphere, 38 (1999), 2003
5 B. C. Pan, Y. Xiong, Q. Su, A. M. Li, J. L. Chen, Q. Z. Zhang,
Chemosphere, 51 (2003), 953
6 S. A. Aljlil, F. D. Alsewailem, Applied Clay Science, 42 (2009), 671
7 Y. H. Zhang, Adsorption, Shanghai Scientific & Technological
Literature Publishing House, Shanghai 1989, 125
8 W. Ruixia, C. Jinlong, C. Lianlong, F. Zheng-hao, L. Ai-min, Z.
Quanxing, Adsorp. Sci. Technol., 22 (2004), 523
9 M. S. El-Geundi, Adsorption Science and Technology, 7 (1990), 114
F. D. ALSEWAILEM, S. A. ALJLIL: RECYCLED POLYMER/CLAY COMPOSITES FOR HEAVY-METALS ADSORPTION
Materiali in tehnologije / Materials and technology 47 (2013) 4, 525–529 529
Figure 8: Freundlich isotherm is used to fit the experimental data of
the virgin polymer composites
Slika 8: Ujemanje Freundlichove izoterme z eksperimentalnimi po-
datki za kompozite iz devi{kega polimera
Figure 7: Freundlich isotherm fit for the experimental data of the
recycled polymer composites
Slika 7: Ujemanje Freundlichove izoterme z eksperimentalnimi po-
datki za kompozite iz recikliranega polimera

F. CAJNER et al.: INFLUENTIAL FACTORS IN THE SURFACE-HARDNESS TESTING OF A NITRIDED LAYER
INFLUENTIAL FACTORS IN THE SURFACE-HARDNESS
TESTING OF A NITRIDED LAYER
VPLIVNI FAKTORJI PRI PREIZKU[ANJU TRDOTE POVR[INE
NITRIRANE PLASTI
Franjo Cajner, Darko Landek, Ivan Kumi}, Amalija Vugrin~i}
Faculty of Mechanical Engineering and Naval Architecture (FMENA), University of Zagreb, Ivana Lucica Street No. 1, 10000 Zagreb, Croatia
darko.landek@fsb.hr
Prejem rokopisa – received: 2012-09-27; sprejem za objavo – accepted for publication: 2012-12-04
Nitrocarburizing is one of the frequently applied processes that significantly improve the service life of steel parts in the
complex activities of mechanical loads, wear and corrosion damages. Characterization and confirmation of the quality of a
nitrided/nitrocarburized layer is a prescribed norm including a determination of the surface hardness, nitriding hardening depth,
compound-zone thickness and its porosity. In the testing of the surface hardness, in spite of determined conditions, there are
additional factors that can affect the obtained result and can lead to a misunderstanding between a customer and a provider of
the service of nitriding. In this paper, possible influential factors related to the surface hardness are considered and statistically
analyzed by the ANOVA two-factor test. The considered influential factors are: the surface-preparation method and the loading
force. The tests carried out during the research indicated a significant influence of the loading force and of the surface
preparation on the results of the nitrided/nitrocarburized-steel surface hardness. Also, the indentation size effect is confirmed in
the hardness tests with a small loading force.
Keywords: nitriding/nitrocarburizing, surface hardness
Nitrocementacija je pogosto uporabljen postopek, ki ob~utno izbolj{a zdr`ljivost jeklenih komponent pri kompleksnem
delovanju mehanskih obremenitev, obrabe in korozijskih po{kodb. Karakterizacija in potrditev kvalitete nitrirane/
nitrocementirane plasti je predpisana in vklju~uje dolo~anje trdote povr{ine, globino utrjevanja z nitridi, debelino spojinske cone
in njeno poroznost. Pri preizku{anju trdote povr{ine so kljub to~no dolo~enim pogojem prisotni {e dodatni faktorji, ki lahko
vplivajo na rezultate in lahko povzro~ijo nesoglasje med uporabnikom in izvajalcem nitriranja. V tem ~lanku so mogo~i vplivni
faktorji obravnavani in statisti~no analizirani z ANOVA-preizkusom dveh faktorjev. Obravnavana faktorja sta: metoda priprave
povr{ine in sila obremenjevanja. Preizkusi so pokazali mo~an vpliv sile obremenjevanja in priprave povr{ine na rezultate trdote
nitrirane/nitrocementirane povr{ine. Pri merjenju trdote z majhnimi obremenitvami pa je bil ugotovljen tudi vpliv velikosti
odtisa.
Klju~ne besede: nitriranje/nitrocementiranje, trdota povr{ine
1 INTRODUCTION
Nitrocarburizing is a thermochemical process for
modifying the work-piece surface where carbon and
nitrogen are diffused into the surface to form a surface
layer consisting of a compound layer and a diffusion
layer.1 The ISO 15787:2001(E) standard prescribes
drawings and characterization of the nitrocarburized
layer. The characterization of the layer includes the
surface-hardness testing, determination of the effective
depth of nitriding, thickness of the compound layer and
its porosity.2 The above mentioned standard is supple-
mented with the ISO 6507-1:2005 standard, according to
which the hardness should be tested with the Vickers
hardness test method.3 However, even in the case when
all the requirements of these standards are fully satisfied,
the results of a nitrocarburized-layer characterization
may differ due to a number of influential factors. These
differences in the results may lead to a misunderstanding
between a customer and a provider of the service of
nitriding. Here, the inaccuracies and imprecision of a
surface-hardness testing are of special importance. The
results of a nitrocarburized-layer surface hardness testing
are significantly affected by the following factors:
indentation load of the indenter, accuracy and precision
of the hardness tester, preparation of the test-sample
surface, porosity of the compound layer, and the
measurer’s experience.4
2 EXPERIMENTS
In order to determine the effect of surface preparation
and of the applied load on the results, hardness tests
were carried out on the surface of the normalized
annealed 21CrMo5-7 steel after it had been nitrocarbu-
rized in a TENIFER salt bath (580 °C/2 h oil cooling).
Before the hardness tests, the surfaces of the nitrocar-
burized samples were prepared by grinding and polishing
according to Table 1 (the treatments marked as: A, B,
C). The aim of the polishing was to reduce or remove the
porous part of the compound layer in order to reduce the
adverse effect of the porosity on the hardness test results.
The hardness of the nitrocarburized layer was tested with
the Vickers hardness test method under three different
loads: 4.9 N, 9.81 N, and 49.03 N (HV0.5, HV1, and
HV5). The nitriding hardness depth (NHD) was deter-
mined on a cross-section of a metallographic test sample
Materiali in tehnologije / Materials and technology 47 (2013) 4, 531–533 531
UDK 621.785.53:519.233.4 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(4)531(2013)
with the Vickers test under the load of 4.9 N (HV0.5)
according to the ISO 15787:2001(E) standard. Optical
microscopy was applied to the same sample to determine
the thickness of the compound layer and its porosity.
Table 1: Preparation of the test sample surface for hardness testing
Tabela 1: Priprava preizkusne povr{ine vzorca za merjenje trdote
Sample Sandpaper grain size/duration (min)
A P1000/10
B P1000/10 + P2000/5
C P600/5 + P1000/5 + P2000/5
3 RESULTS
The surface-hardness test results for the nitrocarbu-
rized samples prepared with three different surface
treatments and with three different loads applied are
presented in Figure 1. The nitriding hardness depth
(Figure 2) is 0.22 mm. From the same figure one can see
that the hardness decreases sharply with an increased
distance from a sample edge. An analysis of variance
(ANOVA) was carried out to determine the significance
of the surface preparation and scale loading on the
results of the hardness tests conducted on a sample
surface. The ANOVA results (Figure 3) confirmed the
significant influence of the surface pre-treatment and the
selection of the indentation load of the indenter on the
results of hardness tests.
Generally, lower hardness values were measured
when a larger amount of the material was removed from
the surface of a nitrocarburized sample, i.e., when sand
paper with larger grains was used before the testing. The
nitrocarburized layer is thin and the hardness decreases
sharply with an increased distance from a sample edge.
Therefore, if the hardest part of the layer is removed by
grinding, the hardness of the surface is thus reduced.
Although samples A and B have been treated in a similar
way, sample B, which has been additionally treated with
F. CAJNER et al.: INFLUENTIAL FACTORS IN THE SURFACE-HARDNESS TESTING OF A NITRIDED LAYER
532 Materiali in tehnologije / Materials and technology 47 (2013) 4, 531–533
Figure 2: Hardness distribution on the cross-section of a nitrocar-
burized EN 21CrMo5-7 steel sample
Slika 2: Potek trdote preko pre~nega prereza nitrocementiranega
vzorca jekla EN 21CrMo5-7
Figure 1: Surface hardness of nitrocarburized 21CrMo5-7 steel
samples after different surface preparations (Table 1) and applied
loads
Slika 1: Trdota povr{ine nitrocementiranega vzorca iz jekla
21CrMo5-7 pri razli~nih pripravah povr{ine (Tabela 1) in uporablje-
nih obremenitvah
Figure 3: Influence of the surface preparation and test load on the
measured values of the surface hardness on the same nitrocarburized
EN 21CrMo5-7 steel sample: a) effect of the surface condition on the
measured values of surface hardness, b) effect of the indentation load
of the indenter on the measured values of surface hardness
Slika 3: Vpliv priprave povr{ine in obremenitve pri preizkusu na
izmerjeno vrednost trdote povr{ine na enakem nitrocementiranem
vzorcu jekla EN 21CrMo5-7: a) vpliv priprave povr{ine na izmerjeno
vrednost trdote povr{ine, b) vpliv obremenitve pri vtiskovanju
vtisnega telesa na izmerjeno vrednost trdote povr{ine
finer grade sandpaper, exhibited a better reflection of the
light, which made the measurement of diagonals easier,
while the hardness of the surface remained almost the
same. Accordingly, the surface preparation of sample B
could be considered optimal (as it shows a good
reflection of the light along with the preserved good
properties of the nitrocarburized layer).
If we consider the change in the surface hardness in
dependence on the indentation load of the indenter, we
can notice that higher values of hardness are obtained
when higher loads are applied. This is in line with the
expectations that the hardness measured under low loads
will depend greatly on the applied force due to the short
diagonals of indentations, which can hardly be measured
accurately. This phenomenon, known as the Indentation
Size Effect (ISE), is of significance at low loads.4 The
compound-layer thickness (CLT) of sample A was
approximately 16 μm, with a porous part of approxima-
tely 8 μm. The mean CST of sample B was approxima-
tely 15 μm and it was reduced to only 5 μm on the
damaged spots produced by rough grinding. If the
ground surfaces of samples A, B, and C are compared
using a light microscope, one can notice that a part of a
compound layer has been removed in the process of
surface preparation. The rougher surface of sample C
after the surface preparation (the surface is damaged,
more material has been removed) is the most probable
cause for the values of the surface hardness that are
lower than those of sample B.
4 CONCLUSION
The following conclusions can be drawn from the
investigation into the effect of a surface treatment and
indentation load of the indenter on the test results for the
nitrocarburized EN 21CrMo5-7 steel surface hardness:
• A nitrided layer with the compound-layer thickness
(CLT) of 16 μm and with the porosity of 8 μm was
obtained. The nitrided hardness depth (NHD) was
0.22 mm.
• The values of the surface hardness (according to the
Vickers test) greatly depend on the quality of the
surface and the applied load. By using sandpaper
with larger grains to grind the spot to be measured,
the porous part of the compound layer is removed,
the layer becomes thinner and the measured values of
the surface hardness are lower. When the loads of
0.981 N and 4.89 N were applied, higher values of
the surface hardness were obtained for the same
measurement spot with the load of 0.981 N than
those obtained in the test with the load of 0.489 N.
• During the preparation of the hardness testing of the
compound-layer porous part, a compromise between
grinding and polishing has to be made in order to
obtain a good light reflection and to preserve a
sufficient thickness of the surface layer.
Acknowledgements
This investigation was carried out within the research
projects "Surface engineering in manufacturing of
engineering components and tools" and "Modelling of
material properties and process parameters" funded by
the Ministry of Science, Education and Sports of the
Republic of Croatia, within the Faculty of Mechanical
Engineering and Naval Architecture, University of
Zagreb.
5 REFERENCES
1 D. Pye, Practical Nitriding and Ferritic Nitrocarburizing, ASM Inter-
national, Metals Park, OH, USA 2003
2 ISO 15787:2001(E) – Technical product documentation – Heat-
treated ferrous parts – Presentation and indication ISO Committee,
Geneve 2001
3 ISO 6507-1:2005 – Metallic materials – Vickers hardness test – Part
1: Test method, ISO Committee, Geneve 2005
4 ASM Handbook, Mechanical Testing and Evaluation, Vol. 8, ASM
International, Metals Park, OH 2000
F. CAJNER et al.: INFLUENTIAL FACTORS IN THE SURFACE-HARDNESS TESTING OF A NITRIDED LAYER
Materiali in tehnologije / Materials and technology 47 (2013) 4, 531–533 533

R. ZEJAK et al.: MECHANICAL AND MICROSTRUCTURAL PROPERTIES OF THE FLY-ASH-BASED ...
MECHANICAL AND MICROSTRUCTURAL PROPERTIES
OF THE FLY-ASH-BASED GEOPOLYMER PASTE AND
MORTAR
MEHANSKE IN MIKROSTRUKTURNE LASTNOSTI
GEOPOLIMERNIH VEZIV IN MALT IZ LETE^EGA PEPELA
Radomir Zejak1, Irena Nikoli}2, Dragoljub Ble~i}2, Vuk Radmilovi}3, Velimir Radmilovi}3
1University of Montenegro, Faculty of Civil Engineering, D`ord`a Va{ingtona bb, 81000 Podgorica, Montenegro
2University of Montenegro, Faculty of Metallurgy and Technology, D`ord`a Va{ingtona bb, 81000 Podgorica, Montenegro
3Faculty of Technology and Metallurgy, University of Belgrade, Karnegijeva 4, 11120 Belgrade, Serbia
irena@ac.me
Prejem rokopisa – received: 2012-12-03; sprejem za objavo – accepted for publication: 2012-12-14
In this paper we have investigated the influence of synthesis parameters on the mechanical properties of the fly-ash-based
geopolymer paste and mortars. Moreover, the influence of an addition of limestone sand on the microstructure of fly-ash-based
geopolymers was investigated as well. An addition of limestone sand increased the compressive strength of the fly-ash-based
geopolymer mortar, in comparison to the strength of the geopolymer paste, by changing the composition of the gel phase in the
geopolymer structure and with the physical strengthening due to the presence of well connected grains of the sand. The results
have shown that the compressive strength and Young’s modulus of elasticity of the fly-ash-based geopolymer mortar are
correlated. The values of these parameters increase as the alkali and silicate dosages increase. At one point, they reach the
maximum values and then start to decrease. On the other hand, the parameters that considerably increase the compressive
strength slightly decrease the flexural strength of the geopolymer mortar.
Keywords: geopolymerization, fly ash, mortar, flexural strength, compressive strength, Young’s modulus
^lanek obravnava raziskavo vplivov parametrov sinteze na mehanske lastnosti geopolimernih veziv in malt iz lete~ega pepela.
Preiskovan je bil tudi vpliv dodatka peska iz apnenca na mikrostrukturo geopolimera iz lete~ega pepela. Dodatek peska iz
apnenca je pove~al tla~no trdnost geopolimerne malte iz lete~ega pepela v primerjavi s trdnostjo geopolimernega veziva s
spreminjanjem sestave `elirne faze v strukturi geopolimera in s fizikalnim utrjevanjem zaradi prisotnosti dobro povezanih zrn
peska. Rezultati so pokazali pri geopolimeru iz lete~ega pepela povezavo med tla~no trdnostjo in Youngovim modulom
elasti~nosti. Vrednosti teh parametrov nara{~ajo z nara{~anjem dodatkov silikatov in alkalij. V dolo~eni to~ki dose`ejo najve~jo
vrednost in se potem za~nejo zmanj{evati. Po drugi strani pa parametri, ki ob~utno pove~ajo tla~no trdnost, rahlo zmanj{ajo
upogibno trdnost geopolimerne malte.
Klju~ne besede: geopolimerizacija, lete~i pepel, malta, upogibna trdnost, tla~na trdnost, Youngov modul
1 INTRODUCTION
Geopolymerization is an innovative technology that
can transform a variety of aluminosilicate materials with
alkali activation into useful, environmentally friendly
materials known as geopolymers (inorganic polymers).
These materials may be successfully applied in civil
engineering as a replacement for cement binders. Geo-
polymerization is also known as a waste-minimization
technology because it may utilize waste material (coal
fly ash and metallurgical slag) as raw material.
The mechanism of geopolymerisation may be
summarized in a few steps:
(1) Dissolution process, in which a breaking of
Al-O-Si and Si-O-Si bonds in the source material and a
liberation of Al3+ and Si4+ in an alkali silicate solution
occur. Together with the dissolution, a hydrolysis of
dissolved Al and Si occurs, leading to the formation of
aluminate and silicate monomeric species;
(2) Polycondensation of the Al and Si species in an
amorphous three-dimensional aluminosilicate network
(gel), in which the Si4+ and Al3+ cations are tetrahedrally
coordinated as [SiO4]4– and [AlO4]5– and linked by
oxygen bridges.1 The negative charge of the AlO4– group
is charge-balanced by alkali cations (Na+ and/or K+) 2;
(3) Hardening of aluminosilicate gel.
The empirical formula of a geopolymer is: Mn
[(-SiO2)z·AlO2]n·wH2O, where Mn is a cation, usually an
alkali (Na+ or K+), n is a degree of polycondensation, w £
3 and z is 1, 2 or 3.3
Geopolymers may be characterized with a variety of
properties like a good thermal resistance4–6 and resi-
stance to aggressive environment,7,8 but the basic engi-
neering request is a high compressive strength. Depend-
ing on the processing conditions, geopolymers may show
different values of compressive strength. Engineering
demands for an improvement in the mechanical pro-
perties of geopolymers have resulted in considering
possible geopolymer applications in the form of mortars
or concrete.9–13
Fly ash, the by-product from coal-fired power
stations, is mainly used for a geopolymer synthesis.
Although fly ash is not considered to be a hazardous
waste, the high quantity of ash produced all over the
Materiali in tehnologije / Materials and technology 47 (2013) 4, 535–540 535
UDK 678.86 ISSN 1580-2949
Professional article/Strokovni ~lanek MTAEC9, 47(4)535(2013)
world imposes the necessity of finding a solution for this
environmental problem. In this sense, a utilization of fly
ash through the geopolymerization process may be
considered as a promising solution.
The goal of this study was to investigate the influence
of the synthesis parameters, the alkali and silicate dosa-
ges, on the mechanical properties of the fly-ash-based
geopolymer paste and mortar and to investigate the
influence of a sand addition on the microstructural
properties of the fly-ash-based geopolymers.
2 EXPERIMENT
The fly ash sourced for this investigation was
supplied from the coal-fired power station Pljevlja in
Montenegro. Its chemical composition is given in Table
1.
Table 1: Chemical composition of fly ash
Tabela 1: Kemijska sestava lete~ega pepela
Compound Content, w/%
SiO2 49.45
Fe2O3 5.23
Al2O3 21.77
TiO2 0.66
CaO 13.34
Na2O 0.46
ZnO 4.5·10–3
MgO 1.29
MnO 0.02
P2O5 0.24
K2O 1.4
LOI* 4.35
*Loss on ignition
The geopolymer paste was prepared by mixing fly
ash with an alkali silicate solution, in a solid-to-liquid
ratio of 1. An alkali silicate activator was prepared by
mixing the NaOH and Na2SiO3 solutions at the mass
ratios of 1, 1.5 and 2. The concentrations of the NaOH
solution were (7, 10 and 13) M and a commercial sodium
silicate solution (w(Na2O) = 8.5 %, w(SiO2) = 28.5 %, a
density of 1.4 g/cm3) was used. The geopolymer paste
was cast in a cylindrical plastic mould, sealed with lead
and left to rest in an oven for 48 h at the temperature of
65 °C.
Geopolymer mortar was prepared by mixing the
geopolymer paste with crushed limestone sand. The
granulometry distribution obtained with the sieve method
is given in Table 2. The prepared mixtures were cast into
the standard 40 mm × 40 mm × 160 mm prism moulds
covered with polyethylene film sheets and left in an oven
for 48 h at 65 °C. After this time, the samples were
allowed to cool down, then removed from the moulds
and left to rest for additional 14 d at ambient temperature
before any testing was performed. The samples were
tested for the compressive and flexural strengths as well
as the modulus of elasticity. All the tests were performed
using a minimum of three samples per mortar mix.
Table 2: Analytical sieving analysis of limestone sand
Tabela 2: Analiti~na sejalna analiza peska iz apnenca
Sieve size opening
mm
Cumulative mass fraction of a
sample pass through each sieve
%
0.063 3.45
0.09 7.2
0.125 10.24
0.25 15.88
0.5 28.41
1.0 54.47
2.0 74.78
4.0 97.90
8.0 100.00
The samples of the geopolymer paste were tested
only for the compressive strength, while the samples of
the geopolymer mortar were subjected to testing for all
the mechanical properties (compressive strength, flexural
strength and Young’s modulus of elasticity).
The flexural strength was determined from the
modulus of the rupture test, while the Young’s modulus
of elasticity was calculated from the linear stress/strain
response prior to failure. The samples remaining from
the flexural-strength tests were used for testing the
compressive strength of the geopolymer mortar.
Microstructural investigations of the geopolymer
mortar were carried out using the FEI 235DB focused-
ion-beam system at the National Center for Electron
Microscopy at Berkeley, equipped with an EDAX
Genesis energy dispersive spectrometer (EDS). The
SEM images were recorded with a secondary electron
detector (SED). A cross-sectioning of the geopolymer
paste (without a sand addition) was carried out using
1000 pA gallium ion beam and a microstructural charac-
terization was performed using an electron beam at the 5
kV operating voltage. In order to minimize electron
charging effects, a through-lens back-scattered electron
detector (TLD-B) was used for image recording.
3 RESULTS AND DISCUSSION
3.1 Mechanical properties of fly-ash-based geopoly-
mers
Mechanical properties of fly-ash-based geopolymers
are of primary importance regardless of possible appli-
cations in civil engineering. One of the basic conditions
for a possible use of geopolymers in construction is a
satisfactory strength, which can be significantly
enhanced with an addition of sand. Two of the most
important factors that greatly influence mechanical
properties are alkali and silicate dosages. The influence
of an alkali dosage on the mechanical properties of the
geopolymer paste and mortar was evaluated through the
R. ZEJAK et al.: MECHANICAL AND MICROSTRUCTURAL PROPERTIES OF THE FLY-ASH-BASED ...
536 Materiali in tehnologije / Materials and technology 47 (2013) 4, 535–540
change in the NaOH concentration, while the influence
of a silicate dosage was investigated through the change
in the w(Na2SiO3)/w(NaOH) ratio.
The results have shown that the mechanical
properties of the fly-ash-based geopolymer paste and
mortar greatly depend on the reaction parameters of the
geopolymerization process. Depending on the reaction
parameters, the compressive strength of fly-ash-based
geopolymers may be considerably increased with an
addition of sand. The change in the compressive strength
of the fly-ash-based geopolymer paste and mortar as a
function of the reaction parameters is given in Figures 1
and 2. It is evident that the compressive strength of the
geopolymer mortar is considerably higher than the com-
pressive strength of the geopolymer paste. One of the
reasons for that is the fact that well connected grains of
sand physically strengthen the fly-ash-based geopoly-
mers. The compressive strengths of both the geopolymer
paste and mortar increase with the increase in the NaOH
concentration (Figure 1) and the w(Na2SiO3)/w(NaOH)
mass ratio (Figure 2). At one point, they reach the
maximum values and then decrease. The maximum
values of the compressive strength were obtained using
the 10 M NaOH at the w(Na2SiO3)/w(NaOH) mass ratio
of 1.5. Moreover, the compressive strength of the
geopolymer mortar is higher than the strength of the
geopolymer paste by 58.7 %.
The results of the investigation of geopolymer mor-
tars have shown that the Young’s modulus of elasticity is
related to the compressive strength. The increase in the
NaOH concentration from 7 M to 10 M leads to an
increase in the Young’s modulus. The maximum value of
the Young’s modulus of elasticity (9.1 GPa) was reached
using 10 M NaOH. A further increase in the alkaline
(NaOH) dosage to 13 M slightly reduced the modulus of
elasticity of the geopolymer mortar (Figure 3). The same
behaviour of the fly-ash-based geopolymer mortar was
observed with the change in the silicate dosage (Figure
4). The fly-ash-based geopolymer mortar reached the
maximum value of the Young’s modulus at the
w(Na2SiO3)/w(NaOH) ratio of 1.5. A further increase in
the silicate dosage, i.e., a change in the w(Na2SiO3)/
w(NaOH) ratio, led to a reduction in the Young’s modu-
lus of elasticity. A similar observation of the influence of
the silicate dosage on the Young’s modulus of elasticity
was observed by Duxon et. al.14
On the other hand, the flexural strength of the
fly-ash-based geopolymer composites is inversely related
to the compressive strength and Young’s modulus of
elasticity. The parameters that considerably increase the
compressive strength and Young’s modulus of elasticity
slightly reduce the flexural strength of the fly-ash-based
geopolymer composites. The minimum values of the
flexural strength correspond to the maximum values of
the compressive strength and Young’s modulus of
R. ZEJAK et al.: MECHANICAL AND MICROSTRUCTURAL PROPERTIES OF THE FLY-ASH-BASED ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 535–540 537
Figure 2: Change in the compressive strength of geopolymer paste
and mortar as a function of the w(Na2SiO3)/w(NaOH) mass ratio
Slika 2: Spreminjanje tla~ne trdnosti geopolimernega veziva in malte
v odvisnosti od masnega razmerja w(Na2SiO3)/w(NaOH)
Figure 1: Change in the compressive strength of geopolymer paste
and mortar as a function of the NaOH concentration
Slika 1: Spreminjanje tla~ne trdnosti geopolimernega veziva in malte
v odvisnosti od koncentracije NaOH
Figure 3: Change in the flexural strength and Young’s modulus of ela-
sticity of geopolymer mortar as a function of the NaOH concentration
Slika 3: Spreminjanje upogibne trdnosti in Youngovega modula ela-
sti~nosti geopolimerne malte v odvisnosti od koncentracije NaOH
elasticity that were reached using 10 M NaOH (Figure
3), at the w(Na2SiO3)/w(NaOH) ratio of 1.5 (Figure 4).
The highest value of the flexural strength was obtained
using 7 M NaOH at the w(Na2SiO3)/w(NaOH) mass ratio
of 1.
3.2 SEM –EDS microanalysis
The microstructure of the geopolymer mortar is
characterized by the presence of sand grains and geo-
polymer paste (Figure 5a). The geopolymer paste
consists of a gel phase (A) and unreacted fly-ash parti-
cles (B), (Figure 5b). The strength of the geopolymer
mortar primarily depends on the structure of the gel,
while the presence of the well connected grains of sand
additionally strengthen the structure of the mortar.
Besides, the unreacted fly-ash particles play the role of
microaggregates contributing to the overall strength of
the geopolymer mortar. The results of an EDS analysis
of the gel phase in the mortar (Figure 6a) have shown a
quantity of Ca that is considerably higher than the Si, Al,
and Na contents. On the other hand, the results of an
EDS analysis of the gel phase of the geopolymer paste
prepared without an addition of sand, but with the same
reaction parameters, have shown the quantities of Si, Al,
and Na that are considerably higher than the Ca content.
R. ZEJAK et al.: MECHANICAL AND MICROSTRUCTURAL PROPERTIES OF THE FLY-ASH-BASED ...
538 Materiali in tehnologije / Materials and technology 47 (2013) 4, 535–540
Figure 6: EDS of the: a) gel phase of fly-ash-based geopolymer
mortar and b) paste (in the absence of sand)
Slika 6: EDS veziva: a) geopolimerne malte iz lete~ega pepela in
b) veziva (brez peska)
Figure 5: a) Microstructure of geopolymer mortar prepared using 10
M NaOH at the w(Na2SiO3)/w(NaOH) mass ratio of 1.5; b) geopoly-
mer paste (as a part of geopolymer mortar); A – gel; B – unreacted fly
ash particles
Slika 5: a) Mikrostruktura geopolimerne malte, pripravljene z 10 M
NaOH, pri masnem razmerju w(Na2SiO3)/w(NaOH) 1,5; b) geopoli-
merno vezivo (kot del geopolimerne malte); A- vezivo; B- nereagirani
delci lete~ega pepela
Figure 4: Change in the flexural strength and Young’s modulus of
elasticity of geopolymer mortar as a function of the w(Na2SiO3)/
w(NaOH) mass ratio
Slika 4: Spreminjanje upogibne trdnosti in Youngovega modula
elasti~nosti geopolimerne malte v odvisnosti od masnega razmerja
w(Na2SiO3)/w(NaOH)
In this case, Ca in the gel phase only originates from the
fly ash (Figure 6b).
The previous research showed that the main gel
forming elements are Si, Al and Na and that the structure
of the geopolymer gel is determined by the w(Si)/w(Al)
ratio.15 But, in the geopolymer mixture with a high Ca
content, the influence of Ca must be taken into account
because the properties of the gel also greatly depend on
the Ca content.16,17
Table 3: Content of elements (w/%) and their ratios in the gel phases
of geopolymer mortar and paste
Tabela 3: Vsebnost elementov (w/%) in njihov dele` v vezivu geo-
polimerne malte in lepila
Ratio Geopolymer mortar Geopolymer paste
w(Na)/w(Si) 0.48 0.42
w(Na)/w(Al) 1.23 1.17
w(Si)/w(Al) 2.65 2.8
w(Ca)/w(Si) 3.04 0.17
Si 6.31 17.14
Al 2.38 6.11
Ca 19.17 2.85
There is no considerable difference between the
w(Na)/w(Si), w(Na)/w(Al) and w(Si)/w(Al) ratios in the
gel phase of the geopolymer paste and mortar (Table 3).
But, in the case of w(Ca)/w(Si), the ratio is considerably
higher in the gel phase of the geopolymer mortar than in
the geopolymer paste, which indicates a strong influence
of Ca on the structure of the geopolymer mortar. On the
other hand, in the gel phase of the geopolymer mortar
the observed quantities of Al and Si are considerably
lower than their contents in the gel phase of the geopoly-
mer paste without an addition of limestone sand (Table
3). It could be that the presence of Ca in some way
depletes the Al and Si dissolution, but this assumption
needs to be further studied.
As the limestone sand was used for the geopoly-
mer-mortar synthesis, it is necessary to take into account
the presence of Ca in the gel phase as a result of the
limestone dissolution in a strongly alkaline medium.
Besides, a certain amount of Ca in the gel originates
from fly ash. The reaction mechanism of the geopolyme-
risation varies significantly, depending on the presence
of Ca in the geopolymer mixture. In the absence of Ca,
the main product of the geopolymerisation process is an
aluminosilicate inorganic polymer (the gel), while in the
presence of Ca, additional Ca-bearing phases are
formed.18 The reaction pathway during the geopoly-
merization is strongly influenced by the Ca presence,
because it strongly reacts with the soluble silicate and
aluminate species provided during the dissolution step.
This leads to the formation of a calcium silica hydrate
and calcium aluminate hydrate gel phases.19 It is
considered that these Ca phases additionally strengthen
the geopolymer structure, but a clear understanding of
how Ca influences the reaction mechanism of the
geopolymerization has not yet been established. It is
assumed that the calcium silica hydrate gel fills the voids
and pores within the geopolymer paste leading to an
increase in the compressive strength.18 Moreover, it is
also considered that Ca is capable of acting as a
charge-balancing cation within the geopolymer structure
participating, in this way, in the geopolymerization
process.20
4 CONCLUSIONS
In this paper, the influence of the synthesis para-
meters and a limestone-sand addition on the mechanical
and microstructural properties of the fly-ash-based
geopolymer paste and mortar was investigated. The
results have shown that the compressive strength of the
geopolymer paste may be increased by adding limestone
sand. Moreover, the compressive strength correlates with
the Young’s modulus of elasticity of the fly-ash-based
geopolymer mortars. Both the compressive strength and
Young’s modulus of elasticity increase with an increase
in the NaOH concentration and the w(Na2SiO3)/
w(NaOH) mass ratio, reaching their maximum values at
one point and then decreasing. On the other hand, the
parameters that increase the compressive strength
slightly decrease the flexural strength of the fly-ash-
based geopolymer mortar. The maximum value of the
compressive strength corresponds to the minimum value
of the flexural strength of the fly-ash-based geopolymer
mortar, indicating that the fly-ash-based geopolymers
may be considered to be brittle materials. An addition of
limestone sand increases the compressive strength of the
fly-ash-based geopolymers by changing the structure of
the geopolymeric gel phase. The gel phase of the
geopolymer mortar is characterized by a content of Ca
that is considerably higher than in the geopolymer paste,
resulting in a higher compressive strength of the
geopolymer mortar. In addition, the physical presence of
the well connected grains of sand also contributes to the
strengthening of the geopolymer mortar.
Acknowledgements
The authors would like to acknowledge the financial
support of the Montenegrin Ministry of Science in the
framework of the project No.01-460. The FIB-SEM
analysis was performed at the National Center for
Electron Microscopy at Lawrence Berkeley National
Laboratory, funded by the U.S. Department of Energy
under the Contract DE-AC02-05CH11231.
5 REFERENCES
1 P. Duxson, A. Fernández-Jiménez, J. L. Provis, G. C. Lukey, A.
Palomo, J. S. J. van Deventer, Geopolymer technology: the current
state of the art, Journal of Materials Science, 42 (2007) 9, 2917–2933
2 F. Pacheco-Torgal, J. Castro-Gomes, S. Jalali, Alkali-activated bin-
ders: A review: Part 1. Historical background, terminology, reaction
mechanisms and hydration products, Construction and Building
Materials, 22 (2008), 1305–1314
R. ZEJAK et al.: MECHANICAL AND MICROSTRUCTURAL PROPERTIES OF THE FLY-ASH-BASED ...
Materiali in tehnologije / Materials and technology 47 (2013) 4, 535–540 539
3 J. Davidovits, Geopolymers, inorganic polymeric new materials,
Journal of Thermal Analysis, 37 (1991), 1633–1656
4 D. L. Y. Kong, J. G. Sanjayan, Effect of elevated temperatures on
geopolymer paste, mortar and concrete, Cement and Concrete
Research, 40 (2010), 334–339
5 D. L. Y. Kong, J. G. Sanjayan, K. Sagoe-Crentsil, Comparative
performance of geopolymers made with metakaolin and fly ash after
exposure to elevated temperatures, Cement and Concrete Research,
37 (2007), 1583–1589
6 D. L.Y. Kong, J. G. Sanjayan, Damage behavior of geopolymer com-
posites exposed to elevated temperatures, Cement and Concrete
Composites, 30 (2008), 986–991
7 V. Sata, A. Sathonsaowaphak, P. Chindaprasirt, Resistance of lignite
bottom ash geopolymer mortar to sulfate and sulfuric acid attack,
Cement and Concrete Composites, 34 (2012), 700–708
8 R. R. Lloyd, J. L. Provis, J. S. J. van Deventer, Acid resistance of
inorganic polymer binders. 1. Corrosion rate, Materials and Struc-
tures, 45 (2012), 1–14
9 M. Steinerova, Mechanical properties of geopolymer mortars in
relation to their porous structure, Ceramics –Silikaty, 55 (2011) 4,
362–372
10 S. Jumrat, B. Chatveera, P. Rattanadecho, Dielectric properties and
temperature profile of fly ash-based geopolymer mortar, Inter-
national communications in Heat and Mass Transfer, 38 (2011),
242–248
11 J. Temuujin, A. van Riessen, K. J. D. MacKenzie, Preparation and
characterization of fly ash based geopolymer mortars, Construction
and Building Materials, 24 (2010), 1906–1910
12 P. K. Sarker, R. Haque, K. V. Ramgolam, Fracture behaviour of heat
cured fly ash based geopolymer concrete, Materials and Design, 44
(2013), 580–586
13 X. S. Shia, F. G. Collins, X. L. Zhaob, Q. Y. Wanga, Mechanical
properties and microstructure analysis of fly ash geopolymeric
recycled concrete, Journal of Hazardous Materials, 237–238 (2012),
20–29
14 P. Duxson, J. L. Provis, G. C. Lukey, S. W. Mallicoat, W. M. Kriven,
J. S. J. van Deventer, Understanding the relationship between geo-
polymer composition, microstructure and mechanical properties,
Colloids and Surfaces A: Physicochemical and Engineering Aspects,
269 (2005), 47–58
15 H. Xu, J. S. J. van Deventer, Effect of Source Materials on Geopoly-
merization, Industrial and Engineering Chemistry Research, 42
(2003), 1698–1706
16 J. Temuujin, A. van Riessen, R. Williams, Influence of calcium com-
pounds on the mechanical properties of fly ash geopolymer pastes,
Journal of Hazardous Materials, 167 (2009), 82–88
17 K. Dombrowski, A. Buchwald, M. Weil, The influence of calcium
content on the structure and thermal performance of fly ash based
geopolymers, Journal of Materials Science, 42 (2007), 3033–3043
18 D. Khale, R. Chaudhary, Mechanism of geopolymerization and
factors influencing its development: a review, Journal of Materials
Science, 42 (2007), 729–746
19 P. Duxson A. Fernández-Jiménez, J. L. Provis, G. C. Lukey, A.
Palomo, J. S. J. van Deventer, Geopolymer technology: the current
state of the art, Journal of Materials Science, 42 (2007), 2917–2933
20 L. Chao, S. Henghu, L. Longtu, A review: The comparison between
alkali-activated slag (Si+Ca) and metakaolin (Si+Al) cements,
Cement and Concrete Research, 40 (2010), 1341–1349
R. ZEJAK et al.: MECHANICAL AND MICROSTRUCTURAL PROPERTIES OF THE FLY-ASH-BASED ...
540 Materiali in tehnologije / Materials and technology 47 (2013) 4, 535–540