A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
903–910
PHASE-TRANSFORMATION BEHAVIOR AND
MICROMECHANICAL PROPERTIES OF A DUAL-PHASE STEEL
AFTER CHEMICAL MODIFICATIONS
FAZNE SPREMEMBE IN MIKROMEHANSKE LASTNOSTI
DVOFAZNIH JEKEL PO KEMIJSKIH PRILAGODITVAH
Aijuan Zhao1, Guoxin Zhao2, Haijie Sun1, Hairong Gao1, Shaoqing Wang1,
Xiuli Chen1
1School of Chemistry and Chemical Engineering, Zhengzhou Normal University, Zhengzhou 450044, Henan, China
2Institute of Chemical Industry and Food, Zhengzhou Institute of Technology, Zhengzhou 450044, Henan, China
mkyj2017@163.com
Prejem rokopisa – received: 2017-01-31; sprejem za objavo – accepted for publication: 2017-05-30
doi:10.17222/mit.2017.017
In this work, the phase-transformation behavior and micromechanical properties of a dual-phase steel after chemical
modifications were investigated theoretically and experimentally. In particular, the micromechanical behavior of the steel was
modeled, based on the effects of the microstructure, phase fractions, local compositions of single phases and their area shapes.
The developed model was used for predicting the damage behavior of a specimen. It was demonstrated that the tensile strength
increased with the increasing temperature due to an increase in the amount of martensite in the steel, but the hardening behavior
of this specimen was affected by the microstructure. Furthermore, the flow curves of the steel under different intercritical
temperatures could be well predicted based on the real microstructures. The subsequent simulation results showed that while
higher stress concentrated on the martensite, the shear-band appearance strongly depended on the microstructures of the phases.
In addition, for the prediction of damage behaviors, the true stress/true strain curves of macroscale simulations showed good
agreement with the experiments involving differently heat-treated steels.
Keywords: intercritical treatment, steel, micro model
V tem delu so teoreti~no in eksperimentalno preiskovali fazne spremembe in mikromehanske lastnosti dvofaznega jekla po
kemijskih prilagoditvah. [e posebej je bilo modelirano mikromehansko obna{anje jekla glede na mikrostrukturo, dele`e
posameznih faz, lokalno sestavo posameznih faz in njihovo morfologijo. Razviti model je bil uporabljen za napoved po{kodb
vzorca. Dokazano je bilo, da se natezna trdnost pove~uje z nara{~ajo~o temperaturo zaradi nara{~anja vsebnosti martenzita v
jeklu, vendar je na kaljivost tega vzorca vplivala mikrostruktura. Nadalje je bilo ugotovljeno, da je mo`no krivulje te~enja jekla
pri razli~nih interkriti~nih temperaturah dobro napovedati na podlagi dejanskih mikrostruktur. Nadalje so rezultati simulacij
pokazali, da so vi{je napetosti koncentrirane na martenzitu in isto~asna prisotnost stri`nih pasov, mo~no odvisne od
mikrostruktur posameznih faz. Poleg tega je za napoved po{kodb pomembno, da se prave krivulje napetost-deformacija,
dobljene s pomo~jo simulacij na makronivoju, dobro ujemajo z eksperimentalnimi, dobljenimi pri razli~no toplotno obdelanih
jeklih.
Klju~ne besede: interkriti~na obdelava, jeklo, mikromodel
1 INTRODUCTION
In the past few decades, HSLA steels were widely
used in the fields of pipelines, pressure vessels, heavy
machinery, buildings, cars, bridges, offshore platforms
and ships.1 However, with the development of the
technology and deterioration of the environment, an even
better performance of advanced steel was required,
especially for the automotive industry.2 As a result, DP
steels were developed. The matrix of DP steels consists
of two different phases: ferrite and martensite; the
former shows great plasticity and toughness, while the
later shows high strength. Combining the two great per-
formances, DP steels show great mechanical properties.3
The intercritical heat treatment, which involves
holding in the two-phase (+) region, followed by
quick cooling, is frequently applied to obtain ferrite and
martensite phases in a low-alloy steel.4–8 Previous
investigations suggested that different chemical compo-
sitions, thermomechanical processing routes or heat
treatments lead to different microstructural configu-
rations, different phase-space distributions and other
microstructure characteristics, affecting the deformation
and failure behaviors of DP steels.9,10
Recently, researchers obtained the flow curves of the
composite phases using micromechanical modelling
based on the RVEs selected from a real microstruc-
ture.11–19 The flow behavior of a DP steel mainly depends
on the properties of ferrite and martensite and the vol-
ume fractions of different phases. As steel has the same
chemical composition, when modelling heat-treated
steels, the model mainly focuses on the effects of
strengthening methods and the grain size on the strength
of different phases. With respect to material modeling,
two equations, the Ashby–Orowan equation and
Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910 903
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
UDK 669.017.3:67.017:669.018.2 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(6)903(2017)
Hall-Petch equation, were taken into account when
describing the flow-stress behavior of steel.12
Moreover, in actual engineering, the predictions of
deformation and failure behaviors of DP steels are the
final purpose. Some of the available damage models
such as the ones of the aforementioned failure mecha-
nisms, the Gurson–Tvergaard–Needleman (GTN) da-
mage model, the extended finite-element model
(XFEM), the cohesive-zone model and so on, were
closely estimated for the experiments.20–23
A. Ramazani et al. investigated the effect of an in-
homogeneous morphology on the mechanical properties
of a welded joint.15 2D RVEs simulated flow curves were
corrected to 3D ones, taking into account the effects of
the microstructure, the chemical composition and the
area fraction on the macroscopic mechanical properties
of the welded joint. Finally, the tensile test of the welded
material with the inhomogeneous morphology was
simulated and good agreement between the experimental
and predicted flow curves was achieved.
The aim of this work was to predict the damage
behavior of an HSLA steel under different intercritical-
temperature conditions and to establish a relational
model presenting the connection between the microstruc-
ture and macromechanics of materials to provide guid-
ance for material design based on the microstructure.
2 EXPERIMENTAL PART
The as-received material was the Gr.65 steel, a high-
strength low-alloy structural steel (with a yield strength
of 460MPa), supplied as a hot-roll plate. Table 1 gives
the chemical composition of the steel. The initial micro-
structure mainly consisted of ferrite with a small amount
of pearlite, shown in Figure 1.
All of the specimens cut along the rolling direction of
the plate, with dimensions of 70 mm × 40 mm × 4 mm
(length, width and thickness) underwent heat treatment
in a resistance furnace, and the heat-treatment regime is
illustrated on Figure 2. The specimens were first auste-
nitized at 900 °C for 15 min, followed by water cooling
to get the martensite phase, then reheated to different
intercritical temperatures (760, 790 and 820) °C for 15
min and water cooled again to get a ferrite-martensite
microstructure. Then, all the specimens were tempered at
450 °C for a holding time of 10 min and air cooled to
room temperature. After the heat treatments, the speci-
mens were ground, polished and etched for metallogra-
phic analyses.
Figure 3 shows the geometry of the specimens for
tensile testing. The tensile specimens were tested at a
strain rate of 0.00033 s–1. The force and displacement
curves were recorded by means of a load cell and exten-
someter (in this study, the length of the extensometer
was 12.5 mm) during the tests. Later, the stress/strain
curves of the steels were calculated at room temperature.
3 NUMERICAL APPROACH
3.1 Micromechanical modeling
Recently, the model based on the Ashby-Orowan
equation and Hall-Petch equation was usually used to
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
904 Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Table 1: Chemical composition of Gr.65 steel
Elements C Si Mn P S Al V Ti Cr Nb Fe
w/% 0.13 0.3 1.4 0.014 0.002 0.03 0.043 0.014 0.06 0.031 Balance
Figure 2: Heat-treatment regime for the samples
Figure 1: Original microstructure of Gr. 65 steel in OM Figure 3: Geometry of a specimen for tensile testing
predict the flow-stress behavior of each phase in the
steel.11–15,18 The former equation represents the material
strengthening due to the carbon content and other
alloying elements’ carbide precipitation, while the latter
is based on the dislocation theory determining the effects
of the grain size.12 The approach can be expressed as
Equation (1):
    

= + +
− −
0
1
Δ s M b
Mk
kL
exp( )
(1)
where  is the flow stress and  is the true strain. The
first term 0 takes care of the Peierls stress and the
effects of the elements in a solid solution (Equation (2)).
The second term is the solid-solution strengthening due
to the carbon content (Equations (3) and (4)).  is the
material constant, M is the Taylor factor, μ is the shear
modulus, b is the Burgers vector, k is the recovery rate
and L describes the dislocation mean free path.
0 (in MPa) = 77 + 750(%P) + 60(%Si) + 80(%Cu) +
+ 45(%Ni) + 60(%Cr) + 80(%Mn) + 11(%Mo) (2)
From the literature, the values for parameters , M, μ,
and b were 0.33, 3, 80000 MPa, and 2.5·10–10 m. For
ferrite, the k value used is 10–5/d, where d is the ferrite
gain size. For martensite, two constants, 41 and 0.038,
were used to define k and L.11,13,15,18
The second term is the solid-solution strengthening
due to the carbon and nitrogen contents. For ferrite, it is:
Δ s ss
f
ss
fC N= +5000(% % ) (3)
while for martensite, it is
Δ s ss
m
ss
mC N= + −3605 161(% % ) (4)
where Css
f and %N ss
f denote the carbon and nitrogen
contents in the solid solution (w/%) in ferrite, Css
m and
%N ss
m are the carbon and nitrogen contents in the solid
solution in martensite, respectively. The values of the
carbon and nitrogen contents were computed with
software JMatPro.
3.2 Macromechanical modeling
2D RVE was generated based on real micrographs
and can involve all the microstructural features in the
calculations. However, a specimen deforms three-dimen-
sionally during the uniaxial tensile test and 2D-modeling
approaches are not able to predict the flow curve of a
material precisely. Therefore, in order to study the
effects of microstructural features on the mechanical
properties of a heat-treated steel, the predicted flow
curves of 2D modeling should be correlated to the 3Ds
by introducing a correlation factor. A. Ramazani et al.11
introduced a function as 3D/2D to describe the correla-
tion between the 2D and 3D flow-curve modelling of DP
steels. The function mainly considered the effect of the
martensite volume fraction and the equivalent plastic
strains under the 2D FE simulation and provided some
influence exponentials, as shown in Equation (5). There-
fore, the developed flow curves from 2D RVE calcula-
tions were corrected to 3D curves using Equation (5).
These corrected flow curves are used as input data for
macromechanical modeling.
  

3D 2D eq
p
m m
eq
p
/ ( )
. (
= × × × + × +
+ ×
− −2 10 1 10
0 0218
4 2 3 7 3V V
) . ( )
. ( )
2 2 5 2 2
2
7 10 018
0 007
× + × × + × ×
× + × ×
−V V
V V
m m eq
p
m eq
p

 m eq
p
m+ × × +0 0036 1
2. ( ) V
(5)
where 3D and 2D are the 3D and 2D flow stresses, Vm
and  eq
p are the martensite volume fraction and
equivalent plastic strains under the 2D FE simulation.
In order to study the microstructure-based failure of
the intercritically heat-treated HSLA steel, a GTN model
was applied to investigate the damage behavior of the
material at the macroscale. The GTN model is formu-
lated as Equation (6):
	



GTN
v H=
⎛
⎝
⎜
⎞
⎠
⎟ + ⋅
⎛
⎝
⎜
⎞
⎠
⎟ − +
y y
q f q q f
2
1 2 3
22
3
2
1* *cosh ( ) = 0
(6)
where v, y and H are the von Mises equivalent stress,
the matrix-material yield stress and the hydrostatic
stress; q1, q2 and q3 are the model parameters, in prac-
tice, q1 = 1.5, q2 = 1 and q3 = (q1)2 = 2.25.24 Function f*
was introduced by V. Tvergaard and A. Needleman25 to
describe the effect of a void interaction starting during
the failure process (Equation (7)):
f
f f f
f
q f
f f
f f f f
*
;
/
( );
=
≤
+
−
−
− >
⎧
⎨
⎪
⎩⎪
c
c
c
F c
c c
1 1 (7)
were f, fc and fF are the void volume fraction, the critical
void volume fraction at the onset of void coalescence,
and the void volume fraction at failure, respectively. The
whole void-volume-fraction development is defined as
the sum of the growth of the existing voids: fgrowth and
the nucleation of new voids fnucleation .
As the matrix material is considered to be incom-
pressible, fgrowth is defined by the volumetric part of the
plastic-strain rate  kk :
 ( ) f fgrowth kk= − ⋅1  (8)
C. C. Chu and A. Needleman26 assumed that the
strain controlled the void-nucleation mechanism,
following a normal distribution. fnucleation is defined by the
rate of the equivalent plastic strain  in Equation (9):
 f Anucleation = ⋅  (9)
where
A
f
s S
= −
−⎛
⎝
⎜
⎞
⎠
⎟
⎡
⎣
⎢
⎤
⎦
⎥⋅
N
N
N
N2
1
2
2
π
exp
 
,
fn is the volume fraction of the secondary voids, N and
SN are the mean value and the standard deviation of the
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910 905
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Hall-Petch equation, were taken into account when
describing the flow-stress behavior of steel.12
Moreover, in actual engineering, the predictions of
deformation and failure behaviors of DP steels are the
final purpose. Some of the available damage models
such as the ones of the aforementioned failure mecha-
nisms, the Gurson–Tvergaard–Needleman (GTN) da-
mage model, the extended finite-element model
(XFEM), the cohesive-zone model and so on, were
closely estimated for the experiments.20–23
A. Ramazani et al. investigated the effect of an in-
homogeneous morphology on the mechanical properties
of a welded joint.15 2D RVEs simulated flow curves were
corrected to 3D ones, taking into account the effects of
the microstructure, the chemical composition and the
area fraction on the macroscopic mechanical properties
of the welded joint. Finally, the tensile test of the welded
material with the inhomogeneous morphology was
simulated and good agreement between the experimental
and predicted flow curves was achieved.
The aim of this work was to predict the damage
behavior of an HSLA steel under different intercritical-
temperature conditions and to establish a relational
model presenting the connection between the microstruc-
ture and macromechanics of materials to provide guid-
ance for material design based on the microstructure.
2 EXPERIMENTAL PART
The as-received material was the Gr.65 steel, a high-
strength low-alloy structural steel (with a yield strength
of 460MPa), supplied as a hot-roll plate. Table 1 gives
the chemical composition of the steel. The initial micro-
structure mainly consisted of ferrite with a small amount
of pearlite, shown in Figure 1.
All of the specimens cut along the rolling direction of
the plate, with dimensions of 70 mm × 40 mm × 4 mm
(length, width and thickness) underwent heat treatment
in a resistance furnace, and the heat-treatment regime is
illustrated on Figure 2. The specimens were first auste-
nitized at 900 °C for 15 min, followed by water cooling
to get the martensite phase, then reheated to different
intercritical temperatures (760, 790 and 820) °C for 15
min and water cooled again to get a ferrite-martensite
microstructure. Then, all the specimens were tempered at
450 °C for a holding time of 10 min and air cooled to
room temperature. After the heat treatments, the speci-
mens were ground, polished and etched for metallogra-
phic analyses.
Figure 3 shows the geometry of the specimens for
tensile testing. The tensile specimens were tested at a
strain rate of 0.00033 s–1. The force and displacement
curves were recorded by means of a load cell and exten-
someter (in this study, the length of the extensometer
was 12.5 mm) during the tests. Later, the stress/strain
curves of the steels were calculated at room temperature.
3 NUMERICAL APPROACH
3.1 Micromechanical modeling
Recently, the model based on the Ashby-Orowan
equation and Hall-Petch equation was usually used to
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
904 Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Table 1: Chemical composition of Gr.65 steel
Elements C Si Mn P S Al V Ti Cr Nb Fe
w/% 0.13 0.3 1.4 0.014 0.002 0.03 0.043 0.014 0.06 0.031 Balance
Figure 2: Heat-treatment regime for the samples
Figure 1: Original microstructure of Gr. 65 steel in OM Figure 3: Geometry of a specimen for tensile testing
predict the flow-stress behavior of each phase in the
steel.11–15,18 The former equation represents the material
strengthening due to the carbon content and other
alloying elements’ carbide precipitation, while the latter
is based on the dislocation theory determining the effects
of the grain size.12 The approach can be expressed as
Equation (1):
    

= + +
− −
0
1
Δ s M b
Mk
kL
exp( )
(1)
where  is the flow stress and  is the true strain. The
first term 0 takes care of the Peierls stress and the
effects of the elements in a solid solution (Equation (2)).
The second term is the solid-solution strengthening due
to the carbon content (Equations (3) and (4)).  is the
material constant, M is the Taylor factor, μ is the shear
modulus, b is the Burgers vector, k is the recovery rate
and L describes the dislocation mean free path.
0 (in MPa) = 77 + 750(%P) + 60(%Si) + 80(%Cu) +
+ 45(%Ni) + 60(%Cr) + 80(%Mn) + 11(%Mo) (2)
From the literature, the values for parameters , M, μ,
and b were 0.33, 3, 80000 MPa, and 2.5·10–10 m. For
ferrite, the k value used is 10–5/d, where d is the ferrite
gain size. For martensite, two constants, 41 and 0.038,
were used to define k and L.11,13,15,18
The second term is the solid-solution strengthening
due to the carbon and nitrogen contents. For ferrite, it is:
Δ s ss
f
ss
fC N= +5000(% % ) (3)
while for martensite, it is
Δ s ss
m
ss
mC N= + −3605 161(% % ) (4)
where Css
f and %N ss
f denote the carbon and nitrogen
contents in the solid solution (w/%) in ferrite, Css
m and
%N ss
m are the carbon and nitrogen contents in the solid
solution in martensite, respectively. The values of the
carbon and nitrogen contents were computed with
software JMatPro.
3.2 Macromechanical modeling
2D RVE was generated based on real micrographs
and can involve all the microstructural features in the
calculations. However, a specimen deforms three-dimen-
sionally during the uniaxial tensile test and 2D-modeling
approaches are not able to predict the flow curve of a
material precisely. Therefore, in order to study the
effects of microstructural features on the mechanical
properties of a heat-treated steel, the predicted flow
curves of 2D modeling should be correlated to the 3Ds
by introducing a correlation factor. A. Ramazani et al.11
introduced a function as 3D/2D to describe the correla-
tion between the 2D and 3D flow-curve modelling of DP
steels. The function mainly considered the effect of the
martensite volume fraction and the equivalent plastic
strains under the 2D FE simulation and provided some
influence exponentials, as shown in Equation (5). There-
fore, the developed flow curves from 2D RVE calcula-
tions were corrected to 3D curves using Equation (5).
These corrected flow curves are used as input data for
macromechanical modeling.
  

3D 2D eq
p
m m
eq
p
/ ( )
. (
= × × × + × +
+ ×
− −2 10 1 10
0 0218
4 2 3 7 3V V
) . ( )
. ( )
2 2 5 2 2
2
7 10 018
0 007
× + × × + × ×
× + × ×
−V V
V V
m m eq
p
m eq
p

 m eq
p
m+ × × +0 0036 1
2. ( ) V
(5)
where 3D and 2D are the 3D and 2D flow stresses, Vm
and  eq
p are the martensite volume fraction and
equivalent plastic strains under the 2D FE simulation.
In order to study the microstructure-based failure of
the intercritically heat-treated HSLA steel, a GTN model
was applied to investigate the damage behavior of the
material at the macroscale. The GTN model is formu-
lated as Equation (6):
	



GTN
v H=
⎛
⎝
⎜
⎞
⎠
⎟ + ⋅
⎛
⎝
⎜
⎞
⎠
⎟ − +
y y
q f q q f
2
1 2 3
22
3
2
1* *cosh ( ) = 0
(6)
where v, y and H are the von Mises equivalent stress,
the matrix-material yield stress and the hydrostatic
stress; q1, q2 and q3 are the model parameters, in prac-
tice, q1 = 1.5, q2 = 1 and q3 = (q1)2 = 2.25.24 Function f*
was introduced by V. Tvergaard and A. Needleman25 to
describe the effect of a void interaction starting during
the failure process (Equation (7)):
f
f f f
f
q f
f f
f f f f
*
;
/
( );
=
≤
+
−
−
− >
⎧
⎨
⎪
⎩⎪
c
c
c
F c
c c
1 1 (7)
were f, fc and fF are the void volume fraction, the critical
void volume fraction at the onset of void coalescence,
and the void volume fraction at failure, respectively. The
whole void-volume-fraction development is defined as
the sum of the growth of the existing voids: fgrowth and
the nucleation of new voids fnucleation .
As the matrix material is considered to be incom-
pressible, fgrowth is defined by the volumetric part of the
plastic-strain rate  kk :
 ( ) f fgrowth kk= − ⋅1  (8)
C. C. Chu and A. Needleman26 assumed that the
strain controlled the void-nucleation mechanism,
following a normal distribution. fnucleation is defined by the
rate of the equivalent plastic strain  in Equation (9):
 f Anucleation = ⋅  (9)
where
A
f
s S
= −
−⎛
⎝
⎜
⎞
⎠
⎟
⎡
⎣
⎢
⎤
⎦
⎥⋅
N
N
N
N2
1
2
2
π
exp
 
,
fn is the volume fraction of the secondary voids, N and
SN are the mean value and the standard deviation of the
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910 905
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
characteristic plastic-strain distribution,  and  are the
equivalent plastic strain and the rate of the equivalent
plastic strain, respectively.
Finally, the damage-evolution law is given as:
f f f f
f
s
= + = − ⋅ +
+ −
−
  ( ) 
exp
growth nucleation kk
N
N
1
2
1
2


π


N
NS
⎛
⎝
⎜
⎞
⎠
⎟
⎡
⎣
⎢
⎤
⎦
⎥⋅
2 (10)
In this damage model, N and SN were assumed to be
0.3 and 0.1, while the other parameters were determined
with a calibration of the numerical results using experi-
mental data. 24
3.3 Macro/microscopic finite-element model and boun-
dary conditions
All the associated numerical models were performed
with the finite-element software ABAQUS. Figure 4a
shows the macroscopic finite-element model for a ten-
sile-test specimen, and the specimen was modeled with
solid, reduced-integration element C3D8R. The boun-
dary condition and the load plan, according to which the
upper part of the specimen was stretched and the bottom
was fixed was consistent with the real condition.
In the case of the macroscopic finite-element simu-
lations, all the specimens were generally considered as
continuous and homogeneous. On the microscale, each
constituent character of the microstructure of a steel
specimen needed to be incorporated. RVE was applied to
take into account the influences of different deformation
behaviors of each phase on the resulting mechanical
properties. The 2D RVE models were generated from
real microstructures modeled with plane strain element
CPE4R in ABAQUS. The selection requirement is that
the volume percent of martensite is the same with the
real microstructure in 2D RVE. The boundary condition
and load plan are shown in Figure 4b; the bottom of 2D
RVE is fixed and the top of it is pulled.
4 RESULTS AND DISCUSSION
4.1 Microscale simulation results
Intercritical temperature T, ferrite grain size d,
martensite fraction Vm, along with the carbon content in
ferrite Css
f and in martensite Css
m are listed in Table 2.
After calculating the nitrogen content, we found that the
effect of the nitrogen content on the solid-solution
strengthening was extremely minimal, so we ignored the
nitrogen contents in ferrite and martensite. In general,
the parameter has a monotonous trend with an increase
in the temperature (760 °C to 820 °C), except for the
ferrite grain.
Table 2: Model parameters for the flow-curve prediction for indivi-
dual phases
Specimen T(°C) d (m) Vm (%) Css
f (%) Css
m (%)
A 760 5.28 48% 0.00688 0.33
B 790 2.31 57% 0.0059 0.21
C 820 4.17 69% 0.00455 0.13
Figure 5 exhibits the stress/strain curve for micro-
mechanical modelling. It can be seen that ferrite shows a
ductile hardening behavior, whereas martensite is more
brittle and also shows a higher strength. In this study, the
grain size and the carbon content played important roles
in the computed yield curves for ferrite and materials.
From Table 2 and Figure 5, the solute carbon content
limits the yield strength of martensite, while the ultimate
stress of ferrite is dependent on the grain size.
The microstructure in OM and the selections of 2D
RVEs from the real microstructures at different inter-
critical temperatures are displayed in Figure 6. The
phases of all the specimens were ferrite and martensite,
and after different intercritical temperatures, the phases
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
906 Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Figure 5: Flow curves for ferrite and martensite phases in a specimen
Figure 4: a) Macrostructural finite-element mode, b) microstructural
finite-element model
presented different microstructures. After the heat treat-
ment (at 760 °C), lath martensite and banded ferrite
coexisted in steel, as shown in Figure 6a. When the
second quenching temperature was 790 °C, the grain
boundaries were clear and the parallel lath martensite
was distributed in the grains (Figure 6b). The amount of
lath martensite increased and the grain boundaries
disappeared at 820 °C (Figure 6c).
After a 5 % deformation, the von Mises stress and
equivalent plastic-strain distributions for the specimens
were simulated; the results are shown in Figure 7. It can
be clearly concluded that the stress is mainly carried by
martensite, and a higher stress difference exists between
the ferrite and martensite phases due to the von Mises
stress distributions. For the first specimen, the maximum
stress mainly focuses on the small connecting belt
among the martensite grains (Figure 7a). Regarding the
second specimen, the maximum stress is mainly distri-
buted along the lath martensite (Figure 7b). The last
specimen’s maximum stress is distributed uniformly in
martensite (Figure 7c). Furthermore, the equivalent
plastic-strain distributions show shear bands in the ferrite
areas; it is easily seen that the amount of shear bands of
specimen (A) is low, while more shear bands consist in
specimen (C).
Because of different microstructures, there are some
differences in the shape of fracture surfaces. On the first
specimen, there are more shear-fracture dimples on the
fracture face and a tearing dimple is shown (Figure 8a).
For the second specimen, tear dimples are obviously
seen from the fracture morphology; its fracture tearing
edge is higher and the tearing dimple is deeper (Fig-
ure 8b). For the third specimen, the micro-fracture mor-
phology shows some large equiaxed dimples (Fig-
ure 8c).
Therefore, the effective stress/strain and work-
hardening-rate/true-strain curves from the 2D RVE
calculations were corrected to 3D curves using Equation
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910 907
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Figure 7: Von Mises stress distributions (left) and equivalent plastic-
strain distributions (right) on 2D RVEs
Figure 6: Microstructure in OM and selections of 2D RVEs from real microstructures at different intercritical temperatures: a) 760 °C, b) 790 °C
and c) 820 °C
characteristic plastic-strain distribution,  and  are the
equivalent plastic strain and the rate of the equivalent
plastic strain, respectively.
Finally, the damage-evolution law is given as:
f f f f
f
s
= + = − ⋅ +
+ −
−
  ( ) 
exp
growth nucleation kk
N
N
1
2
1
2


π


N
NS
⎛
⎝
⎜
⎞
⎠
⎟
⎡
⎣
⎢
⎤
⎦
⎥⋅
2 (10)
In this damage model, N and SN were assumed to be
0.3 and 0.1, while the other parameters were determined
with a calibration of the numerical results using experi-
mental data. 24
3.3 Macro/microscopic finite-element model and boun-
dary conditions
All the associated numerical models were performed
with the finite-element software ABAQUS. Figure 4a
shows the macroscopic finite-element model for a ten-
sile-test specimen, and the specimen was modeled with
solid, reduced-integration element C3D8R. The boun-
dary condition and the load plan, according to which the
upper part of the specimen was stretched and the bottom
was fixed was consistent with the real condition.
In the case of the macroscopic finite-element simu-
lations, all the specimens were generally considered as
continuous and homogeneous. On the microscale, each
constituent character of the microstructure of a steel
specimen needed to be incorporated. RVE was applied to
take into account the influences of different deformation
behaviors of each phase on the resulting mechanical
properties. The 2D RVE models were generated from
real microstructures modeled with plane strain element
CPE4R in ABAQUS. The selection requirement is that
the volume percent of martensite is the same with the
real microstructure in 2D RVE. The boundary condition
and load plan are shown in Figure 4b; the bottom of 2D
RVE is fixed and the top of it is pulled.
4 RESULTS AND DISCUSSION
4.1 Microscale simulation results
Intercritical temperature T, ferrite grain size d,
martensite fraction Vm, along with the carbon content in
ferrite Css
f and in martensite Css
m are listed in Table 2.
After calculating the nitrogen content, we found that the
effect of the nitrogen content on the solid-solution
strengthening was extremely minimal, so we ignored the
nitrogen contents in ferrite and martensite. In general,
the parameter has a monotonous trend with an increase
in the temperature (760 °C to 820 °C), except for the
ferrite grain.
Table 2: Model parameters for the flow-curve prediction for indivi-
dual phases
Specimen T(°C) d (m) Vm (%) Css
f (%) Css
m (%)
A 760 5.28 48% 0.00688 0.33
B 790 2.31 57% 0.0059 0.21
C 820 4.17 69% 0.00455 0.13
Figure 5 exhibits the stress/strain curve for micro-
mechanical modelling. It can be seen that ferrite shows a
ductile hardening behavior, whereas martensite is more
brittle and also shows a higher strength. In this study, the
grain size and the carbon content played important roles
in the computed yield curves for ferrite and materials.
From Table 2 and Figure 5, the solute carbon content
limits the yield strength of martensite, while the ultimate
stress of ferrite is dependent on the grain size.
The microstructure in OM and the selections of 2D
RVEs from the real microstructures at different inter-
critical temperatures are displayed in Figure 6. The
phases of all the specimens were ferrite and martensite,
and after different intercritical temperatures, the phases
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
906 Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Figure 5: Flow curves for ferrite and martensite phases in a specimen
Figure 4: a) Macrostructural finite-element mode, b) microstructural
finite-element model
presented different microstructures. After the heat treat-
ment (at 760 °C), lath martensite and banded ferrite
coexisted in steel, as shown in Figure 6a. When the
second quenching temperature was 790 °C, the grain
boundaries were clear and the parallel lath martensite
was distributed in the grains (Figure 6b). The amount of
lath martensite increased and the grain boundaries
disappeared at 820 °C (Figure 6c).
After a 5 % deformation, the von Mises stress and
equivalent plastic-strain distributions for the specimens
were simulated; the results are shown in Figure 7. It can
be clearly concluded that the stress is mainly carried by
martensite, and a higher stress difference exists between
the ferrite and martensite phases due to the von Mises
stress distributions. For the first specimen, the maximum
stress mainly focuses on the small connecting belt
among the martensite grains (Figure 7a). Regarding the
second specimen, the maximum stress is mainly distri-
buted along the lath martensite (Figure 7b). The last
specimen’s maximum stress is distributed uniformly in
martensite (Figure 7c). Furthermore, the equivalent
plastic-strain distributions show shear bands in the ferrite
areas; it is easily seen that the amount of shear bands of
specimen (A) is low, while more shear bands consist in
specimen (C).
Because of different microstructures, there are some
differences in the shape of fracture surfaces. On the first
specimen, there are more shear-fracture dimples on the
fracture face and a tearing dimple is shown (Figure 8a).
For the second specimen, tear dimples are obviously
seen from the fracture morphology; its fracture tearing
edge is higher and the tearing dimple is deeper (Fig-
ure 8b). For the third specimen, the micro-fracture mor-
phology shows some large equiaxed dimples (Fig-
ure 8c).
Therefore, the effective stress/strain and work-
hardening-rate/true-strain curves from the 2D RVE
calculations were corrected to 3D curves using Equation
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910 907
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Figure 7: Von Mises stress distributions (left) and equivalent plastic-
strain distributions (right) on 2D RVEs
Figure 6: Microstructure in OM and selections of 2D RVEs from real microstructures at different intercritical temperatures: a) 760 °C, b) 790 °C
and c) 820 °C
(5). These corrected curves are shown in Figure 9; speci-
men (B) shows the highest work hardening behaviors,
higher stress level and yield stress.
4.2 Macroscale simulation results
The flow curves were obtained through micromecha-
nical modeling and inputted into the macromechanical
model. Subsequently, the GTN damage model was
applied to investigate the failure behavior of the material.
After macromechanical modeling, the stress/strain
curves of the simulation were compared with the experi-
mental curves, as shown in Figure 10. It can be observed
that the macromechanical model based on the flow
curves of micromechanical modeling provides very good
estimates for the experimental results.
The equivalent plastic-strain distribution and void-
volume-fraction distribution in tensile samples (1/2 part
of the extensometer measurement) at a 6 % deformation
are shown in Figure 11. After the deformation due to the
engineering strain of 6 %, the plastic strain and void
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
908 Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Figure 8: Fracture surfaces of the heat-treated specimens at different intercritical temperatures: a) 760 °C, b) 790 °C and c) 820 °C
Figure 11: a) Equivalent plastic-strain distribution, b) void-volume-
fraction distribution in 1/2 part of the extensometer measurement, at a
deformation of 6 %
Figure 10: Experimental and simulated stress/strain curves of inter-
critically heat-treated steels
Figure 9: True stress/strain and work-hardening curves obtained
through micromechanical modeling
volume fraction mainly focus on the core of a specimen.
For specimen (A), the equivalent plastic strain is the
smallest, but the void volume fraction is the largest for
specimen (C).
In addition, development curves of the equivalent
plastic strain and the mean void volume fraction were
calculated during the simulation process. With an in-
crease in the displacement, the development trends of the
equivalent plastic strain and the mean void volume
fraction for different specimens were similar, as shown
in Figure 12. It can be seen that there are three processes
(we took the third specimen as an example).
During the first stage, the growth rate of the
equivalent plastic strain is almost constant and the mean
void volume fraction initially remained unchanged,
followed by a slow increase, which shows that the spe-
cimen is in the state of uniform deformation. After the
first stage, the equivalent plastic strain increases rapidly
and the volume fraction of voids increases greatly, which
means the specimen goes into the inhomogeneous defor-
mation stage and a lot of voids begin to form and grow
up. During the last stage, the mean volume fraction of
voids and the equivalent plastic strain continue to
increase, followed by a slow increase, and come to be
maximum in the end. This illustrates that the coalescence
of voids plays an import part in the third stage. For the
specimen, large cracks are formed and the sample finally
breaks. After the comparison between the equivalent
plastic strain and the mean void volume fraction, we can
see that specimen (B) shows a great elongation perfor-
mance because of the lowest void-volume-fraction
growth rate (Figure 12).
5 CONCLUSIONS
The Gr.65 steel underwent an intercritical heat-treat-
ment process. The micromechanical properties of the
heat-treated steel were predicted, taking into account the
effects of the microstructure, phase fractions, local com-
positions of single phases and their area shapes. The
GTN model was used for predicting the damage behavior
of the specimens for the macromechanical model. The
following points could be drawn:
1) The tensile strength increased with the increasing
temperature due to an increase in the amount of
martensite in the steel, but the hardening behavior of
this specimen was affected by the microstructure.
2) The flow curves of the Gr.65 steel at different
interitical temperatures could be well predicted using
the 2D FE simulation based on real microstructures.
Simulation results showed that higher stress con-
centrated on the martensite; at the same time, the
shear-band appearance strongly depended on the
microstructures of the phases.
3) For the prediction of damage behaviors, the true
stress/true strain curves of macroscale simulations
showed good agreement with the experiments in-
volving differently heat-treated steels.
6 REFERENCES
1 H. J. Jun, J. S. Kang, D. H. Seo, F. K. Kim, Effects of deformation
and boron on microstructure and continuous cooling transformation
in low carbon HSLA steels, Materials Science and Engineering A,
422 (2006) 1, 157–162, doi:10.1016/j.msea.2005.05.008
2 S. Oliver, T. B. Jones, G. Fourlaris, Dual phase versus TRIP strip
steels: comparison of dynamic properties for automotive crash
performance, Materials Science and Technology, 23 (2007) 4,
423–431, doi:10.1179/174328407X168937
3 Y. I. Son, Y. K. Lee, K. T. Park, Ultrafine grained ferrite–martensite
dual phase steels fabricated via equal channel angular pressing:
microstructure and tensile properties, Acta Materialia, 53 (2005) 11,
3125–3134, doi:10.1016/j.actamat.2005.02.015
4 L. Shi, Z. Yan, Y. Liu, D. G. Li, Improved toughness and ductility in
ferrite/acicular ferrite dual-phase steel through intercritical heat
treatment, Materials Science and Engineering A, 590 (2005), 7–15,
doi:10.1016/j.msea.2005.10.006
5 J. Kang, C. Wang, G. D. Wang, Microstructural characteristics and
impact fracture behavior of a high-strength low-alloy steel treated by
intercritical heat treatment, Materials Science and Engineering A,
553 (2012), 96–104, doi:10.1016/j.msea.2012.05.098
6 Z. J. Xie, S. F. Yuan, W. H. Zhou, G. D. Wang, Stabilization of
retained austenite by the two-step intercritical heat treatment and its
effect on the toughness of a low alloyed steel, Materials & Design,
59 (2014), 193–198, doi:10.1016/j.matdes.2014.02.035
7 M. A. Maleque, Y. M. Poon, H. H. Masjuki, The effect of inter-
critical heat treatment on the mechanical properties of AISI 3115
steel, Journal of Materials Processing Technology, 153 (2004),
482–487, doi:10.1016/j.jmatprotec.2004.04.391
8 W. H. Zhou, X. L. Wang, P. K. C. Venkatsurya, R. F. Shi, Struc-
ture–mechanical property relationship in a high strength low carbon
alloy steel processed by two-step intercritical annealing and inter-
critical tempering, Materials Science and Engineering A, 607 (2014),
569–577, doi:10.1016/j.msea.2014.03.107
9 M. Azuma, S. Goutianos, N. Hansen, D. F. White, Effect of hardness
of martensite and ferrite on void formation in dual phase steel,
Materials Science and Technology, 28 (2012) 9, 1092–1100,
doi:10.1179/1743284712Y.0000000006
10 B. C. Hwang, T. Y. Cao, S. Y. Shin, F. H. Kim, Effects of ferrite grain
size and martensite volume fraction on dynamic deformation
behaviour of 0.15C–2.0Mn–0.2Si dual phase steels, Materials
Science and Technology, 21 (2005) 8, 967–975, doi:10.1179/
174328405X47609
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910 909
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Figure 12: Comparison between the equivalent plastic strain and
mean void volume fraction during loading
(5). These corrected curves are shown in Figure 9; speci-
men (B) shows the highest work hardening behaviors,
higher stress level and yield stress.
4.2 Macroscale simulation results
The flow curves were obtained through micromecha-
nical modeling and inputted into the macromechanical
model. Subsequently, the GTN damage model was
applied to investigate the failure behavior of the material.
After macromechanical modeling, the stress/strain
curves of the simulation were compared with the experi-
mental curves, as shown in Figure 10. It can be observed
that the macromechanical model based on the flow
curves of micromechanical modeling provides very good
estimates for the experimental results.
The equivalent plastic-strain distribution and void-
volume-fraction distribution in tensile samples (1/2 part
of the extensometer measurement) at a 6 % deformation
are shown in Figure 11. After the deformation due to the
engineering strain of 6 %, the plastic strain and void
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
908 Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Figure 8: Fracture surfaces of the heat-treated specimens at different intercritical temperatures: a) 760 °C, b) 790 °C and c) 820 °C
Figure 11: a) Equivalent plastic-strain distribution, b) void-volume-
fraction distribution in 1/2 part of the extensometer measurement, at a
deformation of 6 %
Figure 10: Experimental and simulated stress/strain curves of inter-
critically heat-treated steels
Figure 9: True stress/strain and work-hardening curves obtained
through micromechanical modeling
volume fraction mainly focus on the core of a specimen.
For specimen (A), the equivalent plastic strain is the
smallest, but the void volume fraction is the largest for
specimen (C).
In addition, development curves of the equivalent
plastic strain and the mean void volume fraction were
calculated during the simulation process. With an in-
crease in the displacement, the development trends of the
equivalent plastic strain and the mean void volume
fraction for different specimens were similar, as shown
in Figure 12. It can be seen that there are three processes
(we took the third specimen as an example).
During the first stage, the growth rate of the
equivalent plastic strain is almost constant and the mean
void volume fraction initially remained unchanged,
followed by a slow increase, which shows that the spe-
cimen is in the state of uniform deformation. After the
first stage, the equivalent plastic strain increases rapidly
and the volume fraction of voids increases greatly, which
means the specimen goes into the inhomogeneous defor-
mation stage and a lot of voids begin to form and grow
up. During the last stage, the mean volume fraction of
voids and the equivalent plastic strain continue to
increase, followed by a slow increase, and come to be
maximum in the end. This illustrates that the coalescence
of voids plays an import part in the third stage. For the
specimen, large cracks are formed and the sample finally
breaks. After the comparison between the equivalent
plastic strain and the mean void volume fraction, we can
see that specimen (B) shows a great elongation perfor-
mance because of the lowest void-volume-fraction
growth rate (Figure 12).
5 CONCLUSIONS
The Gr.65 steel underwent an intercritical heat-treat-
ment process. The micromechanical properties of the
heat-treated steel were predicted, taking into account the
effects of the microstructure, phase fractions, local com-
positions of single phases and their area shapes. The
GTN model was used for predicting the damage behavior
of the specimens for the macromechanical model. The
following points could be drawn:
1) The tensile strength increased with the increasing
temperature due to an increase in the amount of
martensite in the steel, but the hardening behavior of
this specimen was affected by the microstructure.
2) The flow curves of the Gr.65 steel at different
interitical temperatures could be well predicted using
the 2D FE simulation based on real microstructures.
Simulation results showed that higher stress con-
centrated on the martensite; at the same time, the
shear-band appearance strongly depended on the
microstructures of the phases.
3) For the prediction of damage behaviors, the true
stress/true strain curves of macroscale simulations
showed good agreement with the experiments in-
volving differently heat-treated steels.
6 REFERENCES
1 H. J. Jun, J. S. Kang, D. H. Seo, F. K. Kim, Effects of deformation
and boron on microstructure and continuous cooling transformation
in low carbon HSLA steels, Materials Science and Engineering A,
422 (2006) 1, 157–162, doi:10.1016/j.msea.2005.05.008
2 S. Oliver, T. B. Jones, G. Fourlaris, Dual phase versus TRIP strip
steels: comparison of dynamic properties for automotive crash
performance, Materials Science and Technology, 23 (2007) 4,
423–431, doi:10.1179/174328407X168937
3 Y. I. Son, Y. K. Lee, K. T. Park, Ultrafine grained ferrite–martensite
dual phase steels fabricated via equal channel angular pressing:
microstructure and tensile properties, Acta Materialia, 53 (2005) 11,
3125–3134, doi:10.1016/j.actamat.2005.02.015
4 L. Shi, Z. Yan, Y. Liu, D. G. Li, Improved toughness and ductility in
ferrite/acicular ferrite dual-phase steel through intercritical heat
treatment, Materials Science and Engineering A, 590 (2005), 7–15,
doi:10.1016/j.msea.2005.10.006
5 J. Kang, C. Wang, G. D. Wang, Microstructural characteristics and
impact fracture behavior of a high-strength low-alloy steel treated by
intercritical heat treatment, Materials Science and Engineering A,
553 (2012), 96–104, doi:10.1016/j.msea.2012.05.098
6 Z. J. Xie, S. F. Yuan, W. H. Zhou, G. D. Wang, Stabilization of
retained austenite by the two-step intercritical heat treatment and its
effect on the toughness of a low alloyed steel, Materials & Design,
59 (2014), 193–198, doi:10.1016/j.matdes.2014.02.035
7 M. A. Maleque, Y. M. Poon, H. H. Masjuki, The effect of inter-
critical heat treatment on the mechanical properties of AISI 3115
steel, Journal of Materials Processing Technology, 153 (2004),
482–487, doi:10.1016/j.jmatprotec.2004.04.391
8 W. H. Zhou, X. L. Wang, P. K. C. Venkatsurya, R. F. Shi, Struc-
ture–mechanical property relationship in a high strength low carbon
alloy steel processed by two-step intercritical annealing and inter-
critical tempering, Materials Science and Engineering A, 607 (2014),
569–577, doi:10.1016/j.msea.2014.03.107
9 M. Azuma, S. Goutianos, N. Hansen, D. F. White, Effect of hardness
of martensite and ferrite on void formation in dual phase steel,
Materials Science and Technology, 28 (2012) 9, 1092–1100,
doi:10.1179/1743284712Y.0000000006
10 B. C. Hwang, T. Y. Cao, S. Y. Shin, F. H. Kim, Effects of ferrite grain
size and martensite volume fraction on dynamic deformation
behaviour of 0.15C–2.0Mn–0.2Si dual phase steels, Materials
Science and Technology, 21 (2005) 8, 967–975, doi:10.1179/
174328405X47609
A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910 909
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
Figure 12: Comparison between the equivalent plastic strain and
mean void volume fraction during loading
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A. J. ZHAO et al.: PHASE-TRANSFORMATION BEHAVIOR AND MICROMECHANICAL PROPERTIES ...
910 Materiali in tehnologije / Materials and technology 51 (2017) 6, 903–910
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
C. B. ZHENG et al.: EIS AND SKP STUDY ON IMPROVEMENT OF THE PROTECTION PERFORMANCE ...
911–918
EIS AND SKP STUDY ON IMPROVEMENT OF THE PROTECTION
PERFORMANCE OF AN ALKYD-VARNISH COATING MODIFIED
WITH AIR-PLASMA TREATMENT ON Q235 STEEL
EIS IN SKP [TUDIJA IZBOLJ[ANJA ZA[^ITE Z ALKIDNO
PREVLEKO, MODIFICIRANO S PLAZEMSKO OBDELAVO NA
Q235 JEKLU
Chuanbo Zheng1, Haoyang Qu1, Wei Wang2
1Jiangsu University of Science and Technology, School of Materials Science and Engineering, Mengxi Road 2, Zhenjiang, Jiangsu Province,
212003, China
2State Key Laboratory for Marine Corrosion and Protection, Luoyang Ship Material Research Institute (LSMRI), 149-1 Zhuzhou Road,
Laoshan District, Qingdao, 266101, China
15952802516@139.com
Prejem rokopisa – received: 2017-02-16; sprejem za objavo – accepted for publication: 2017-05-30
doi:10.17222/mit.2017.022
Electrochemical impedance spectroscopy (EIS) and scanning Kelvin probe (SKP) were used to study the failure of the coating
modified with air-plasma treatment in 3.5 % mass fraction of NaCl. The results show that the failure process can be divided into
three stages including water penetration, accumulation of corrosion products at the location of a defect, water penetration of the
entire coating and coating failure. The EIS results show that the plasma-treated samples exhibit an additional
electrical double-layer capacitor that delays the coating failure during the second stage. The potential of the non-plasma-treated
samples decrease faster than that of the plasma-treated samples according to the SKP study. As the transition layer, air plasma
delays the coating failure due to the chemical bond between the metal substrate and the coating.
Keywords: air-plasma treatment, electrochemical impedance spectroscopy, scanning Kelvin probe, coating failure, hydrostatic
pressure
Elektrokemijsko impedan~no spektroskopijo (angl. EIS) in vrsti~no Kelvinovo sondo (SKP) smo uporabili za preu~evanje
po{kodb na prevleki, modificirani s plazmo, v 3,5 mas. % raztopini NaCl. Rezultati ka`ejo, da se nastajanje po{kodb lahko
razdeli na tri faze, vklju~no s penetracijo vode: na akumulacijo korozijskih produktov na mestu napake, na prodiranje vode v
dele celotne prevleke in na odpoved prevleke. Rezultati EIS ka`ejo, da imajo vzorci, obdelani v plazmi, dodaten elektri~ni
dvoslojni kondenzator, ki v drugi fazi upo~asni po{kodbo prevleke (premaza). SKP-analize so pokazale, da se potencial
plazemsko neobdelanih vzorcev zmanj{uje hitreje kot tistih vzorcev, ki so bili obdelani s plazmo. Kot prehodna plast, obdelava z
zra~no plazmo upo~asni po{kodbo premaza zaradi kemijske vezi med kovinskim substratom in prevleko.
Klju~ne besede: obdelava s plazmo, elektrokemijska impedan~na spektroskopija, skeniranje s sondo Kelvin, odpoved prevleke,
hidrostati~ni tlak
1 INTRODUCTION
With the exploration and development of marine
research, problems of the corrosion and protection of
structural materials and organic coatings in marine
environments have increasingly gained attention.1–3
Thus, the corrosion problems of the materials under
deep-ocean conditions must be considered. The environ-
ment of the deep ocean is a complicated system,
including hydrostatic pressure, differently dissolved
oxygen (DO), all kinds of salts, water velocity and
suspended silt. Coating deterioration, delamination,
blistering and penetration are likely to occur in the
deep-sea environment.4–6 How to improve the coating
quality is an urgent need for deep-ocean resource
exploitation.
Due to the fact that air-plasma treatment7–10 does not
alter the overall performance of the substrate, the
chemical and physical activities of a material surface re-
mains constant. During a reaction, polar oxygen groups
with single or double bonds can be incorporated into the
surface of the substrate, enhancing the wettability of the
substrate surface.11,12 After the air-plasma treatment, the
hydrophobicity and spreading ability of the material
surface were also found to improve. In addition, porosity
is effectively reduced on the substrate surface due to a
better wettability of the organic coating. Therefore, the
coating in combination with the surface of the substrate
allows a better modification process.7,8,13
Many researchers effectively modified surfaces with
plasma-processing techniques in order to increase
hydrophobicity and alter only the surface properties of
the material. L. Wang et al.14 prepared a column-like nano/
micro-scale topography surface via trichloro(octyl)silane
(TCOS) vapor deposition on a polydimethylsiloxane
oxidized with air plasma. Their results showed a
successful assembly of TCOS on a polydimethylsiloxane
surface. An addition of n-heptanol to alkylsiloxane also
helps regulate the scale and roughness of the column-like
nano/micro-scale topography. C. Riccardi et al.15 induced
Materiali in tehnologije / Materials and technology 51 (2017) 6, 911–918 911
MATERIALI IN TEHNOLOGIJE/MATERIALS AND TECHNOLOGY (1967–2017) – 50 LET/50 YEARS
UDK 621.793.5:620.1:669.018.26 ISSN 1580-2949
Original scientific article/Izvirni znanstveni ~lanek MTAEC9, 51(6)911(2017)