Corrosion Deterioration of Steel in Cracked SHCC
International Journal of Concrete Structures and Materials
Corrosion Deterioration of Steel in Cracked SHCC
Suvash Chandra Paul
Gideon Pieter Adriaan Greeff van Zijl
The presence of cracks is unavoidable in reinforced concrete structures and also a gateway for chloride into concrete, leading to corrosion of steel reinforcing bars. So, crack control, crack width limitation and chloride threshold levels are wellestablished concepts in durability of reinforced concrete structures. This paper reports on accelerated chloride-induced corrosion in cracked reinforced strain-hardening cement-based composites and reinforced mortar beams, both in loaded and unloaded states. Corrosion rates are monitored and loss of mass and yield force, as well as corrosion pitting depth in steel bars are reported. The chloride content at different depths in specimens is also determined through XRF, and through chemical testing of acid and water soluble chloride content by titration. Finally, different relationships are drawn between crack properties, mass loss, yield force, corroded depth and chloride levels at the steel surface for different cover depths. It is found that the crack spacing and free chloride at the steel surface level are best correlated to the corrosion damage in the specimens.
chloride; corrosion; cracks; steel yield force; pitting depth; mass loss; SHCC
Cracks in concrete act as pathways for quick ingress of
chloride and water. There is evidence that a threshold crack
width of about 0.05 mm exists for water permeability. For
cracks wider than this threshold, water permeability
increases by orders of magnitude
(Wang et al. 1997)
diffusivity has been shown to increase linearly with crack
width increase in the range 0.03–0.2 mm
(Djerbi et al.
. Little or no chloride appears to be transported through
cracks with widths below the lower threshold of 0.03 mm,
while the diffusivity stabilises at crack widths beyond the
upper threshold (0.2 mm). Once oxygen also reaches the
steel, an electrochemical cell forms and corrosion of steel
reinforcement commences. In reinforced concrete structures
(RCS) chloride ions act as a catalyst in the localised
breakdown of the passive film on the steel surface. The
passive film initially forms on steel as a result of the alkaline
nature of the pore solution in concrete. A minimum
concentration of chlorides on the steel, known as the chloride
threshold level, is required to de-passivate the reinforcement
and allow corrosion to occur
(Angst et al. 2009; Pacheco and
. Once the steel has been fully de-passivated,
the rate of corrosion depends on the availability of oxygen
and moisture, as well as structural geometry. After corrosion
initiation, corrosion products (iron oxides and hydroxides)
form and accumulate. These products have a greater volume
than the original iron, which leads to internal stresses and may
result in cracking and spalling of the concrete cover
et al. 2016; Broomfield 2007)
. Corrosion initiation periods
were investigated by the
in RC containing cracks
of various widths. A maximum of 30 days to corrosion
initiation was observed in specimens containing 0.025 mm wide
cracks. For a crack width of 0.09 mm, the corrosion initiation
period was about 14 days. Note that in this study, the
specimens were exposed to a 10% salt aqueous solution
continuously for 49 days of testing. It illustrates that the presence of
wider cracks in concrete specimens accelerates the corrosion
process and reduces the initiation period. However, in the same
study no relationship was found between the corrosion rate and
crack widths. The blockage of cracks with the corrosion
products, and possible self-healing of cracks may reduce the
corrosion rate in time. Another study by
Mohammed et al.
on microcell corrosion in cracked concrete found a
linear relationship between the corrosion current density and
crack widths in the range 0.1–0.7 mm at early stages. In a later
stage, after 15 weeks of testing, no relationship was found.
In the chloride-induced corrosion process, a certain
amount of chlorides must reach the steel to initiate the
corrosion process. This value is significant because it is an input
parameter in the service life design and service life
prediction models of RCS. However, it is still uncertain what the
threshold amount of chloride for corrosion initiation is, and
thus guidelines for durability design, and preventative repair
of RCS are unclear.
suggested that the
critical chloride content (% by cement wt) is about
proposed ranges of 0.9–1.8% and 0.48–2.02%.
The threshold chloride content depends on several factors
like the concrete pH level, cement and admixture types and
water/binder ratio. Another important issue is whether all the
chloride inside concrete contributes to the corrosion process.
Depending on the matrix binding capacity, some chloride
may be bound in matrix compositions and do not contribute
to the corrosion. So, the amount of free chloride in concrete
may be more representative than the total chloride content.
German and Zaborski (2011)
suggested a threshold value for
free chloride of about 0.35% (by concrete wt) in concrete.
For the service life prediction of RCS, time to initiation of
corrosion is typically modelled by Fick’s Law
(Liang et al.
2009; Tang et al. 2015)
. The challenge in corrosion
modelling is rigorous calibration at the hand of sufficiently large
experimental data sets, as well as field validation.
The availability of research data on corrosion in fibre
reinforced concrete (FRC) is also limited and not thoroughly
understood due to the complex mechanism of corrosion
(Yoo and Yoon 2016)
. A particular class of FRC is
discussed in this paper, namely strain-hardening
cementbased composites (SHCC)
. The main feature of
SHCC is the formation of multiple fine cracks and increased
or maintained tensile resistance upon increased deformation
whilst more cracks appear. Limited field and laboratory data
exist for a new material like SHCC, and embedded steel bar
reinforcement (R/SHCC), and only limited researches have
been performed on total chloride content and corrosion in
(Sahmaran et al. 2008; Kobayashi et al. 2010;
Mihashi et al. 2011; Wittmann et al. 2011; Kobayashi and
Rokugo 2013; Paul and van Zijl 2014, 2016)
SHCC and R/SHCC have the ability to form multiple
cracks with limited crack widths during the strain-hardening
(Paul and van Zijl 2013)
. It is further known that crack
width and cover depth play an important role in corrosion of
the steel reinforcing
(Bashir et al. 2017; Soltani et al. 2013)
However, what is unknown is the effect of crack spacing and
whether an optimum range of crack spacing can be found
that will limit the corrosion rate. Therefore it is the aim here
to investigate the effect of crack spacing on chloride-induced
corrosion of R/SHCC. This research paper presents the test
program and results of accelerated chloride-induced
corrosion of pre-cracked R/SHCC specimens with three levels of
cover depths, and one level of steel bar reinforcement. As
reference, reinforced mortar (R/mortar) specimens were also
made with two different cover depths. Results are presented
of laboratory tests over a minimum of 6 months to a
maximum of 2 years of cyclic wetting and drying with chloride
aqueous solution, on a total of 64 R/SHCC and R/mortar
beam specimens. The corrosion rate, consequences of
corrosion of steel bars in terms of mass loss and pitting depths,
and total and free chloride profiles are studied. As far as the
authors could establish, these are the first reports of free
versus total chloride in SHCC, as are the proposed relations
between loss of mass and resistance, as well as pitting depths
in steel bars in R/SHCC. The outcome of this research is
intended to contribute to durability design guidelines with
respect to chloride-induced corrosion in R/SHCC.
2. Research Background
Previous studies on corrosion of R/SHCC specimens
found that little or no corrosion can be seen in the
specimens, compared with significant corrosion observed in
comparative R/mortar or reinforced steel fibre reinforced
cementitious composites (R/SFRC)
(Kobayashi et al. 2010;
Mihashi et al. 2011; Kobayashi and Rokugo 2013)
60 days of accelerated chloride exposure (5 min spraying
with 3% NaCl aqueous solution every 6 h),
Kobayashi et al.
found no sign of corrosion discoloration on the
surface of steel in cracked (0.02–0.06 mm crack widths)
patched specimens made from high performance fibre
reinforced cement composites (HPFRCC) with minimum of
0.75 to maximum 1.5% polyethylene (PE) fibre content.
Mihashi et al. (2011)
reported that only 11.8% of steel bar
surface area was corroded, but with zero corrosion depth in
hybrid fibre reinforced cementitious composite (HFRCC)
specimens, compared with 100% corroded surface area and
3.1 mm corroded depth in reference mortar specimens. In
their case, steel fibres were combined with polymeric fibre.
The steel fibre acted as sacrificial anode, whereby rebar
corrosion was significantly reduced. In this case specimens
were subjected to a 3% NaCl solution by cyclic wetting and
drying, while a potential of 3 V was applied continuously to
the specimens for a period of about 1 year. In neither of
these researches information of the loss of steel bar yield
force due to corrosion was reported.
In a recent work it was found that corrosion reduces the
capacity of steel bars in SHCC
(Paul and van Zijl 2016)
which pitting corrosion could also be observed.
et al. (2008)
investigated the loss of flexural load capacity
due to accelerated corrosion in both R/SHCC and R/mortar
specimens. Reinforced specimens were kept in 5% NaCl
solution for a minimum of 25 h, to a maximum of 300 h,
while a potential of 30 V was applied to the steel in both the
SHCC and mortar specimens. Four-point bending was
performed on the specimens after a period of 25, 50, 75, 100
and 300 h of accelerated corrosion. In mortar specimens,
37% lower load capacity was found after 25 h of exposure
than in control specimens (without any applied current). No
significant change in flexural resistance was found in the
R/SHCC specimens after 50 h of exposure. After 300 h of
exposure, more than 55% loss in flexural capacity was found
in R/SHCC, but in R/Mortar an 85% loss was found after
just 75 h of exposure. In a similar type of experiment
Paul et al. (2017)
also found lower corrosion
mass loss in R/SHCC specimens when compared with
R/Mortar specimens. In none of the researches mentioned
here information about the number of cracks, crack spacing
and their influence on corrosion deterioration was reported.
It should be noted that the electrically enhanced corrosion
exposure as applied by
Mihashi et al. (2011)
et al. (2008)
replicates specific, aggressive corrosion
conditions. A purely cyclic wetting–drying chloride exposure and
a Coulostatic corrosion measuring technique have been
performed by several researchers and are followed here. It is
argued to be a replication of in situ corrosion in coastal
(Glass 1995; Gonzalez et al. 2001; Andrade and
. The applied current and duration in the
Coulostatic method is miniscule, causing negligible
influence on the internal electrochemical corrosion mechanism in
3. Materials and Test Methods
3.1 Mix Design of SHCC
The mix ingredients and proportions are shown in Table 1.
Two different types of sand and cement were used. A fine
sand (FS-SHCC) with fineness modulus (FM) of 1.9 and
local, natural, coarse sand (CS-SHCC) with FM of 3.3 were
used. Polyvinyl alcohol (PVA) fibres with length 12 mm and
diameter 0.04 mm were used.
3.2 Specimen Preparation for Corrosion Testing
The size of beam specimens prepared for one series of
beams, i.e. FS2, CS2, FM1 and CM1 was 100 9
100 9 500 mm, with a single steel reinforcing bar inside.
Three R/SHCC specimens were prepared for each of three
cover depths 15, 25 and 35 mm (C15, C25 and C35), for
both types of sand, i.e. in total 18 specimens. In case of
mortar, only 15 mm cover depth was chosen; i.e. in total 6
R/mortar specimens were made. In all specimens a 10 mm
diameter (Y10) high tensile ribbed bar (B1) with a tensile
yield stress of about 510 MPa was used. Electrical
connection for corrosion testing was facilitated by letting the steel
protrude at the top of one series of specimens (see Fig. 1a).
Note that in these specimens a stainless plate (15 mm thick
and 150 mm long) was used as a counter electrode (CE). For
FS31, FS32 and FM2 the beam specimen size was
80 9 100 9 490 mm, also with a single Y10 bar (see
Fig. 1c). Here a 5 mm stainless steel with the same length as
B1 was cast into the specimens as a CE, protruding 5 mm
from each beam end—see Fig. 1b–d. The three cover depths
were also used for FS31 in two levels of deflection, giving in
total 18 specimens (nine for each level of deflection) for
FS31. For FS32 and FM2 specimens, only C15 and C25
cover depths were used. A total of 22 specimens were made
from both FS32 and FM2. Details of specimens FS32 and
FM2 are shown in Table 2, and full details can be found in
3.3 Method of Forming Cracks in the Specimens
Three-point flexural testing was performed in the
specimens for creating cracks in FS2 and CS2 specimens. In this
regard, specimens were loaded up to a vertical deflection of
3.5 mm (D3.5) in a Zwick Z250 materials testing machine
(MTM) and then unloaded. In case of FM1 and CM1,
specimens were loaded to a deflection of only 1.5 mm
(D1.5). The number of cracks and crack widths were
recorded. FS31 specimens were loaded up to two different
levels of vertical deflections 5 mm and 7 mm (D5 and D7)
respectively in the Zwick Z250 MTM, in order to study the
effect of two different crack patterns. After unloading from
the Zwick, all FS31 specimens were placed in special steel
frames for sustained loading at the same level of deflections
(D5 and D7) by tightening the bolts on both ends of each
specimen as shown in Fig. 1c. In case of FS32 and FM2,
cracks were formed by making notches in the specimens as
shown in Table 2 and then specimens were placed and
cracked in similar steel frames as FS31 specimens. The FS32
specimen cracks were formed while measuring the
deformation over a central 100 mm gauge length on the notched
surface as shown in Fig. 1c. Details of specimen surface
strain and the number of notches are shown in Table 2.
Figure 2a shows the way in which the number of crack and
crack spacing were measured in the specimen and Fig. 2b
shows the pitting depth measurement in the corroded steel
bars which will be discussed in the research results section.
3.4 Corrosion Test Setup
In this study the cracked specimens were subjected to the
cyclic wetting (3 days) with 3.5% NaCl aqueous solution,
a Notches of 3 mm wide and 10 mm deep were sawn to create artificial cracks in the middle of each specimen, as shown in Fig. 1c.
b During pre-cracking, the deformation was controlled to the value shown over a central 100 mm gauge length.
followed by drying (4 days). FS2, CS2, FM1 and CM1
specimens were subjected to capillary absorption in
unloaded state (up to 108 weeks for SHCC and 85 weeks for
mortar). Note that, only one directional capillary absorption
of the chloride solution through the cracked face of
specimens was allowed. In this case, the central 200 mm cracked
length of each specimen was placed in contact with chloride
solution, and other surfaces were sealed. In contrast, a
ponding method was followed in case of FS31 (up to
57 weeks), FS32 (up to 28 weeks) and FM2 (up to
28 weeks) on specimens in the loaded state. Here a 200 mm
long and 100 mm wide pond was made on the cracked
surface of each specimen by non-absorbent plastic as shown
in Fig. 1b. All the connections between plastic and the
specimen surface were sealed to prevent leakage.
A Coulostatic method was used to measure the corrosion
rate. It replicates the Randles circuit, a type of polarisation
measuring technique, where a small amount of a known
current (DI) is passed through the steel for a known amount
of time (Dt) while the potential decay (DE) is observed. The
polarisation resistance (Rp) of the concrete can then be
determined from the DI/DE ratio. The Y10 steel reinforcing
bar acted as the working electrode (WE), the stainless steel
as the CE, and an Ag/AgCl half-cell was used as the
reference electrode as seen in Fig. 1d. For the corrosion rate
measurement, a Spider8 data logger was used while a current
of 4 mA was applied to each of the specimens for a period of
5 mS by a laboratory built current pulse generator. The
reading from each specimen was collected on the last day of
each wet and dry cycle. Note that in the case of FS31
specimens with 15 mm cover, lower perturbation was found
for the 4 mA applied current and as a result a different
current (10 mA) was applied for a 6 mS time period for
Once Rp is known, the corrosion current (Icorr) (lA/cm2)
and from that corrosion rate (Vcorr) (mm/year) can be
determined. So during corrosion testing, if no further current
is applied to the steel bar, the potential decays can be
represented exponentially with time as follows:
where g0 is the initial potential shift and gt the potential shift
(DE) at time t. Rp is then obtained from the time constant (sc)
and interfacial capacitance (C) as follows:
gt ¼ g0 exp
Rp ¼ C
where g0 and sc values can be determined by fitting an
exponential function to the perturbation (mV) vs time (sec)
curve. B is the Stern-Geary constant varying from 26 to
52 mV depending on whether passive or active corrosion is
occuring. In this research project the B value was considered
to be 26 mV. This corrosion rate measuring technique was
broadly explained by
Gonzalez et al. (2001)
Andrade and Alonso (2004)
After a certain period of cyclic wetting and drying
exposure, some of the specimens were broken to observe actual
corrosion in the steel bar. The outcomes like loss of rebar
tensile yield force and pitting depths were then estimated and
compared with experimental outcomes. In this regard,
Eqs. (5–6) were used to estimate uniformly corroded depth
(dc) and change in rebar yield force due to uniform cross
section reduction (DFy) as follows:
dcðtÞ ¼ dc0 þ
¼ dc0 þ
2 Vcorr þ Vcorr;i 1 ðti
Icorr ¼ Rp
DFyðtÞ ¼ p4ry nds2
with dc0 the initial corroded depth, ti the corroding period, qs
the steel density, ds and Ls the steel diameter and length (here
assumed to be the full 500 mm) and ry is the steel nominal
yield stress. Note that only for FS32 and FM2 the rebar mass
loss was determined by subtracting the weight measured
after testing from the initial weight. Note that Eq. (6) does
not consider stress concentrations due to the geometrical
changes caused by pitting corrosion.
3.5 Method of Determining Chloride Content
Chloride penetration profiles were drawn for both
R/SHCC and R/mortar specimens after a certain period of
accelerated testing. More than 20 specimens were chosen
from different SHCC and mortar types after different
exposure periods. Chloride profiles were then obtained by
drilling from the exposed surface in layers of 3 mm each, up
to a depth of 45 mm in most specimens. A 16 mm drill was
used to collect powder samples at the different layers in
specimens. For a single layer, a minimum of 6 g powder
sample was collected. Powder samples were obtained from
at least 4–6 different drilled positions in a specimen, in order
to collect the required sample size. All the drillings were
performed on cracked positions, for the powder samples to
be representative of cracked regions in the specimens
. So, the chloride content obtained here is
one-dimensional, and it is the average value of different cracked
positions in a single layer. X-ray fluorescence (XRF) was
used to determine the total chloride content in each 6 g
powder sample. Chemical analysis was also used to
determine both total and free chloride content in each layer. For
these chemical analyses the RILEM TC 178-TMC
(2002a, b) recommendations were followed.
4. Research Results
The slump flow value and mechanical properties of
ultimate compressive strength (fcu) and E-modulus (E-mod) of
SHCC and mortar specimens at the age of 28 days are
shown in Table 3. The ultimate uniaxial tensile strength
(fu,st), first cracking strength (fcr,st) and ultimate tensile strain
(eu,st) of dumbbell shaped specimens at 14 days from
different SHCC mixes are also shown in Table 3. In Table 3 no
significant difference in strength can be observed for SHCC
with fine and coarse sand. However, a significantly reduced
ultimate tensile strain is found for SHCC containing coarse
sand. This may be due to the irregular distribution of fibre
caused by the larger particles of coarse sand in a small
dumbbell section. Also for the particular mix design used in
this study, the target compressive strength of SHCC was in a
range of 25–28 MPa, and a tensile strain capacity of 1–2%.
4.1 Flexural Cracks in the R/SHCC Specimens
Figure 3a, b show the crack properties, including average
crack width (ACW) and maximum crack width (MCW) in
the R/SHCC specimens for different cover depths. Larger
crack widths were found in specimens with C15 and this
trend was clearer at higher level of deflections in the
specimens (Fig. 3b). Although a scatter was noticed for C35
specimens at lower deflection (FS2 and CS2), it appears that
there is no significant difference in crack widths for cover
depths of C25 and C35.
4.2 Corroded Depth in R/SHCC and R/Mortar
Corroded depths calculated from Eq. (5) in different
R/SHCC and R/mortar specimens are shown in Fig. 4a–d. In
all cases specimens with C15 show larger corroded depths.
The higher corrosion in the C15 specimens can be explained
by the shorter distance for both chloride and oxygen to
penetrate to the steel level. For interpretation of the higher
corroded depths of R/FS31 specimens with C15 in Fig. 4b,
the application of higher current (10 mA) for a longer period
(6 mS) should be kept in mind, as mentioned in Sect. 3.4. It
is also worth mentioning that corrosion rate measurements in
the FS2, CS2, FM1 and CM1 specimens were started later,
while others were taken from the beginning. The corroded
depths of these specimens do not reflect the total values
since the dc0 value in Eq. (5) was set to zero at the start of
corrosion rate measuring. No significant difference was
found in the corroded depths for different sand types and
also for C25 and C35 specimens. This can be explained by
the similar strain-hardening behaviour in terms of strength
and cracks in FS2 and CS2 specimens. Further research with
different grain sizes of CS is recommended to come up with
Figures 5a–c show the relationships between uniformly
corroded depths Eq. (5), actual pitting depths measured in
the steel bars, loss of steel bar yield force obtained from
tensile testing, and cover depths in R/FS2 and R/CS2
specimens. The results show a power relationship between
the corroded depths and cover depths. The loss of steel yield
force was found to be exponentially related to the cover
depths in the R/SHCC specimens. It must be kept in mind
that the loss of yield force may depend strongly on position,
size and shape of the pitting corrosion.
4.3 Visual Observation of Corrosion Damage
Corrosion damage in the steel bars in both R/SHCC and
R/mortar specimens was also observed visually after being
removed from the specimens. Different SHCC and mortar
specimens were broken after different durations of chloride
exposure. Interestingly, because of multiple fine cracks, in
most FS2, CS2 and FS31 specimens more distributed
corrosion (also called general corrosion) was observed as it is
shown in Fig. 6. A similar trend was also found in a study by
. This is postulated to be a result of
microcell corrosion, for which it is known that the cathode to
anode area ratio is small, and as a result the severity of
corrosion is reduced. Concentration of chloride ions is likely
not localized in SHCC due to the multiple fine and finely
spaced cracks in SHCC, so the corroded depths caused by
the ions are not deep. In this case the capacity of steel bar is
not reduced significantly. Multiple fine cracks and smaller
crack spacing of SHCC (see Fig. 6a) are believed to be the
main reason for micro-cell corrosion in R/SHCC specimens.
Note that Fig. 6b shows the steel bar before cleaning and
Fig. 6c shows the same steel bar after cleaning with HCl
Contrary to R/SHCC specimens, it appears that macro-cell
corrosion occurred in all R/mortar specimens, with a typical
one shown in Fig. 7. The limited number of cracks and
larger crack widths in mortar specimens (Fig. 7a) are
believed to be the reason for macro-cell corrosion. In this
case the ion concentration is likely localized in the position
of large cracks. In macro-cell corrosion the distance between
anode and cathode is larger, resulting in localized attack of
ions and formation of deeper corrosion. Steel capacity is
clearly more significantly reduced when the corroded depths
are greater. Greater corrosion depths are visible in Fig. 7c
after cleaning the steel bar. Figure 7b shows the steel bar
before cleaning and Fig. 7d shows the same, clean steel bar
surface at the opposite side of cracks.
4.4 Loss of Yield Force and Pitting in Steel Bars
Due to Corrosion
After cleaning the corroded steel bars, the tensile yield
force was also determined in FS32 and FM2 specimens as
reported in Fig. 8. Virgin steel bars were also tested for the
reference yield force capacity (Fy). Equation (6) was used to
find the calculated yield force capacity [Fy (Cal)] of
corroded steel bars. All the steel bars were tested under uniaxial
tensile testing to find the experimental yield force capacity
[Fy (Exp)] which is the remaining yield capacity of the
corroded steel bars. In all cases the yield force capacity from
the tensile test was found to be lower than the calculated
capacity. The reasons include lack of calibration of the
parameters as described in the previous section, but
dominantly the localised nature of actual corrosion as opposed to
the assumption of uniformly distributed corrosion damage in
the calculated values.
Fig. 8 Calculated and experimental yield forces of steel in
R/FS32 and R/FM2.
From Fig. 8 it is interesting to see that the experimental
residual yield force capacity for SHCC specimens with
exactly the same exposure times is greater for N3 and N5
specimens than for N1. Higher standard variations are seen
in the N5 specimens. In case of C15 specimens, 5.4, 4 and
4.7% lower steel yield force were observed in N1, N3 and
N5 specimens in comparison with the virgin steel. Similarly
for C25 specimens, the losses of yield forces were 5, 3.9 and
4% respectively for N1, N3 and N5 specimens than virgin
steel. So, with the limited data set reported here, it is noticed
that there is a trend of increasing retained yield force for
smaller crack spacing. This is in agreement with Fig. 4c, i.e.
smaller corroded depth, more distributed corrosion and
lower pitting depth. Also, a slightly higher retained yield
force is found for the larger cover of 25 mm, than for the
specimens with 15 mm cover. No noticeable difference was
observed in FM2 for different cover depth of the N1
Measured actual pitting depths (dp) in the steel bars are
shown in Figs. 9a–d. Since the corrosion rates of the
specimens were low, the pitting depths were also low. Accurate
measurement of these depths (see Fig. 2b) is complicated by
pitting shape and size, despite the high dial gauge resolution
(10 lm). Because of the sensitivity of measuring these small
pitting depths, not all pitting depths were measured, but the
deeper pitting depths (larger than 10 lm) in the steel bars
were chosen by visual inspection. A minimum of three to a
maximum of six pitting depths on each steel bar were
measured and their average value is reported in Fig. 9, while
the minimum and maximum values of these 3–6 largest
pitting depths are indicated by error bars. It is clear that the
average pitting depths in the FS2 and CS2 specimens that
were exposed for a longer time, are higher compared with
those in the FS31 and FS32 specimens with lower exposure
durations. In SHCC, lower pitting depths were observed for
larger cover depth of the steel bar. An unexpected result is
that D7 specimens had marginally lower pitting depths than
D5 specimens. In FS31, at D5 level, almost 35% and 43 and
at D7 level, almost 66 and 5% lower pitting depths were
observed in C25 and C35 specimens than in C15 specimens.
However, because of the lower corrosion damage in the steel
bars, no strong correlations were found for different sand
type and number of notches in the R/SHCC specimens tested
Pitting area in the steel bars were also measured and
reported in Fig. 10. Because of the difficulty of measuring
the pitting area, steel bars were selected from only FS2, CS2
and FS31 specimens. After cleaning the steel bars, pitting
areas were marked on the steel surface as circular,
rectangular and trapezoidal shapes. A Vernier scale was then used
to measure the dimensions of the marked areas. The pitting
areas on the steel surface were added together per steel bar. It
is clear from Fig. 10 that a larger number of cracks in the
specimens in the gauge length of 200 mm, cause lower total
Figure 11 shows a relationship between average pitting
depths and corrosion-induced loss of tensile yield force in
R/FS31 and R/FS32 specimens. Note that these are the
average values of three specimens for the same cover depth.
An upward trend in loss of yield force with increased pitting
depth is clear. Of importance in rebar yield force is the
reduction in its cross-sectional area, which is clearly not
represented by the pitting depth only, in light of the rather
poor correlations shown in the figure. More accurate
representation of loss of cross-sectional area is required to
improve correlation with, and estimation of corrosion
damage to steel reinforcement. It is also important to note that
the steel bars used in R/FS2 and R/CS2 specimens were bent
to have the U-shape shown in Fig. 6, while in other cases
steel bars were straight. The purpose was to facilitate the
necessary cable connection during the corrosion rate
measurement on these specimens. However, for the tensile
testing the bent portions of these bars were straightened
again. It is reported by the
Choi et al. (2003)
procedure (bend and straighten) can cause significant loss of
yield force (20–25%) of the steel bar. For this reason higher
loss of yield force was found in the steel bars of R/FS2 and
R/CS2 specimens. However, these specimens are not
included in Fig. 11, since the properties of these steel bars
were influenced by the described bending process.
4.5 Chloride Content in the R/SHCC and R/
Total chloride profiles determined by XRF are shown in
Fig. 12a–d. Note that the different types of specimens were
tested after different durations of exposure, which are also
shown in the figures. The typical trend of high chloride
content at the surface, and gradual reduction with depth is
clear in the graphs. From the results found in this research it
can be said that the maximum total chloride content for this
specific matrix can be up to 1.8% of binder weight. The
chloride profile through the whole depth of both cracked and
un-cracked SHCC specimens is also shown in Fig. 12c. For
the un-cracked (UC) chloride profile, powder samples were
collected in 5 mm layer intervals up to a depth of 50 mm on
the specimen surface where there were no cracks. In this
regard, a single notch specimen from FS32 type SHCC was
chosen. It can be seen that the chloride penetration is
significantly lower un-cracked SHCC and from a depth of about
20 mm a low total amount of chloride is found in the
4.6 Relationship Between XRF and Chemical
For the chloride profiling, the XRF method was chosen for
its ease and short duration. To verify the outcome of the
XRF results, chemical analysis according to the RILEM
TC178 recommendation (2002a) was also followed to
determine the total chloride content in the specimens. A total
of 30 different samples which were tested with XRF were
used also in the chemical analysis. Figure 13 shows the
relationship between the XRF and total chloride chemical
test results. A linear correlation is found, which deems XRF
to be a suitable alternative for the determination of total
chloride profiling in SHCC.
Fig. 12 XRF total chloride profile obtained in different R/SHCC and R/mortar specimens up to a depth of 45 mm, at different ages
in weeks (W).
4.7 Total and Free Chloride Content at Different
Depth of Specimens
A total of 42 samples were tested by chemical analysis for
determining the free chloride content according to RILEM
TC178 (2002b). The relationships between the free and total
chloride measurement results from the chemical analysis, as
well as those from XRF total chloride measurement are
shown in Fig. 14. Note that the chloride content showed
there are the average chloride content in the specimens from
the top surface to bottom surface of each individual steel bar.
For instance, chloride content in a C15 specimen steel bar
was determined by taking the average XRF chloride value
from a depth of 15–25 mm reported in Fig. 12. Lower
chloride content (both total and free) was found in the mortar
specimens. This may be explained by the specimens’ lower
total crack width. Although multiple fine cracks were found
in the SHCC specimens, the total crack width in SHCC
specimens were significantly higher than in the mortar
specimens. Another reason is that when drilling was
performed to collect the powder samples of mortar specimens,
only one crack was covered by the drill diameter (16 mm)
while in SHCC 2–4 cracks were covered due to their small
spacing. In this regard, two-dimensional chloride profiling is
envisaged in future work, to improve the characterisation of
chloride content distribution in cracked cement-based
composites. Also, the presence of fibre in the SHCC matrix may
form pores which ultimately help chloride ions to be stored,
and as a result higher free chloride concentrations can be
expected in SHCC than in mortar specimens.
Relationships were also drawn from the chemical testing
of total and free chloride content with respect to the depths
in cracked R/SHCC specimens as shown in Figs. 15a–b.
Both chloride contents appear to form an exponential
relationship with depth in cracked R/SHCC.
In this research the difference between the chemical total
and free chloride at different depths in SHCC specimens was
found to be in a range of 5–65%. In case of NC, this
difference was found to be in a range of 10–53% by
So, the results obtained in this research show a similar trend
in the matrix of SHCC used here with that of NC.
5. Discussion of Results
Typically, the corrosion rate/current is captured by the
electric current (or migration of ions) driven by the
potential differences between the cathodic and anodic areas
(Markeset and Myrdal 2008)
. So, the position of the
cathode and anode in the steel bar is very important. When
the number of cracks is limited and the crack spacing is
large, the distance between anode and cathode area is large
and significant pitting corrosion can form in the steel bar
due to the fact that the smaller anode area in the limited
crack region is fed by the large cathode area. This can be
seen in the Fig. 16, where a trend is found of larger pitting
depths forming when average crack spacing in the
specimen is increased. On the other hand when there are many
cracks and cracks are spaced closer together, a lower
cathode: anode ratio exists. When the cathode area reduces,
less OH- becomes available for the reaction with Fe? in
the anodic area and as a result a lower corrosion rate can
be expected in the steel bar.
Depending on the availability of oxygen and moisture in
the specimen, the corrosion rate of steel can vary with time.
It is also difficult to know whether the steel is in an active or
passive state of corrosion inside concrete, since it is not
possible to see the real corrosion damage in the steel
embedded in the concrete. As a result the assumption of
Stern and Geary constant (B) value may under or
overestimate the real corrosion rate in steel bars. Therefore, the
Fig. 16 Influence of average crack spacing in average pitting
depths in steel bar.
corrosion in steel is indeed a complex process and depends
on many factors.
5.1 Influence of Chloride Content and Mass
Loss of Steel in Corrosion
Figure 17 shows a relationship between the estimated
corroded depths with total and free chloride contents at the
surface of steel bars in the different R/SHCC specimens.
Free chloride content appears to be more representative of
the corrosion of steel bars. Note that in chloride content
testing only selected specimens were used from the different
SHCC reported here. So the chloride content corresponding
to the corroded depths reported in Figs. 4, 5 may not be seen
in Fig. 17.
Measured corrosion rate using the Coulostatic technique,
reported as the calculated corroded depth here, was also
verified with the actual mass loss due to corrosion of steel in
R/FS32 and R/FM2 specimens and the outcome is shown in
Fig. 18. Note that the values reported here are the average of
three SHCC and two mortar specimens for the same number
of notches and for two different cover depths. It appears that
the mass loss of steel is exponentially increased as the
corroded depth (in Fig. 18a) and loss of yield force (in
Fig. 18b) increase in the specimens. It is postulated that the
measured corrosion rate as reported by corroded depths
using the Coulostatic method is relevant since it is expected
that the mass loss of steel increases as the corrosion rate
increases. Note that the exposure period for R/FS32 and
R/FM2 was relatively short (28 weeks) and because of that
the magnitudes of the corroded depth, yield force loss and
mass loss values from the different specimens are small.
However, the difference between the lower and upper values
of corroded depth and yield force loss in Fig. 18a, b are
about 35 and 28% respectively, which are significant.
5.2 Effect of Cracking in Chloride Penetration
The influence of average crack widths on total chloride
penetration measured by XRF was also observed in
R/SHCC, as shown in Fig. 19a, b. No clear relationship
appears to exist between crack width and chloride content
for the average crack widths below 0.05 mm in Fig. 19a. In
Fig. 19b, a trend of higher total chloride with higher average
crack width is apparent. To interpret these apparently
contradicting results, it must be noted that the total exposure
time of the specimens in Fig. 19a was 108 weeks, i.e. near
double the exposure time of those specimens shown in
Fig. 19b. Thereby, it is likely that the influence of crack
widths in the Fig. 19a diminish through gradual chloride
ingress into the matrix from the crack faces subsequent to
quick ingress into the cracks. Also, two dimensional chloride
profiling will be used in future work, in order to more
accurately relate crack width and chloride concentration in
cracked SHCC. Chloride ingress is also related to the matrix
diffusivity. Furthermore, it should also be kept in mind that
two different chloride exposures of capillary absorption and
ponding (see Sect. 3.4) might influence the chloride profiles
in the respective specimens.
This research work has revealed some important
parameters which are necessary for durability modelling of chloride
induced corrosion in cracked SHCC. The Coulostatic
technique used here is found to reasonably represent mass loss of
steel due to corrosion, although no attempt to improve
calibration has been made here. In terms of the corrosive
deterioration for the observed and enforced (notched) crack
properties in R/SHCC, the following conclusions can be
drawn from the experimental work of this research:
Average and maximum crack widths and crack spacing
in the R/SHCC are smaller for larger cover depth of the
steel bar. From the particular type of SHCC used in this
research, it appears that a 25 mm cover depth is the
threshold cover depth for limiting the crack width in
• Mass loss, pitting depth and loss of yield force are
considered to be low in all specimens, despite up to
108 weeks of cyclic wetting and drying chloride
exposure. After 108 weeks of such exposure, a maximum
pitting depth of 1.4 mm was found in the 10 mm
diameter steel bars, with the average pitting depth in the
range 0.1–0.5 mm. For lower exposure durations, the
average pitting depths were lower at 0.1–0.4 mm after
57 weeks, and 0.1–0.3 mm after 28 weeks of cyclic
chloride exposure. This led to a maximum loss of yield
force of the bars of about 17% in a CS2_C15B1
• In chloride-induced corrosion performed here, higher
corroded depths (measured by a Coulostatic method),
actual measured pitting depths, and higher loss of yield
force in the steel were found in the specimens with cover
depth of 15 mm than for 25 and 35 mm cover depths. No
significant difference was observed in the specimens
with 25 and 35 mm cover depths.
• A significant reduction in tensile strain capacity was
found for SHCC produced with coarse sand (CS2)
compared with fine sand. However, there were no
significant differences in the pitting depths and loss of
yield force of steel in coarse sand specimens CS2 and
CM1 than in fine sand specimens FS2 and FM1.
• Free chloride content in the specimen at the level of the
steel bar appears to correlate better with the corrosion
damage than total chloride content. However, while the
role of chloride in corrosion initiation has been studied
widely, its role in corrosion propagation and corrosion
rate must be investigated further.
• The XRF method can be an alternative method for
chemical testing in determining total chloride content in
SHCC. Both total and free chloride content reduced with
depth into the specimens and the difference between
chemical total and free chloride content was found to be
in the range of 5–65%, depending on the depth in the
The actual mass loss of steel bars is related to the
corroded depths and loss of yield force of R/FS32 and
R/FM2 specimens. The pitting depth in the steel bars was
larger for larger average crack spacing in the R/SHCC
A larger number of cracks, associated with finer crack
spacing, lead to significantly lower corrosion damage in
This article is distributed under the terms of the Creative
Commons Attribution 4.0 International License (http://
creativecommons.org/licenses/by/4.0/), which permits un
restricted use, distribution, and reproduction in any medium,
provided you give appropriate credit to the original author(s)
and the source, provide a link to the Creative Commons
license, and indicate if changes were made.
Altoubat , S. , Maalej , M. , & Shaikh , F. U. A. ( 2016 ). Laboratory simulation of corrosion damage in reinforced concrete . International Journal of Concrete Structures and Materials , 10 ( 3 ), 383 - 391 .
Andrade , C. , & Alonso , C. ( 2004 ). Test methods for on-site corrosion rate measurement of steel reinforcement in concrete by means of the polarization resistance method . Materials and Structure , 37 , 623 - 643 .
Angst , U. , Elsener , B. , Larsen , C. K. , & Vennesland , Ø. ( 2009 ). Critical chloride content in reinforced concrete-a review . Cement and Concrete Research , 39 ( 12 ), 1122 - 1138 .
Bashir , H. , Osman , B. H. , Wu , E. , Ji , B. , & Abdulhameed , S. S. ( 2017 ). Repair of pre-cracked reinforced concrete (RC) Beams with openings strengthened using FRP sheets under sustained load . International Journal of Concrete Structures and Materials , 11 ( 1 ), 171 - 183 .
Blagojevic´ , A. ( 2016 ). The Influence of Cracks on the Durability and Service Life of Reinforced Concrete Structures in relation to Chloride-Induced Corrosion: A Look from a Different Perspective . PhD Thesis , Delft University of Technology, Delft, The Netherland.
Broomfield , J. P. ( 2007 ). Corrosion of steel in concrete understanding, investigating and repair . Book 2nd edition , Taylor & Francis, USA & Canada.
Choi , H. , Kim , H. , Seo , D. , & Kang , K. ( 2003 ). The study on the capacity transform and alternative plan of reinforcing bar with straightening after bending . Journal of the Architectural Institute of Structural Systems , 19 ( 9 ), 181 - 188 .
Djerbi , A. , Bonnet , S. , Khelidj , A. , & Baroghel-Bouny , V. ( 2008 ). Influence of traversing crack on chloride diffusion into concrete . Cement and Concrete Research , 38 , 877 - 883 .
German , M. , & Zaborski , A. ( 2011 ). Numerical analysis of chloride corrosion of reinforced concrete . Technical Transections , 3 , 47 - 59 .
Glass , G. K. ( 1995 ). An assessment of the coulostatic method applied to the corrosion of steel in concrete . Corrosion Science , 37 ( 4 ), 597 - 605 .
Gonzalez , J. A. , Cobo , A. , Gonzalez , M. N. , & Feliu , S. ( 2001 ). On-site determination of corrosion rate in reinforced concrete structures by use of galvanostatic pulses . Corrosion Science , 43 ( 4 ), 611 - 625 .
Hausmann , D. A. ( 1967 ). Steel corrosion in concrete . Materials Protection , 6 , 19 - 23 .
Huang , Q. ( 2006 ). Influence of cracks on chloride-induced corrosion in reinforced concrete structures . MSc thesis , Chalmers University of Technology, Sweden.
Kobayashi , K. , Iizuka , T. , Kurachi , H. , & Rokugo , K. ( 2010 ). Corrosion protection performance of high performance fiber reinforced cement composites as a repair material . Cement and Concrete Composite , 32 , 411 - 420 .
Kobayashi , K. , & Rokugo , K. ( 2013 ). Mechanical performance of corroded RC member repaired by HPFRCC patching . Construction and Building Materials , 39 , 139 - 147 .
Li , V. C. ( 2012 ). Tailoring ECC for Special Attributes: A Review . International Journal of Concrete Structures and Materials , 6 ( 3 ), 135 - 144 .
Liang , M. T. , Huang , R. , Feng , S. A. , & Yeh , C. J. ( 2009 ). Service life prediction of pier for the existing reinforced concrete bridges in chloride-laden environment . Journal of Marine Science and Technology , 17 ( 4 ), 312 - 319 .
Liu , Y. ( 1996 ). Modelling the Time-to-Corrosion Cracking of the Cover Concrete in Chloride Contaminated Reinforced Concrete Structures . PhD thesis , Virginia Polytechnic Institute and State University, USA.
Markeset , G. , & Myrdal , R. ( 2008 ). Modelling of reinforcement corrosion in concrete- State of the art , COIN Project report no7, SINTEF Building and Infrastructure , ISBN 1891 - 1978 .
Mihashi , H. , Ahmed , S. F. U. , & Kobayakawa , A. ( 2011 ). Corrosion of reinforcing steel in fibre reinforced cementitious composites . Journal of Advanced Concrete Technology , 9 ( 2 ), 159 - 167 .
Mohammed , T. U. , Otsuki , N. , Hisada , M. , & Shibat , T. ( 2001 ). Effect of crack width and bar types on corrosion of steel in concrete . Journal of Materials in Civil Engineering , 13 , 194 - 201 .
Pacheco , J. , & Polder , R. B. ( 2016 ). Critical chloride concentrations in reinforced concrete specimens with ordinary Portland and blast furnace slag cement . HERON , 61 ( 2 ), 99 - 119 .
Paul , S.C. ( 2015 ). The role of cracks and chlorides in corrosion of reinforced strain-hardening cement-based composites (R/SHCC), PhD Thesis , Stellenbosch University, Stellenbosch, South Africa.
Paul , S. C. , Babafemi , A. J. , Conradie , K. , & van Zijl, G. P. A. G. ( 2017 ). Applied voltage on corrosion mass loss and cracking behaviour of steel reinforced SHCC and mortar specimens . Journal of Materials in Civil Engineering . doi: 10 .1061/(ASCE)MT. 1943 - 5533 . 0001807 .
Paul , S. C. , & van Zijl, G. P. A. G. ( 2013 ). Mechanically induced cracking behaviour in fine and coarse sand strainhardening cement based composites (SHCC) at different International Journal of Concrete Structures and Materials (Vol. 11 , No.3, September 2017 ) | 571 load levels . Journal of Advanced Concrete Technology , 11 , 301 - 311 .
Paul , S. C. , & van Zijl, G. P. A. G. ( 2014 ). Crack formation and chloride induced corrosion in reinforced strain hardening cement-based composite (R/SHCC) . Journal of Advanced Concrete Technology , 12 , 340 - 351 .
Paul , S. C. , & van Zijl, G. P. A. G. ( 2016 ). Chloride-induced corrosion modelling of cracked reinforced SHCC . Archives of Civil and Mechanical Engineering , 16 ( 4 ), 734 - 742 .
Pettersson , K. ( 1993 ). Corrosion of steel in high performance concrete . In Proceedings of the 3rd International Symposium on Utilization of High Strength Concrete , Lillehammer, Norway (published by the Norwegian Concrete Association) .
Rilem , T. C. ( 2002a ). Testing and modelling chloride penetration in concrete. Analysis of total chloride content in concrete . Materials and Structure , 35 , 583 - 585 .
Rilem , T. C. ( 2002b ). Testing and modelling chloride penetration in concrete. Analysis of water soluble chloride content in concrete . Materials and Structure , 35 , 586 - 588 .
Sahmaran , M. , Li , V. C. , & Andrade , C. ( 2008 ). Corrosion resistance performance of steel reinforced engineered cementitious composite beams . ACI Materials Journal , 105 ( 3 ), 243 - 250 .
Schiessl , P. , & Raupach , M. ( 1997 ). Laboratory studies and calculations on the influence of crack width on chlorideinduced corrosion of steel in concrete . ACI Materials Journal , 94 ( 1 ), 56 - 62 .
Soltani , A. , Harries , K. A. , & Shahrooz , B. M. ( 2013 ). Crack Opening Behavior of Concrete Reinforced with High Strength Reinforcing Steel . International Journal of Concrete Structures and Materials , 7 ( 4 ), 253 - 264 .
Tang , L. , Utgenannt , P. , & Boubitsas , D. ( 2015 ). Durability and Service Life Prediction of Reinforced Concrete Structures . Journal of the Chinese Ceramic Society , 43 ( 10 ), 1408 - 1419 .
Wang , K. , Jansen , D. C. , Shah , S. P. , & Karr , A. F. ( 1997 ). Permeability study of cracked concrete . Cement and Concrete Research , 27 ( 3 ), 381 - 393 .
Wittmann , F. H. , Wang , P. , Zhang, P. , Zhao , Tie-Jun. , & Betzung , F. ( 2011 ). Capillary absorption and chloride penetration in neat and water repellent SHCC under imposed strain . Paper presented at the 2nd International RILEM Conference on Strain Hardening Cementitious Composites, Brazil , pp. 165 - 172 .
Yoo , D. Y. , & Yoon , Y. S. ( 2016 ). A review on structural behavior, design, and application of ultra-high-performance fiber-reinforced concrete . International Journal of Concrete Structures and Materials , 10 ( 2 ), 125 - 142 .