Stainless Steel Bonded to Concrete: An Experimental Assessment using the DIC Technique
International Journal of Concrete Structures and Materials
Stainless Steel Bonded to Concrete: An Experimental Assessment using the DIC Technique
Hugo Biscaia 0
Noel Franco 0
0 , and Carlos Chastre
The durability performance of stainless steel makes it an interesting alternative for the structural strengthening of reinforced concrete. Like external steel plates or fibre reinforced polymers, stainless steel can be applied using externally bonded reinforcement (EBR) or the near surface mounted (NSM) bonding techniques. In the present work, a set of single-lap shear tests were carried out using the EBR and NSM bonding techniques. The evaluation of the performance of the bonding interfaces was done with the help of the digital image correlation (DIC) technique. The tests showed that the measurements gathered with DIC should be used with caution, since there is noise in the distribution of the slips and only the slips greater than one-tenth of a millimetre were fairly well predicted. For this reason, the slips had to be smoothed out to make it easier to determine the strains in the stainless steel and the bond stress transfer between materials, which helps to determine the bond-slip relationship of the interface. Moreover, the DIC technique allowed to identify all the states developed within the interface through the load-slip responses which were also closely predicted with other monitoring devices. Considering the NSM and the EBR samples with the same bonded lengths, it can be stated that the NSM system has the best performance due to their higher strength, being observed the rupture of the stainless steel in the samples with bond lengths of 200 and 300 mm. Associated with this higher strength, the NSM specimens had an effective bond length of 168 mm which is 71.5% of that obtained for the EBR specimens (235 mm). A trapezoidal and a power functions are the proposed shapes to describe the interfacial bond-slip relationships of the NSM and EBR systems, respectively, where the maximum bond stress in the former system is 1.8 times the maximum bond stress of the latter one.
stainless steel; concrete; bond failure; digital image correlation
The first studies on the external bonded reinforcement
(EBR) technique using steel plates were carried out in the
late 1960s in France by L’Hermite and Bresson, who
analyzed the steel-epoxy-concrete connection
Bresson 1967; L’Hermite 1977)
. Since then, there have been
many studies characterizing the bonding behaviour of
strengthening elements using the EBR technique, initially
with steel plates
(Ladner 1978, 1983; Jones et al. 1980;
Swamy and Jones 1980; Chastre Rodrigues 1993; Ta¨ljsten
and more recently with fibre reinforced polymers
(Blaschko and Zilch 1999; De Lorenzis et al. 2000;
De Lorenzis and Teng 2007; Lorenzis et al. 2001; Nakaba
et al. 2001; Chen et al. 2005; Aiello and Leone 2008;
Martinelli et al. 2011; Dehghani et al. 2012; Biscaia et al.
. As for the near surface mounted (NSM)
technique, although there are some references of its use, only at
the end of the 1990s did studies on the performance of this
technique associated with the use of FRP rods
Zilch 1999; De Lorenzis et al. 2000)
begin to appear.
Therefore, in most of the studies that can be found in the
(e.g. Xia 2005; Akbar et al. 2010; Smith 2010;
Wan 2010; Wan et al. 2014; Biscaia et al. 2016a, b, 2017b)
the focus is on bonded joints between FRP composites and
concrete but recently there have been more studies on both
steel (only with EBR technique) and timber structures (with
both EBR and NSM techniques). Still, in reinforced concrete
(RC) structural strengthening, stainless steel (SS) is a
possible alternative to mild steel or FRP composites due to its
durability. Compared to mild steel, the durability
performance of SS is higher, but there is no significant difference
between them in terms of weight-strength ratio. However,
despite its lower weight/strength ratio, stainless steel has
ductile behaviour, and good corrosion resistance which are
important and decisive factors in choosing it for
strengthening structures instead of FRP composites. Nevertheless, a
significant and common drawback, whatever the bonded
materials are, is the premature debonding of the material
used in the bonding strengthening technique. Several authors
have been studying the premature debonding phenomenon
on FRP composites and concrete joints
(e.g. Arduini et al.
1997; Neubauer and Rosta´sy 1997; Bizindavyi and Neale
1999; Harmon et al. 2003; Smith and Teng 2002; Yao et al.
2005; Teng et al. 2006; Wu and Yin 2003)
, FRP and timber
(e.g. Smith 2010; Wan 2010; Wan et al. 2014; Biscaia
et al. 2016a, b, 2017)
, FRP and steel joints
(e.g. Xia and
Teng 2005; Akbar et al. 2010; Fawzia et al. 2006; Wang
et al. 2016; Yu et al. 2012; Al-Mosawe et al. 2015; Fernando
et al. 2014)
or steel and concrete joints
(e.g. L’Hermite and
Bresson 1967; L’Hermite 1977; Ladner 1978; Ladner 1983;
Jones et al. 1980; Swamy and Jones 1980; Chastre
Rodrigues 1993; Ta¨ljsten 1997; Gomes and Appleton 1999; Van
Germet 1990; Aykac et al. 2013)
Therefore, researchers have studied the debonding
phenomenon between two bonded materials through different
approaches whether they are experimental, analytical,
numerical or embracing part or all of these three procedures.
Furthermore, the test setup configuration assumed for this
kind of study may vary considerably
(Wu et al. 2002)
the most commonly used configurations being the
doublelap pull or push shear tests, the single-lap pull tests, the
double strap tests or the 3-point bending tests. Independently
of the procedure followed, researchers seem to be fairly
unanimous that the debonding failure process of a
structurally bonded joint can be analyzed and predicted through
the relationship between the interfacial bond stress and the
slip (i.e. the relative displacement between bonded
Unlike FRP composites, stainless steel has a more
complex constitutive behaviour, which may change or even
eliminate conventional ways of finding the bond–slip
relationship. Typically, the bond stresses and the slips are
experimentally found in the data collected from strain
gauges that, before the testing of the samples, were bonded on
the strengthening material along their bond length. Thus, to
determine the bond stresses, it is assumed that the bond
stresses developed between two consecutive strain gauges
are constant. To determine the slips, it is first assumed that
the strains in the concrete are zero and, by integrating the
strains with respect to (and along) the bonded length, the
slips within the interface can then be calculated.
Dai et al.
proposed an alternative procedure, which eliminates
the need to use strain gauges. To determine the interfacial
bond–slip relationship, they said that knowing only the
displacement at the most loaded end of the strengthening
material and the load transmitted to that material is enough
to obtain the bond–slip relationship. In both cases, a
sufficiently long bond length should be considered, but the
former method makes the process more expensive because it
requires the use of several strain gauges that should not be
bonded too far away from each other in order to obtain
A more recent alternative to obtain the slips developed
within a bonded joint is digital image correlation (DIC),
which allows the monitoring of an entire surface
et al. 2016)
instead of a single point, as conventional strain
gauges do. Once again, researchers have been studying the
bond between an FRP composite and concrete
Martinelli et al. 2011; Czaderski et al. 2010; Cruzet al. 2016;
Ghiassi et al. 2013; Zhu et al. 2014)
. The use of commercial
DIC techniques is quite expensive but, nowadays, its use
became very economical due to the powerful digital cameras
currently available on the market, plus the free software that
can be easily found on the web
(e.g. Wang and Vo 2012;
http://www.ncorr.com/; GOM Correlate)
to perform the DIC
analysis. However, the reliability of these free software for
the assessment of the debonding phenomenon between
stainless steel and concrete was not demonstrated so far but
its use is unlimited and besides that, to start monitoring
laboratory structures under loading, only the initial cost of
the digital camera and the corresponding free software
installed in a laptop, is needed.
Although some work suggests that the DIC technique can
be used to determine the interfacial bond–slip relationship of
(Ghiassi et al. 2013; Zhu et al.
, the bond stresses are generally smoothed through a
mathematical function that predicts the slips or the strain
distributions in the FRP composite. This procedure bypasses
the difficulties with obtaining a smoothed displacement result
from the DIC technique and the slips fluctuate instead (Zhu
et al. 2014). Therefore, when it comes to determining the
strains and especially the bond stresses in the interface, the
fluctuations in the slip distributions are amplified, which
increases the error in calculating the interfacial bond–slip
relationship. For this reason, a smoothed and previously
known function of the slips or strain distributions is used
instead of the real ‘‘peaks and valleys’’ obtained from the DIC
technique. However, it is important to note that determining
the interfacial bond–slip relationships with DIC can only be
viable if the results gathered by using the DIC technique can
reproduce the same bond–slip relationship accurately enough
and on its own, as the results obtained from other means, such
as those reached by the two procedures mentioned earlier.
Knowing the interfacial bond–slip relationship within the
stainless steel and the concrete due to their bonding with an
adhesive is important because it will open up the possibility
to analyze and study, using an analytical or a numerical
approach, the debonding failure between the stainless steel
and concrete by means of either a closed-form solution or by
an approximation procedure, respectively.
However, to the best knowledge of the authors, there are
no studies covering the bond behaviour of the stainless steel
bonded to concrete and therefore, the present work aims to
present an experimental study in which the bonded interface
between the stainless steel and the concrete is tested with
different bond lengths. Furthermore, stainless steel strips and
rods were used. The stainless steel strips were bonded on the
surface of several RC samples using the Externally Bonded
Reinforcement (EBR) bonding technique, whilst the rods
were bonded into a groove previously made on the surface of
the RC samples as per the Near Surface Technique (NSM).
To monitor the strains in the EBR specimens, several strain
gauges were used. In order not to affect the bonded area of
the NSM specimens, no strain gauges were used in these
samples. In both cases, the DIC technique was used and its
viability was checked with the EBR specimens only. In some
cases, only the most loaded bonded region could be
monitored with the DIC technique due to the range of the bonded
lengths covered in this work (between 50 and 800 mm).
Still, throughout the duration of the tests the slips were quite
accurate compared with those obtained from the strain
gauges and the load–slip response at the stainless steel
loaded end was sufficiently reproduced using the DIC
technique. Moreover, the slip distributions observed with the
DIC technique showed similarities to those obtained from
the strain gauges, despite some fluctuations, i.e. with higher
slips at the SS loaded end and decreasing towards the SS free
end. However, as initially suspected, the differences between
the strains from the DIC technique and the strain gauges
increased throughout. The interfacial bond–slip relationship
between the stainless steel and concrete of the EBR
specimen was then determined from the strain gauges bonded on
the stainless steel strips. The results obtained from the EBR
specimens, allowed a first attempt to be shown to represent
and qualitatively identify the bond–slip relationship of the
NSM specimens. Nevertheless, it was also found that the
NSM specimens and the stainless steel rods performed better
due to the rupture of the stainless steel rods when the bonded
length was equal to or higher than 200 mm long.
2. Experimental Program
In order to evaluate the Mode II bond transfer between
stainless steel (SS) and concrete, an experimental program
including two different bonding techniques was idealized.
The Externally Bonding Reinforcement (EBR) and the near
surface mounted (NSM) techniques were herein considered.
Several bond lengths were tested and their influence on the
strength of the interface was analyzed. Table 1 shows all the
tests carried out as well as the designation of the specimens
given to each one. Additionally, the instrumentation used on
each test is also briefly mentioned in Table 1.
2.1 Mechanical Properties of the Materials
The specimens used for the present experimental program
were taken from RC T-beams previously tested to a 4-point
(Franco and Chastre 2016; Chastre et al.
. The regions of the beams with negligible
bending moments, i.e. at the vicinities of the supports
(pinrolled), were used and the stainless steel was bonded at the
bottom region of the flange of the T-beam with their ends
free of any additional mechanical anchorages. This
procedure ensures that the concrete used was not sufficiently
tensioned to develop or even initiate any cracks that could
affect the results obtained now from the single-lap shear
tests. The strength of the concrete was evaluated at 28 days
of age and 3 concrete cubes were subjected to uniaxial
compression until failure accordingly to the standard NP EN
12390-3 (CEN 2003). The results allowed the average
maximum compression stress of the concrete to be
calculated by fcm = 24.1 MPa, which represents, accordingly to
Eurocode 2 (Eurocode 2 (EC2) 1992), a C20/25 concrete.
The steel reinforcements of the RC T-beams were also
tested under uniaxial tension and its mechanical properties
are briefly reported in Table 2. More details about the tests
of the steel reinforcements can be found elsewhere
and Chastre 2016; Chastre et al. 2017; Chastreet al. 2016)
The mechanical properties of the stainless steel were also
determined from the tensile tests carried out on 7 strips with
a cross section of 20 9 5 mm (width 9 thickness) and from
6 tests on rods with 8 mm diameter, according to the
European standard EN ISO 6892-1
. The results
obtained from these tests are shown in Table 2.
For the bonding of the stainless steel to the concrete, an
epoxy resin was used with the commercial designation S&P
Resin 220. The mechanical properties of the epoxy resin
given by the supplier were herein considered
, i.e. compression strength higher than 70 MPa,
shear strength higher than 26 MPa, Young modulus higher
than 7.1 GPa, bond stress when used with concrete and at
20 C higher than 3 MPa and bond stress to steel at 20 C
higher than 14 MPa (after 3 days).
The yielding point of the stainless steel strips is not clearly
identified due to its constitutive nonlinearities, existing at a
low strain level, whereas the constitutive behaviour of the
stainless steel rods is elastic–plastic with a yielding point
very clear and easy to identify. Therefore, the constitutive
behaviour of the stainless steel strips was approximated to
the Ramberg–Osgood relationship
ess ¼ Ess þ a
where a and n are constants obtained from experimental
tensile test of the stainless steel strip and r0 is the axial stress
in the stainless steel at 0.2% strain. Hence, in the present
study, the values determined for a and n are 0.05 and 9.7,
respectively. Figure 1 shows the stress–strain relationships
of the stainless steel in strips and rods obtained from the
simple tensile tests carried out in a universal tensile machine
with a capacity of 100 kN.
2.2 Geometry and Preparation of the Specimens
As mentioned earlier, for the single-lap shear tests, the
specimens were based on short lengths of RC T-beams and
stainless steel bonded on uncracked concrete regions. The
cross section and the corresponding geometry of the
T-beams are shown in Fig. 2a, whereas Fig. 2b shows the
cross section of the specimen in which the EBR technique
was used. It should be noted that all of the specimens had the
same concrete covering of 20 mm and the stirrups were
spaced between them at every 150 mm. The reinforced
concrete had a total height of 305 mm, in which 105 mm
Bond length, Lb (mm)
2 LVDTa, 3 SGb, 1 DCc
2 LVDTa, 4 SGb, 1 DCc
2 LVDTa, 3 SGb, 1 DCc
2 LVDTa, 5 SGb, 1 DCc
2 LVDTa, 7 SGb, 1 DCc
2 LVDTa, 8 SGb, 1 DCc
2 LVDTa, 11 SGb, 1 DCc
2 LVDTa, 15 SGb, 1 DCc
2 LVDTa, 16 SGb, 1 DCc
2 LVDTa, 21 SGb, 1 DCc
2 LVDTa, 1 DCc
2 LVDTa, 1 DCc
2 LVDTa, 1 DCc
2 LVDTa, 1 DCc
2 LVDTa, 1 DCc
2 LVDTa, 1 DCc
form to the flange. The flange was 405 mm wide and the
web measured 150 mm wide. The stainless steel was always
bonded along the mid line that equally divides the flange in
half. For the EBR system, the concrete surface was
pretreated with grinder, whereas for the NSM system, a small
groove with 12 mm deep was made in the concrete surface
in order to insert the stainless steel rod. The stainless steel
strips and rods were all pre-treated with wire brush and then
cleaned with compressed air and acetone before starting to
bond the SS to the concrete in order to remove contaminants
on the surface of the stainless steel (e.g. oil, grease, water,
(Fernando et al. 2013)
The EBR system was completed when the epoxy resin was
placed on the concrete surface along the bond length and the
stainless steel positioned on the resin. For the NSM system,
the groove was filled with epoxy resin and then the stainless
rod was introduced into the groove. These operations were
repeated for all the specimens herein considered.
2.3 Measurements and Procedures Followed
During the Tests
The single-lap shear test was the configuration used in the
current work. Figure 3 shows an overview of the test setup
adopted in this study. Actually, this configuration was
previously used in the work developed in references
et al. 2016a, b, 2017b)
for the analysis of carbon fibre
reinforced polymers (CFRP) bonded to other structural
materials such as concrete, steel and timber allowing the
necessary and sufficient experimental data for the evaluation
of the bond stress transfer in these joints to be collected. This
test apparatus consists of a steel frame where a hydraulic
jack is installed. A small steel profile is placed at the rear of
the hydraulic jack providing the reaction needed when the
SS is pulled out. A pressure cell with a maximum capacity of
200 kN was placed in the front of the hydraulic jack (see
Detail A in Fig. 3). A mechanical anchorage device
consisting of a hollow metallic cylinder with two-piece anchor
wedges was installed at the front of the pressure cell (see
Details A and C in Fig. 3). This device proved not to be
sufficient to ensure that the SS would not slip inside the
metallic cylinder, between the two anchor wedges, when the
hydraulic jack started to push it out. Therefore, another
metallic device with two metallic bolts was placed in front of
the hollow metallic cylinder. The metallic bolts, when
attached, were efficient because they prevented the
twopiece anchor wedges from slipping inside the cylinder
allowing the loads to be transmitted to the SS-to-concrete
Along the bonded length, several strain gauges
TMLFLA-5-17-5L were bonded to the SS strips. Two Linear
Variable Displacement Transducers (LVDT), were placed at
both edges of the interface. One measured the displacements
at the SS loaded end (see Detail B in Fig. 3), and the other
one measured the displacements at the SS free end. A data
logger was used to collect and send all the data to a desktop
Furthermore, a spray paint with a granite speckle effect
was used to paint the bonded area of the monitoring area.
Figure 4 shows the concrete surface before and after
spraying the bonded area to be monitored during the test. A
digital camera captured photos with 3456 9 5184 pixels at
intervals of 5 s during the test. In order to avoid undesired
shadows in the pictures, a 100 W artificial spotlight was
used. The digital camera was synchronized with the other
monitoring devices such as the LVDT and strain gauges.
This synchronization allowed the DIC technique to be
examined to see if it produces sufficiently accurate results
when compared to the monitoring equipment and to test its
feasibility for evaluating the debonding process of the
SS-toconcrete interface. Therefore, the relative displacements
between bonded materials, whether measured by the DIC
technique or calculated using the strains gauges, combined
with the loads measured through the pressure cell installed at
the front of the test setup, allowed the load–slip response to
be obtained for stainless steel bonded to concrete, which is a
very important relationship for the understanding of the
debonding failure process between two bonded materials,
e.g. Biscaia et al. 2013a, b, 2016, 2017b; Dehghani et al.
2012; Caggiano et al. 2012;
Carrara et al. 2011
The commercial GOM Correlate software was used to
measure the displacements of the painted area. Figure 5
shows, as an example, the displacements measured with the
GOM Correlate software of specimen SS-EBR-L640 and
SS-NSM-L300 at four different stages of the load–slip
response obtained from each sample. Figure 5 clearly shows
the range of displacements measured along the bond length,
the SS loaded end being the region with the highest
displacements, whilst the other end registered smaller
3. Failure Modes and Rupture Loads
The common failure modes observed from the single-lap
shear tests are briefly shown in Fig. 6. In total, five different
failure modes were observed and were classified as follows:
(i) adhesive rupture of the stainless steel-to-resin interface
(Type I); (ii) cohesive rupture within a surface layer of the
concrete (Type II); (iii) mixed rupture, i.e. cohesive in
concrete and adhesive within the SS-to-adhesive interface (Type
III); (iv) cohesive rupture within the concrete (Type IV); and
(v) rupture of the stainless steel rod (Type V).
Mostly, the rupture observed in the EBR samples with
shorter bond lengths was interfacial between the SS strip and
the epoxy resin. However, as the bond length in these
specimens increased, the ruptures began to occur within a
thin layer of concrete. In the NSM specimens tested, the
failure modes observed with shorter or longer bond lengths
were quite different. The failure mode detected in the
samples with shorter bond lengths were all cohesive within the
concrete, whereas the rupture of the SS rod was observed in
the two specimens with the longest bond length, i.e. with
200 and 300 mm. Comparing the EBR and the NSM
techniques, the rupture of the SS observed in the NSM technique
shows that this is more efficient than the EBR technique.
Despite being beyond the scope of this study, the failure
modes herein observed show that an improvement to the
EBR technique must be considered in the future. Amongst
other possible solutions for increasing the bond strength
capacity between the stainless steel and concrete, the
installation of mechanical fasteners or adopting other
(Almeida et al. 2016)
considered. Of course, the best solution for achieving this would be
one that is able to maximise the full mechanical behaviour of
the SS strip. In other words, the ideal solution is the one that
leads to the rupture of the strip. This has been achieved in
recent studies with a new bonding technique designated as
Continuous Reinforcement Embedded at Ends (CREatE)
which was developed by the authors with other reinforcing
(Biscaia et al. 2016c, 2017)
and it consists to
embed both free ends of the reinforcing material into the
Table 3 presents the rupture modes observed and the
rupture loads reached in each tested specimen. In Table 3 it
can be seen that the rupture loads associated to the EBR
technique tend to increase with the bond length and in the
cases where this doesn’t occur the ruptures modes are mixed
modes, i.e. parts of the bond length had adhesive failure
within the SS-to-resin interface and other parts of the bond
length ruptured within a superficial layer of concrete.
Therefore, cohesive ruptures within the concrete are most
efficient because, as should be expected from a bonding
technique, the adhesive interfaces cannot be the weakest link
in the bonding between two materials and the rupture should
take place in one of the two bonded materials instead. For
this reason, the NSM technique was considered the one that
led to the best interface performance because the rupture of
the SS rod was reached when a sufficient bond length was
considered, i.e. 200 and 300 mm, with 48.9 kN (973 MPa)
and 47.8 kN (951 MPa), respectively.
Fig. 5 Displacement field obtained from the DIC software at different points of the load–slip responses of the specimens: a
SSEBR-L640; and b SS-NSM-L300.
4. Accuracy of the DIC Technique
4.1 DIC vs. Strain Gauge-Based Measurements
The strains developed in the stainless steel strips bonded to
the concrete accordingly to the EBR technique were all
collected from the strain gauges bonded along the bond
length. Since the strains developed in the stainless steel are
much larger than those developed in the concrete, the strains
developed in the concrete can be ignored. Therefore, the
slips were determined based on the data collected from the
strain gauges according to
(Biscaia et al. 2013; Ferracuti
et al. 2007)
sðxÞ ¼ sðxiþ1Þ
where x corresponds to the axis parallel to the bond length;
(ei?1 - ei) and (xi?1 - xi) are, respectively, the strain and
the distance between two consecutive strain gauges.
Thus, the slips obtained at the SS loaded end are either
from Eq. (2) or from the DIC technique during the tests, i.e.
the slip at x = Lb versus duration of the test is presented in
Fig. 7. As can be seen from this figure, the accuracy of the
DIC technique with the slip derived from the strain gauges is
quite remarkable, especially in those specimens with larger
bond lengths. However, a non-smooth slip distribution can
be easily seen from the specimens with shorter bond lengths,
i.e. with a bonded length shorter than the effective bond
length, which induces a relevant noisy signal in determining
the strains in the stainless steel.
In terms of relative displacements between materials, the
Absolute Deviation (AD) and the Mean Absolute Deviation
(MAD) between the DIC technique and the slips obtained
from the strain gauge measurements were calculated in each
test according to:
AD ¼ X s0D;IiC
MAD ¼ n
where s0 and s0DIC are the slips measured at x = 0 obtained
from Eq. (2) and from the DIC technique, respectively; and
n is the number of measurements carried out during the test.
The results showed that the specimens with the shortest bond
lengths have the lowest MAD, whereas the specimens with
the longest bond lengths have the highest MAD. Thus, the
lowest MAD was found in specimen SS-EBR-L50 with a
calculated MAD of 0.002 mm and specimen SS-EBR-L560
had the highest MAD of 0.019 mm.
In terms of relative errors, the Absolute Percent Error
(APE) and the Mean Absolute Percent Error (MAPE) were
also determined according to:
APE ¼ 100
MAPE ¼ n
The results showed that the MAPE tends to decrease with
the increase of the bond length adopted for the specimens.
However, it is important to keep in mind that these results
are very scattered but still, some tests showed that when the
displacements increased the values for the MAPE tended to
decrease. This may indicate that the use of the DIC
technique could be used if the displacements to be measured are
not too small. Therefore, the use of the DIC technique may
require some prudence and/or methodologies that should be
considered for the experimental assessing of the bond
between SS and concrete. In the following sections those
aspects are highlighted and developed with the help of the
samples initially considered in this work.
4.2 Load–Slip Response
The load–slip response of the stainless steel bonded to
concrete will have different characteristics, depending if the
bond length is longer or shorter than the effective bond
length. Thus, in order to help with the description of those
differences, Fig. 8 shows all the load–slip responses
obtained from the EBR specimens. Moreover, the slips
obtained from Eq. (1) and from the DIC technique are also
shown in Fig. 8. As already mentioned and previously
demonstrated, the slips at the SS loaded end obtained either
from the strain gauges or from the DIC technique are much
alike. However, it seems that the slips measured from the
DIC technique tend to deviate from the slips measured from
the strain gauges in those EBR with SS specimens with the
shortest bonded lengths. This was probably due to the very
low values (\ 0.05 mm) of the slips involved in the
debonding processes of those EBR specimens
(SS-EBRL50, L100a, L100b and L160). If another digital camera
with a better pixel resolution was used instead, then it would
be sufficient to improve such measurements at such slip
Nevertheless, both measurements are helpful for the
definition of the different states that the interface undergoes until
its failure. For instance, in the EBR specimens with the
longest bonded lengths, three different states can be
identified. In the order of their appearance in the load–slip curve
they are easily characterized by: (i) an initial linear branch
which corresponds to an elastic stage of the interface; (ii) a
nonlinear branch which may correspond to a combined
elastic and softening stage of the interface; and (iii) a plateau
that reveals the debonding initiation of the interface, i.e. the
full detachment of the SS strip from the concrete. However,
as the bond length decreases, these states tend to reduce and
for too short bonded lengths it becomes probable to identify
only the elastic stage from their load–slip responses.
Much like the EBR specimens with SS strips, the load–slip
response of the NSM specimens also allows three different
states to be identified. Figure 9 shows the load–slip
responses obtained from the NSM samples. In Fig. 9a, the
load–slip responses of the shortest specimens are shown,
whereas Fig. 9b shows the responses of the specimens where
the rupture of the SS rod occurred. Distinct from what
happened in the EBR samples with the longest bonded
lengths, where the plateau corresponds to the debonding
propagation along the bond length, the plateau observed
from specimens SS-NSM-L200 and SS-NSM-L300
corresponds to the yielding of the SS rod instead. So, the isolated
and exclusive use of the DIC technique for the
measurements of the displacements of the interface proved that by
itself it is able to identify the distinct phases of the load–slip
responses and the yielding stage of the NSM samples which
helps provide better understanding in the analysis and
interpretation of the interfacial behaviour between the SS
and the concrete.
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7
Slip at x = 0, s0 (mm)
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7
Slip at x = 0, s0 (mm)
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7
Slip at x = 0, s0 (mm)
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7
Slip at x = 0, s0 (mm)
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7
Slip at x = 0, s0 (mm)
00.00 0.05 0.10 0.15 0.20
0.6 0.8 1.0 1.2 1.4
Slip at x = 0, s0 (mm)
4.3 Slips Developed Within the Interface
The relative displacements between bonded materials (or
slips) developed along the bond length of the interface are
analysed next, in accordance to Eq. (2). Whether for the sake
of simplicity of the analysis or to avoid increasing the text
unnecessarily, only two specimens were selected to be
presented, since the bond length of an interface has an important
effect on its load–slip response: (i) the specimen with the
largest bond length; and (ii) the specimen with the shortest
bond length. Also, both strengthening bond techniques are
contemplated in this analysis. Hence, Fig. 10 shows the slip
distributions obtained from the specimens SS-EBR-L50,
SSEBR-L800, SS-NSM-L35 and SS-NSM-L300. Furthermore,
the slips developed within the interface were calculated
taking into consideration that:
s ¼ uss
where uss and uc are the displacements in the stainless steel
and in the concrete. Thus, since the DIC Correlate software
provides only the displacements, the slips measured using
the DIC technique were calculated from the differences
between the displacements measured along a line that
embraces the bond length and another one that considers and
measures the displacements along the concrete surface at the
vicinity of the interface.
The slips developed within the EBR interface chosen to be
presented in Fig. 10a correspond to the debonding loads of
the specimens and in the case of the EBR sample with the
longest bonded length, an intermediate slip at x = 0 was
randomly selected. Therefore, the distribution corresponding
to ‘‘1’’ in Fig. 10a corresponds to the same slip at the SS
loaded end when the debonding of the specimen
SS-EBRL50 occurred. In these cases, the results obtained either from
the strain gauges or from the DIC technique are presented.
Despite the noisy signal obtained from the DIC technique,
the comparison between the two monitoring methods at least
allows us to check the capability of the DIC to follow the
same trend obtained from the strain gauges. Thereby, the
results shown in Fig. 10a indicate that the DIC technique is
capable of following the same slip distributions of those
obtained from Eq. (2), i.e. with highest slips at the SS loaded
end with a decrease of the slips towards the SS free end.
The same criterion was followed in Fig. 10b to show the
slip distributions obtained from the NSM selected samples.
However, the middle slip distribution in the specimen
SSNSM-L300 corresponds to the initiation of the yielding of
the SS rod (see Fig. 10b). As can be seen from these results,
a relevant discontinuity of the slip distribution can be
observed at the vicinities of the SS loaded end, which is
explained by the yielding of the SS rod. At the same time,
the slip distributions corresponding to numbers ‘‘2’’ and ‘‘3’’
are quite similar, which can be explained, once again, by the
yielding of the SS rod outside of the bonded length. Thus,
when the SS rod yields, the load transmitted to the SS rod
remains the same and the slips along the bond length should
remain almost unchanged from then. Consequently, the
displacements increase elsewhere outside the SS-to-concrete
interface and the failure will also be localized there.
4.4 Axial Stresses and Strains Developed in the Stainless Steel
As mentioned above, the distribution of strains in the
stainless steel used on the EBR samples was obtained from
the strain gauges. In addition to the measurements collected
from the strain gauges, the strains in the stainless steel were
also measured with the DIC technique. However, as shown
in the previous subsection, the slips obtained from the DIC
technique are not quite smooth enough to obtain a smooth
strain distribution. Consequently, the axial stress
distributions obtained from both measurements are not much alike,
as shown in Fig. 11a. Besides that, from Fig. 11 it can also
be seen that the maximum axial stress in the stainless steel
used on the EBR with SS specimens and measured by the
strain gauges is quite far away from its rupture value.
Therefore, the mechanical properties of the stainless steel
were not fully used, which shows how inefficient the EBR
technique is. However, using the DIC technique, the
maximum axial stress was, as expected due to its long bonded
length, registered in specimen SS-EBR-L800, which reached
271.0 MPa at 190 mm away from the SS loaded end.
In the NSM samples, the axial stress distributions are at
least consistent with what would be considered acceptable,
i.e. the axial stresses developed in specimens SS-NSM-L35
are almost uniform and didn’t reached the yielding value of
the SS rod. The debonding failure process of short bonded
lengths is characterized by a slip distribution where there are
no undeformed bonded regions (see Fig. 10). In addition,
from the first derivative of Eq. (5) with respect to x and
ignoring the strains developed in the concrete, it could be
concluded that there are no regions where the strains could
be zero unless, of course, precisely at the SS free end, where
there are no assigned external loads to the stainless steel.
In sum, and despite the differences and the procedures that
are needed to overcome the difficulties raised by the use of
the DIC technique, the overall view of the results obtained
with this monitoring technique is positive. In particular, the
fact that several aspects observed in the experiments were
identified and validated by the DIC technique. Even when
the yielding of the SS rod was observed in specimen
SSNSM-L300, the DIC technique was capable of predicting
this as shown in the graph at the bottom of Fig. 11b.
5. Data Interpretation
In this section, the experimental results are discussed and
analyzed. Based on the experiments, the effective bond
length, i.e. the length beyond which the debonding load
cannot increase any more, is defined as the debonding load
vs. bond length graph. Moreover, the interfacial bond–slip
relationships obtained from the experiments both from the
EBR system or NSM system are herein presented.
5.1 Definition of the Effective Bond Length
The notion of effective bond length (Leff) of an interface is
widely assumed as the length beyond which the debonding
(or maximum) load cannot increase with the increase of the
bond length. In the literature, amongst other proposed
(e.g. Biscaia et al. 2013; Teng et al. 2001)
proposal made by Neubauer and Rosta´sy
can be used to estimate the debonding load
and the effective bond length of CFRP-to-concrete
interfaces. This has special importance in those cases where the
bonded length is short and the effective bond length of the
interface is not ensured. Therefore, in order to predict the
effective bond length of the SS-to-concrete interfaces, the
model proposed by Neubauer and Rosta´sy
was used here. The results are shown in
Fig. 12, where the continuous black lines represent the
curves obtained by fitting Neubauer and Rosta´sy’s model
(Neubauer and Rosta´sy 1997)
with the experimental data by
a minimization process of the maximum loads for the tested
samples with different bonded lengths. In the particular case
of the NSM system, the rupture observed from the two
specimens with the longest bond length was due to the
rupture of the SS rod and therefore, the limit of the curve
obtained with the Neubauer and Rosta´sy’s model
and Rosta´sy 1997)
corresponds to the failure load of the
stainless steel rods (48.5 kN), whilst the debonding load of
the EBR system stayed only at 17.7 kN, which represents a
reduction of 63.5% when compared to the failure load of the
30 40 50
Bond length, Lb (mm)
Short bond length, SS-EBR-L50
Short bond length, SS-NSM-L35
rods. Thus, in the case of the EBR system, the effective bond
length was found to be 235 mm, whereas the NSM system
showed an effective bond length of 168 mm. Based on these
results and regarding the maximum loads reached in both
cases, the NSM system with stainless rods showed itself to
be the best bonding system in this case too because a
classical rupture in the stainless steel is reached for a lower
effective bond length than that estimated by the EBR system
with SS strips.
Despite being beyond the purpose of this work, the
performance of the EBR systems can be improved by using an
additional anchorage system. This issue is of great
importance, as can be attested to by the numerous researches found
in the literature (Biscaia et al. 2014;
20 25 30 35
Bond length, Lb (mm)
et al. 2012;
Bre n˜a and McGuirk 2013
; Realfonzo et al. 2013;
Wu and Liu 2013
) seeking for a valid and alternative
solution for enhancing the strength of EBR systems. Still, it is
important to bear in mind that the use of steel mechanical
fasteners involves making a hole on the SS strip which may
lead to an important setback due to the reduction of the SS
cross section of the strip to install the fasteners.
Consequently, the rupture load of the stainless strip is reduced.
Therefore, an alternative method to anchor the stainless steel
is recommended instead and in this way, the innovative
CREatE technique developed by the authors
(Chastre et al.
2016; Biscaia et al. 2016c, 2017)
using CFRP laminates
instead of SS strips could be an important contribution to
this topic because once the FRP or the SS free ends are
embedded in concrete, the rupture of the FRP composite or
the yielding of the SS are always reached.
5.2 Interfacial Bond–Slip Relationship of the EBR System
To determine the local bond–slip relationship between the
stainless steel and concrete, the bond stress developed within
the interface is obtained from the equilibrium of an
infinitesimal length dx of the SS-to-concrete interface, which
gives the following equation:
ð Þ ¼ tss
where tss is the thickness of the stainless steel and drss/dx is
the variation of the axial stress in the stainless steel in the
infinitesimal length dx. The determination of the bond stress
between two consecutive strain gauges was accomplished
through the axial stressed developed in the stainless steel in
which Eq. (1) was used. So, Eq. (6) can be rewritten as a
function of the difference between two consecutive
calculated axial stresses:
s xiþ1=2 ¼ tss
where (rss,i?1 - rss,i) and (xi?1 - xi) are, respectively, the
stress in the stainless steel and the distance between two
consecutive points. It is worth keeping in mind that Eq. (7)
assumes, therefore, that the bond stresses developed between
two consecutive points are constant and, in order to ensure a
precise calculation of the bond stress, it is important to avoid
high distances (xi?1 - xi) and distances shorter than 50 mm
The slip distribution is determined by Eq. (2) and the
bond–slip relationship is then determined by coupling the
bond stress calculated from Eq. (6) and the average slip
s xiþ1=2 ¼
sðxiþ1Þ þ sðxiÞ
where s(xi?1) is the slip at point xi?1 and s(xi) is the slip at
Figure 13 shows the interfacial bond–slip relationships
obtained from EBR with SS specimens with a bond length
greater than the effective bond length, i.e. over 235 mm. In
each specimen, the different curves are presented and each
one corresponds to the mid-point of an interval between two
consecutive strain gauges at a fixed distance from the SS
loaded end. Despite some visible differences between
samples, the results show that the bond–slip relationships have
some points in common. For instance, for every single
specimen, an initial increase of the bond stress is observed
until a maximum bond stress is reached. This first stage is
usually designated in the literature
(e.g. Martinelli et al.
2011; Dehghani et al. 2012; Biscaia et al. 2013, 2014; Xia
and Teng 2005)
as an elastic stage. After this elastic stage, a
softening and nonlinear stage develops. The shape of this
softening stage is not always the same and, for instance in
specimens SS-EBR-L240, SS-EBR-400 and SS-EBR-L800
the softening stage seems to be quite symmetric with the
elastic stage, whereas in the other samples the softening
stage decays quickly after the maximum bond stress value
and tends to be almost parallel to the x-axis as the slip within
the interface approaches its ultimate value. The debonding
stage, i.e. the stage with zero bond stress transfer, is not
clearly observed in any of the EBR specimens studied here.
Also the maximum bond stress and its corresponding
maximum slip developed within the interface was not always the
5.3 Interfacial Bond–Slip Relationship of the NSM System
Unlike the EBR specimens with stainless steel which were
monitored with several strain gauges, the NSM specimens
were monitored only with the DIC technique. Hence, the
relative displacement within the SS-to-concrete interface
measured with the DIC technique was the only information
gathered from the single-lap shear tests carried out in this
work. For this reason, the methodology followed in the
previous section for determining the interfacial bond–slip
relationship of the EBR samples cannot be used in this case
of NSM specimens. The methodology followed in this
situation is described next and it was based on the strain vs.
slip curve obtained from the test at the SS loaded end (@
x = 0). This methodology was proposed by Dai et al.
et al. 2005)
and has been used since then by the authors with
(Biscaia et al. 2016c, 2017)
to evaluate the
bond–slip relationship developed within an FRP-to-concrete
interface. To determine the interfacial bond–slip relationship,
this method requires that the slip and bond stress
measurements at one specific point, which, in the present case,
corresponds to x = 0. Hence, for the current NSM samples,
the prediction of the interfacial bond–slip relationship of the
SS-to-concrete interface begins by considering, once more,
the equilibrium of a segment dx, leading to the following
(Biscaia et al. 2015a, b)
where /ss is the diameter of the stainless steel rod. Eq (9)
can be rewritten as:
sðxÞ ¼ 4 dx
/ss drss ds
sðxÞ ¼ 4 ds dx
where ds/dx is defined as the strain in the stainless steel,
since the strains developed within the concrete are ignored
due to their negligible values when compared to the strains
developed in the stainless steel. Therefore, Eq. (10) can be
rewritten as a function of the axial stresses and strains in the
SS according to:
where ess is the strain in the stainless steel. Eq (11) is then
numerically solved according to:
/ss drss ess
sðsÞ ¼ 4 ds
siðsÞ ¼ 4
rss;i 1 ess;i
where (rss,i?1 - rss,i-1) and (s0,i?1 - s0,i-1) are,
respectively, the stress in the stainless steel and the slip between
points i ? 1 and i - 1 of the stress versus slip curve
obtained from the experiments at the SS loaded end. It is
worth noting that for i = 0 the bond stress and the slip are,
respectively, s0 = 0 MPa and s0 = 0 mm. Thus, from the
stress versus strain behaviour of the stainless steel rods
already defined in Sect. 2, the interfacial bond–slip
relationship is obtained. As already shown from the EBR
samples, the DIC technique is capable of reasonably reproducing
the slips within the interface at the most loaded regions.
Taking into account the stresses in the SS rod obtained from
the load pressure cell, the reproduction of the stress versus
slip curve at the SS loaded end is defined and, based on the
stress versus strain behaviour of the SS, the interfacial bond–
slip relationship is predicted by Eq. (12). Still, when it
comes to the differentiating a noisy signal such as that
obtained from the strain distribution in the stainless steel (see
Sects. 4.3 and 4.4), the results will have an even nosier
signal and that is why an exponential smoothing with a
smooth constant of a = 0.2 was used in this case, in order to
smooth out the ‘‘peaks and valleys’’ found in the results
obtained from Eq. (12) and to show their trend is close to the
real curve. Hence, Fig. 14 shows the smoothed results for
the interfacial bond–slip relationships obtained from all the
NSM specimens. Based on these results, it can be seen that
despite the visible differences, the bond–slip relationship
determined in the NSM samples has an initial elastic stage,
as well. However, at the end of this elastic stage, the bond–
slip relationships in the NSM samples seems to show a
plateau at a peak bond stress value and then, the bond stress
decays until it reaches a zero value. Thereof, bond–slip
relationships such as elastic with fragile rupture,
rigid-plastic, rigid with linear softening or other reported in
et al. 2013)
are herein excluded and a trapezoidal shape may
describe better, even if approximately, the interfacial
behaviour shown in the NSM samples. In terms of the values of
the nuclear points needed to define the bond–slip
relationship, it can be stated that the maximum bond stress
determined from the NSM samples approximately reached
20.0 MPa. Moreover, the maximum slip, i.e. the slip at
maximum bond stress, determined in the NSM samples is
approximately 0.2 mm more than the maximum slip
calculated in the EBR samples.
5.4 Interfacial Bond–Slip Relationships: EBR
System Versus NSM System
Based on the two NSM samples that failed due to the
rupture of the SS rod (specimens SS-NSM-L200 and
SSNSM-L300), it seems that the full detachment of the SS rod
from the concrete occurs at a finite ultimate slip (sult),
whereas in the EBR samples the separation between
materials takes place with a smoother transition. To help in these
comparisons, Fig. 15 shows both curves where the bond–
slip relationship determined either from the NSM samples or
from the EBR samples are represented by areas limited by
their minimum and maximum values. As can be seen from
Fig. 15, the differences between both interfacial behaviours
Given the differences found here, the use of the NSM
bonding technique by itself is not sufficient to justify those
differences. Still, conjugated with the NSM bonding
technique, the use of ribbed stainless steel rods may justify such
differences in the bond stress transfer between samples with
different bonding techniques because the ribs increase the
friction within the interface leading to the improvement of
the bond stress transfer between materials. In fact, in Model
(Fe´de´ration Internationale du Be´ton 2010)
influence of the steel rib area is recognized as one of the
aspects that significantly affects the bond–slip relationship.
Moreover, a trapezoidal shape of the bond–slip relationship
to simulate the local bond behaviour between a ribbed steel
rod and concrete is a possibility that is covered in Model
(Fe´de´ration Internationale du Be´ton 2010)
The mid-curves shown in Fig. 15 are intended to represent
the interfacial bond–slip relationships of each bonding
technique studied here. Based on the maximum and
minimum values obtained from each bonding technique, the
midrange curves in Fig. 15 are intended to cover a mid-range
value for the EBR and the NSM bonding techniques. Hence,
the mid-range bond–slip relationship for the EBR bonding
technique is defined according to
1Þ þ smax
where smax is the maximum bond stress; smax is the slip at
maximum bond stress; and n is a constant to be defined in
order to approximate the shape of the bond–slip relationship
to the experimental results. The mid-range values needed for
the definition of Eq. (13) are: smax = 9.0 MPa,
smax = 0.031 mm and n = 2.5. The mid-range bond–slip
curve for the NSM bonding techniques is defined according
>>: s0max;2 smax;1 ðsult
if 0 s smax;1
if smax;1\s smax;2
s2Þ if smax;2\s sult
if s [ sult
where smax,1 and smax,2 are, respectively, the slips at the end
of the elastic stage and at the end of the constant stage; and
sult is the ultimate slip, i.e. the slip beyond which no further
bond transfer between materials is ensured. For the
definition of Eq. (14) shown in Fig. 15, a mid-range value for the
maximum bond stress was found smax = 16.3 MPa and with
slips smax,1 = 0.060 mm, smax,2 = 0.280 mm and sult =
An experimental work was developed in order to study the
performance of stainless steel strips and rods bonded to
concrete. As well as the use of strain gauges to determine the
interfacial bond–slip relationship between the stainless steel
and the concrete, the DIC technique was also used which
allowed a bonded area to be analysed instead of a local strain
provided by the use of single strain gauges. As an overview
of the results achieved, the following conclusions can be
The use of ribbed SS rods showed that it is possible to
obtain the rupture of the rod if an appropriate bonded
length is used. In the present experimental work, it was
found that for 200 mm the rupture of the SS rod is
reached. Thereby, the premature debonding phenomenon
of the SS rod is avoided and the mechanical properties of
the SS rod are fully used;
• the EBR samples performed poorly when compared to
the NSM samples. In all the tests carried out, the
premature debonding of the SS strip was observed at a
strain somewhat lower than its rupture value. In the EBR
samples with a short bond length, i.e. with a bonded
length shorter than the effective bond length, the rupture
occurred within the SS-to-adhesive interface, which
means that the resin has poor properties for bonding
SS strips. However, when the bond length of the
SS-toconcrete interface increases, a mixed failure mode was
observed with the separation of a thin layer of concrete
from the substrate with 2–3 mm of depth and, at the
same time, with an adhesive rupture within the
• the DIC technique can be used, although carefully, to
evaluate the bond transfer between the SS and concrete.
The displacements measured with the DIC technique and
the slips calculated from these results were reasonably
well estimated. Mainly when those values were greater
than one tenth of a millimetre, the DIC proved to be
capable of predicting the results fairly well. However, the
noisy signal obtained for the slips make it difficult to
determine the strains and bond stresses due to its higher
order, i.e. due to the first and second derivatives of the
slips with respect to x (axis parallel to the bond length)
for the calculation, respectively. Still, the methodologies
followed permitted the yielding of the SS rods in the
NSM samples to be identified and allowed us to get a fair
perspective of the strain distribution in the SS strip in the
• the DIC technique also allowed the load–slip distribution
to be captured accurately. This weighs heavily in the
evaluation of the bond between two materials because,
based on the load–slip response, the interfacial behaviour
can be predicted. Thus, depending on the load–slip
response until failure, the different stages that
characterize the bond–slip relationship can be estimated. For
instance, an initial linear load–slip response means that
the interfacial bond–slip relationship has a linear and
elastic stage as well. Afterwards, the nonlinear load–slip
response observed from the samples means that the
interfacial bond–slip relationship has a softening stage.
This transition between the linear and the nonlinear
load–slip response corresponds to a maximum bond
stress value of the bond–slip relationship;
• The effective bond length of the EBR samples was
235 mm, whereas the NSM samples had an effective
bond length of 168 mm, which represents 71.5% of the
value obtained for the EBR samples;
• the bond–slip relationships obtained for the two types of
samples studied here are different. In the EBR samples, a
power function was able to describe a mid positioning of
the experimental bond stresses (i.e. the corresponding
mid-range values between the maximum and the
minimum experimental bond stresses) obtained along the
slips within the interface at the SS loaded end. However,
in the NSM samples, a trapezoidal shape to describe the
bond–slip relationship was proposed to approximate the
experimental findings. Comparing the limit points of
both bond–slip relationships, it can be concluded that the
mid-range maximum bond stress found for the NSM
samples reached 1.8 times of that found for the EBR
samples. In term of slips, the NSM samples had higher
values with the mid-range value of the ultimate slip
developed within the interface of the NSM samples
being approximately 0.5 mm, whilst an ultimate slip of
0.4 mm was never exceeded in the EBR samples.
The first author of this work would like to express his
deepest gratitude to Fundac¸a˜o para a Cieˆncia e Tecnologia
for the partial financing of this work under the UNIDEMI
Strategic Project PEst-OE/EME/UI0667/2014 and for the
post-doctoral grant SFRH/BPD/111787/2015. The second
author is also grateful to UNIDEMI for his scientific
research grant under the Strategic Project UID/EMS/
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provided you give appropriate credit to the original
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