Bond Strength of GFRP Bars Embedded in Engineered Cementitious Composite using RILEM Beam Testing
International Journal of Concrete Structures and Materials
Bond Strength of GFRP Bars Embedded in Engineered Cementitious Composite using RILEM Beam Testing
Khandaker Muhammed Anwar Hossain 0
0 Department of Civil Engineering, Ryerson University , 350 Victoria Street, Toronto, ON , Canada
This paper presents a study on the bond characteristics of glass fiber reinforced polymer (GFRP) bars in engineered cementitious composite (ECC). Ninety beam specimens having variable parameters namely bar diameter, GFRP bar types (standard low modulus 'LM' and high modulus 'HM'), two concrete types (ECC and normal concrete 'NC') and embedded length (5, 7 and 10 times bar diameter) were tested as per RILEM specifications. Bond stress-slip characteristics and failure modes of specimens as well as influence of variable parameters on bond strength are described. The performance of various Codes and other existing equations in predicting bond strength of both low/high modulus GFRP bars embedded in ECC compared to NC is described based on experimental results. The bond strength decreased with the increase of embedment length-maximum bond strength reduction of 36% was observed. For both ECC and NC, bond strength reduced with the increase of bar size and ECC produced maximum 1.6 times higher bond strength compared to NC. Code based and other existing equations were found conservative in predicting bond strength of GFRP bars embedded in ECC.
glass fiber reinforced polymer bar; engineered cementitious composite; RILEM beam test; bond strength; codes
During the last decades, tremendous progress has been
made on the high performance concrete (HPC). Such HPC
technology involves the family of highly durable fibre
reinforced engineered composites (ECCs). ECCs have high
ductility, tight crack width and can be tailored for low to
high strength. Micro-mechanical design allows optimization
of ECC for high performance. ECC strain hardens after first
cracking, like a ductile metal, and demonstrates a strain
capacity 300 to 500 times greater than conventional concrete
(Li and Kanda 1998; Li 2003; Sahmaran et al. 2009; Ozbay
et al. 2012)
. Even at large deformation, crack widths of ECC
remain less than 60 lm. With intrinsically tight crack width
associated with self-healing potential and high tensile
ductility, ECC is the material of future which offers significant
potential to resolve durability problems of reinforced
concrete (RC) structures. Given the worldwide demand for
infrastructure systems, the potential application of ECC
either in new construction or as repair/retrofitting material is
enormous. Research over the years has contributed to the
development of green cost-effective ECC mixtures
incorporating locally available sand (instead of relatively
expensive and difficult to obtain micro-silica sand), natural
pozzolans, industrial wastes and self-healing agents
(Sahmaran et al. 2009; Ozbay et al. 2012; Sherir et al. 2015;
Hossain and Anwar 2014; Sherir et al. 2017)
as well as their
potential applications in bridge and building structures
confirming superior structural and durability performance
(Hossain et al. 2015a, b; Rafiei et al. 2013; Issani and
Corrosion is a serious problem in RC structures which
damage the steel bar-concrete interface thus degrading bond
strength and ultimately shortening the service life. Periodic
maintenance, repairs and rehabilitations of corroded RC
structures lead to substantial economic burden to authorities
in the United States, Canada and other countries in the
world. There has been an increasing demand for the alternate
materials and techniques for reinforcement in RC structures
(Hossain et al. 2014; Lee et al. 2008)
. Glass fibre reinforced
polymer (GFRP) bars are recognized as a superior
alternative to ordinary steel bars for their high strength, lightweight,
noncorrosive and nonconductive characteristics. The use of
GFRP bars can prevent deterioration due to corrosion,
improve durability and increase the service life of structures.
Researches have been conducted to evaluate the durability
performance in aggressive environment and to develop
models for predicting long-term strength retention of GFRP
bars embedded in concrete
(Benmokrane et al. 2002; Chen
et al. 2006; Robert et al. 2009; Mufti et al. 2007)
. GFRP bars
were proved to be a durable face to the concrete environment
and it was shown that the service lifetime allowed to reach
tensile strength retention of less than 50% should be infinite
(Robert et al. 2009; CSA S6-06 2006)
which also supported
the findings of
Mufti et al. (2007)
who tested few GFRP
reinforced structures after up to 9 years in service
However, it is also reported that the GFRP bars do
deteriorate over time when embedded in concrete or when
immersed in concrete pore solutions due to chemical
(Kim et al. 2012)
Gardoni et al. (2013)
time-variant probabilistic model to predict the tensile
capacity of GFRP bars embedded in concrete. The model
was based on a general diffusion model, in which water or
ions penetrate the GFRP bar matrix and degraded the glass
fiber-resin interface. The model indicated that GFRP bars
with larger diameters exhibited lower rates of capacity loss.
Park et al. (2014)
investigated the long-term flexural
behavior and ductility of GFRP and steel bar reinforced
concrete (RC) members subjected to sustained loads and
accelerated aging conditions (for example, 47 C and 80%
relative humidity). Results indicated that the accelerated
aging conditions reduced flexural capacity in not only
RCsteel, but also RC-GFRP specimens. Different types of
GFRP reinforcement exhibited different rates of degradation
of the flexural capacity when embedded in concrete under
the same exposure conditions.
Extensive research studies have been conducted on the use
of GFRP bars in structural elements and on bond
characteristics of GFRP bars with conventional concrete
et al. 2014; Lee et al. 2008; Choi et al. 2012; Masmoudi
et al. 2011; Yan et al. 2016; Gu et al. 2016; Chaallal and
Benmokrane 1993; Robert and Benmokrane 2010;
Tighiouart et al. 1998; Hao et al. 2009)
. The bond behaviour of
GFRP bars is different from that of deformed steel bars.
Bond strength of GFRP bars in conventional concrete
depends on bar diameter, surface condition (sand coated,
ribbed, helically wrapped or braided), embedment length,
bar mechanical properties and environmental conditions.
Current Canadian Code CSA S806-12 (2012) and Canadian
Highway Bridge Design Code (CHBDC) CSA S6-06 (2006)
provide design equations for the development length of
GFRP bars in conventional concrete taking into account of
bar location, bar surface, clear cover, and distance between
Okelo and Yuan (2005)
suggested an equation relating
bond capacity and concrete compressive strength based on
151 test specimens comprising concretes with compressive
strength varying from 29 to 60 MPa and GFRP bars of
6–19 mm diameter.
Lee et al. (2008)
also suggested an
equation for bond strength of sand-coated and helically
wrapped GFRP bars embedded in concrete whose
compressive strength ranged between 25 and 92 MPa.
The GFRP reinforced ECC is a new technology and can
improve strength, ductility, durability, resistance to
deterioration against aggressive environment, self-healing
capability and overall service life of structures compared to their
conventional concrete counterparts
(Sherir et al. 2015, 2017;
Hossain et al. 2015a, b)
. Full understanding of the bond
characteristics of GFRP bars in ECC is important for this
technology to be adopted in structural applications.
Currently, the existing standards, including the Canadian Codes
CSA S806-12 (2012) and CSA S6-06 (2006), do not provide
any specific design models for ECC. However, the
performance of Code based equations developed for
conventional concrete as well as those developed by other
researchers should be studied in order to validate their
applicability to GFRP reinforced ECC.
Limited research has been conducted on the bond
performance of embedded GFRP bars in ECC
(Hossain et al.
. All such research studies used traditional pullout
tests to determine GFRP bond characteristics. No research
has been conducted to study the bond strength of GFRP bars
in transversely reinforced confined ECC using beam tests as
which simulates the reinforcement
behavior in real structural elements used in building systems.
It should be noted that the beam test normally achieves
higher bond strength compared to traditional pullout test due
to auxiliary transverse reinforcement providing confinement
to the GFRP bars over the bonded length
. The confining effects of fibers and transverse
reinforcement are included in the bond strength equations
provided the Canadian Codes, which is ignored in ACI
This paper presents the results of a research conducted on
the bond characteristics of both low modulus (LM) and high
modulus (HM) GFRP bars embedded in ECC using
transversely reinforced beam specimens of two configurations
having variable parameters such as: bar diameter, bar types,
embedded length and concrete types (based on strength class).
The influences of each of these parameters on bond strength
are described. Bond strengths derived from existing Code
based design equations are compared with those obtained
from experiments. This research contributes to GFRP-ECC
bond technology where knowledge is limited and provides
data for the Code writers and professionals. The findings and
conclusions of this research will surely benefit engineers,
builders and local authorities involved in designing and
constructing structures with GFRP bar reinforced ECC.
2. Experimental Investigation
Bond tests were conducted using ninety reinforced beam
specimens in accordance with the
specifications having variable parameters as shown in Table 1. The
variable parameters for the tests were: three nominal
diameter of sand coated GFRP bars (12.5, 15.9 and 19.1 mm),
two GFRP bar types (low/standard modulus ‘LM’ and high
modulus ‘HM’), one concrete cover thickness
(approximately 40 mm), two types of concrete: traditional normal
concrete (NC) and ECC and three embedment lengths (5, 7
and 10 times bar diameter, ‘D’). A total of 90 RILEM beam
specimens were cast with ECC and NC and cured for
specific duration before they were tested. Three identical
specimens were used for each parameter. The details of the
test parameters are given in Table 1 with the specimen
parameters. In typical beam specimen designation (for
example: ECC-15.9-5D-LM or NC-15.9-5D-HM as shown
in Table 1) 15.9, 5D and LM/HM represent bar diameter,
embedment length (5 9 nominal bar diameter) and bar type,
Typical beam specimen designation: ECC-15.9-5D-LM or NC-15.9-5D-HM.
LM low modulus, HM high modulus, ECC engineered cementitious composite, NC normal concrete, D bar diameter.
* 12.7 mm GFRP bars were not used for NC.
2.1 Materials and Properties
Two concrete mixtures had been used—a Ryerson
produced green Engineered Cementitious Composite (ECC) and
a commercial normal concrete (NC). ECC was made of
polyvinyl alcohol (PVA) fibers (8 mm length, 39 lm
diameter, 1600 MPa tensile strength and 1300 kg/m3
density), local mortar sand (instead of silica sand), Portland
cement, fly ash (as 55% replacement of cement), admixtures
and water to binder ratio of 0.27. ECC mixture contained
2.6% PVA fibre content by volume. Commercial NC was
made of Portland cement, silica fume, air-entraining
admixture, 10 mm maximum size stone and other carefully
The compressive and flexural strength of the concrete was
determined from the average of ten 100 9 200 mm cylinder
and 300 9 50 9 75 mm beam control specimens that were
cast and cured under the same laboratory conditions as the
beam specimens and tested at the time of pullout specimen
testing. The strength properties of ECC and commercial NC
(determined from control specimens at the age of testing (at
28 days) as per ASTM C39/C39 M (2011) and ASTM C78/
C78 M (2010) are presented in Table 1. The cylinder
compressive strengths of ECC and NC (mean values of 10
control specimens) were 57 and 63 MPa, respectively
whereas beam flexural strengths of ECC and NC were 6.7
and 4.7 MPa, respectively.
The nominal bar diameter, tensile strength, tensile strain
and modulus of elasticity of the GFRP bars are shown in
Table 2. The actual bar diameter (db) was more than the
nominal bar diameter due to sand coating. The sand coating
was included in calculating db since it affects the concrete
surface area in contact with the bar. For each bars, the actual
bar diameter was chosen from the average of 10 diameter
readings measured with a micrometer accurate to 0.01 mm.
2.2 Beam Specimen Geometry, Configuration and Casting
The beam geometry was based on the beam test
recommendation established by
. The beam test was
comprised of two parallelepiped reinforced concrete blocks,
interconnected at the bottom by the rebar whose bond was to
be investigated and at the top by a specially fabricated steel
hinge. The dimensions of the bond test beams were
dependant on the diameter of the rebar being investigated. In the
recommendation, two beam types are given which are
dependent on the diameter of the rebar: Type A and Type B.
For the smaller 12.7, 15.9 mm diameter GFRP rebars, Type
A beams were used, whereas Type B beam specimens were
used for the larger 19.1 mm diameter GFRP bars. Table 3
and Fig. 1 show the dimensions of the specimens for Type A
and Type B. The unbonded length was created by placing
foam pipe insulation around the GFRP rebar to prevent the
concrete from bonding to the bar. As shown in Fig. 1, the
bonded portion of the GFRP rebar was located at the center
of each beam block.
The auxiliary confining reinforcement for the beam
specimens consisted of plain, mild steel bars. Details of the
reinforcement are given in Figs. 2 and 3 for Type A and
Type B beam specimens, respectively. The longitudinal steel
reinforcement was 8 mm in diameter and the transverse
reinforcement was 6 mm in diameter. The stirrups at the end
of the reinforcement cages were spot welded at the top and
bottom to the longitudinal reinforcement such that the cage
would not become distorted during casting. Cable ties were
used to secure the transverse reinforcement to the
a Calculated based on nominal bar diameter; Actual bar diameter (db) with sand coating.
Bar size, D (mm)
Nominal bar, dn (mm)
Embedment lengths, 5D/7D/10D (mm)
Thickness of concrete blocks (mm)
Depth of concrete blocks (mm)
Length of concrete blocks (mm)
Distance between concrete blocks (mm)
Total length of beams (mm)
Length of bars tested (mm)
Distance between centre-line of bar and centerline of hinge (mm)
Distance between centre-line of bar and underside of beam (mm)
Spacing of the loads (mm)
Spacing of the bearing supports (mm) Type A 12
longitudinal reinforcement. Figure 4 shows a reinforcement
cage for a Type B specimen.
2.3 Casting of Beam and Control Specimens
The beam specimens were cast using 400 L capacity
concrete mixer in the Structures laboratory of Ryerson
University. For ECC, initially the sand, cement, and fly ash
were added in the mixture machine and mixed for 2 min.
Water and HRWRA were then added into the dry mixer and
mixed for 2 min. Slight adjustments in the amount of the
HRWRA in each mixture were made to achieve consistent
and uniform matrix for better fiber distribution and
workability. Finally, PVA fiber was added and mixed for
3 min. ECC had excellent fresh and workability properties
with slump flow, V-funnel time and L-box index of 900 mm,
2.1 s and 0.98, respectively. Immediately after mixing,
flowable ECC was poured into the moulds of beam and
control specimens without consolidation (Fig. 4).
The commercial pre-packaged ready mix NC was mixed
as per the guidelines provided by the Manufacturer. The
slump for the mix was conducted according to ASTM C143/
C143 M (2015) was 120–180 mm. Air content values
performed according to ASTM C231/C231 M (2011) ranged
from 3 to 5%. Immediately after mixing, the NC was placed
into the beam and control specimen moulds and consolidated
using a vibrating table according to ASTM C192/C192 M
(2011). Figure 4 shows casting of specimens with NC.
Control specimens in the form cylinders and beams were
cast from each batch to determine compressive and flexural
strengths. Beam and control specimens were covered with
wet burlap for the first 24 h in order to maintain high
humidity on the exposed surface of the specimens. 24 h after
casting, the specimens were demoulded and left in the
laboratory (at 21 ± 2 C and 70 ± 5% RH) without
providing any special curing until tested. Figure 5 shows casted
beam and control specimens.
2.4 Beam Test-Setup, Testing and Failure
The beam test set-up (shown in Fig. 6) and testing were
designed according to the recommendation by
. The loading was applied under load control
corresponding to stresses in the bar in increments of 80, 160, 240,
320, 400 MPa, etc. The load was increased from one
increment to the next increment evenly over the span of 30 s.
Once the next increment was reached, the load was kept
constant for 2 min. After 2 min, the load was increased to
the next increment over a span of 30 s. This process was
repeated until bond failure of the specimen.
A data acquisition system was used to record data from the
load cell. Two linear variable displacement transducers
(LVDTs) located at each end of the beam specimen provided
free end slip measurements and strain gauge installed at the
centre of the GFRP bar (between the two concrete blocks)
provided the strain development during the entire loading
history. All beam tests exhibited sudden pullout mode of
3. Results and Discussion
3.1 Failure Modes
All ECC specimens (with both LM and HM bars) showed
pullout mode of failure. All NC specimens also showed
pullout mode of failure with some signs splitting cracks
around the bar only in NC-19.1-10D-LM and
NC-19.1-10DHM specimens at the beam end. From the visual inspection
of the interface between the GFRP bar and the concrete, it
could be seen that the sand coating had been partially
sheared off of the GFRP rebar.
3.2 Bond Strength, Bond-Slip Relationship and GFRP Bar Strain from Beam Tests
Based on the geometry of the beam specimens, as well as
the locations of the applied loads (F/2), supports/support
reactions and hinge as shown in Fig. 1, the tensile load (PA
or PB for Type A and B specimens, respectively) in the
GFRP rebar was derived as F(B - H)/4G by making
summation of moment about hinge equal to zero. Using the
values of B, H and G for Type A and B specimens (as shown
in Fig. 1), the bar tensile loads (PA and PB) were derived as:
PA = 1.25F; for Type A specimens; PB = 1.50F; for Type
B specimens where F is the total applied load determined by
the load cell.
Following the assumption of constant distribution of bond
stress, the average bond stress and bond strength (s) over the
embedment length (le) were determined by Eq. (1):
where P is the bar load, Ppeak is the peak bar load (maximum
value of PA or PB), db is the actual GFRP bar diameter, Ab is
the cross-sectional area of the bar and fs is the peak stress in
Table 4 shows the test results including LM/HM GFRP
bar parameters, concrete strength (fc0), bar peak load, bar
Bond stress ¼ pledb ;
Ppeak Abfs dbfs
s ¼ pledb ¼ pledb ¼ 4le
peak load as % of ultimate GFRP bar load and calculated
mean bond strength from RILEM beam tests.
3.3 Peak load development and bond stressslip response
Peak load increased with the increase of embedment length
(5D to 10D) and bar diameter (D) as indicated in Table 4. The
peak load ranged between 25.5% (observed on NC) and 74.2%
(observed in ECC) of ultimate load for specimens with LM
GFRP bars compared to the range between 20.6% (observed in
NC) and 61.2% (observed in ECC) for specimens with HM
GFRP bars. The highest peak load of 74.2% was recorded in a
12.7 mm bar (ECC-12.7-10D-LM beam specimen). This
indicates that the maximum stress developed were well below
the ultimate strength of both LM and HM bars. Thus, it can be
concluded that all beam specimens failed due to bond failure
and not due to rebar rupture.
Typical bond-slip curves from the beam tests are shown in
Fig. 7. Bond stress-slip curves exhibited similar pattern for NC
and ECC specimens and the specimens exhibited pullout mode of
failure with post peak slip development. Following the peak load
(descending branch), the pullout load dropped
quickly—indicating that as slip increased, the bonding decreased sharply.
Bond-slip curves of both ECC and NC showed similar trend of
variation, however, ECC showed lower slip at peak load and
lower bond stress reduction with slip (a measure of higher
ductility) as evident from the post-peak descending branch
(Fig. 7). ECC showed certain improvement in bond strength with
some enhanced post-peak behavior (more ductility) due to high
fiber confinement. It should be noted that the use of different fiber
ratios in ECC can further improve the post-peak behaviour.
Due to the presence of transverse reinforcement as stated
in the ACI 408R-03 (2003), the concrete is confined to
prevent a splitting failure and thus develops higher bond
stress causing bar pullout failure. As a result, bond strengths
from beam test having transverse reinforcement are higher
than those of pullout tests. This is evident from Fig. 8 which
compares the bond strength of 15.9 mm LM GFRP bars
embedded in NC from current beam tests with those
obtained from pullout tests (having similar clear cover and
concrete compressive strength) reported by
Hossain et al.
. Beam bond strength, on average, was 1.36 times
higher than those obtained from pullout test (Fig. 8).
The GFRP bar embedded in ECC also showed higher
bond strength and ductility (in terms of lower rate of post
peak slip development) compared to NC (Fig. 8) similar to
that reported by
Harajli et al. (2002)
3.4 Effect of Bar Diameter, Bar Type, Embedded
Length and Concrete Type
The effects of bar diameter and embedded length on bond
strength are shown in Fig. 9. In general, the bond strengths
Mean cylinder compressive strength of ECC and NC—57 and 63 MPa, respectively.
All specimens failed due to bar pullout.
* Mean value of three identical tests—maximum range variation from the mean = ± 1.1%.
5D 7D 10D
Embedment length ( 5, 7 and 10 mes bar dimater)
in ECC specimens were higher than their NC counterparts
for both LM and HM bars. Also, bond strengths of larger
diameter bars were lower compared to their small diameter
counterparts (Table 4 and Fig. 9). In the current study, bond
strengths of 12.7 mm diameter bar were higher than
15.9 mm bar ones and bond strengths of 15.9 mm diameter
bar were higher than 19.1 mm diameter ones
(s12.7 [ s15.9 [ s19.1). Previous studies also reported that the
bond strength of GFRP bars embedded in NC decreased with
the increase of the bar diameter
(Hossain et al. 2014; Hao
et al. 2009)
Kotynia et al. (2007)
also observed a 21% bond
strength reduction with the increase of ribbed GFRP bar
diameter from 12 to 16 mm having 10D embedment length
in 35 MPa normal concrete (made of crushed sand, stone
and cement) showing rib failure in RILEM beam test. This
could be attributed to the greater amount of bleed water
trapped beneath the larger bar diameter for NC creating
voids which consequently reduces the bond strength by
reducing rebar-concrete the contact area. For the case of
ECC, the amount of bleed water present in the fresh state is
significantly less than the normal NC. Therefore, it is
expected that the effect of bleed water should be reduced for
ECC and hence, higher bond strength should be expected
(Davalos et al. 2008)
. Another possible reason for the
reduction in bond strength with the increase of the bar size is
the Poisson’s ratio effect. Studies have shown that the
diameter reduction increases with the bar size (indicating
that the Poisson’s effect has a greater influence on the larger
diameter bar) leading to a reduction in frictional and
mechanical locking stresses producing lower bond strength
(Hossain et al. 2014; Davalos et al. 2008)
. The better bond
behavior of small diameter bars than large diameter ones for
a given embedment length can be explained by fracture
mechanics—if one considers debonding as propagation of
tunnel crack around the rebar. In this case, the debonding
force is not proportional to the rebar area and not mainly
depend on the ratio of bond surface to bar area but mostly on
the size effect. Since the size effect in plain smooth bars is
mostly related to the localized debonding at bar-concrete
interface, the balance between the energy required to
increase bar debonding and that released by the concrete
embedment has to be considered. The embedment
lengthdiameter ratio plays a major role—the larger the ratio the
larger the size effect
(Stang et al. 1990; Bamonte et al.
On the other hand, for high strain capacity of ECC
(300–500 times greater than the NC) the reduction in
frictional and mechanical locking stresses due to the reduction
in bar diameter for Poisson’s effect could be minimized by
deformation in concrete. This should increase the bond
strength of ECC specimens compared to their NC
Bar dia (mm)
counterparts. Table 5 shows that the bond strength of both
LM and HM bars in ECC was consistently higher (maximum
1.60 times higher for LM and 1.50 times higher for HM bars
as evident from the ratio of bond strength) than their NC
counterparts irrespective of embedded length and bar
Bond strength of HM bars are comparatively lower than
there LM counter parts for both ECC and NC irrespective of
embedded length and bar diameter. This is evident from the
ratio of bond strength of HM bar (sHM) to LM bar (sLM)
ranging between 0.67 and 1.06 with a mean value of 0.87
(Table 4). The lower bond strength of HM bar was attributed
to the premature detachment of the sand coating from the
core compared to only small area of delamination for LM
bar. This could due to the higher strength and lower interface
bond between sand coating and FRP core for the HM bars
used in this study. This was confirmed from the higher
interface delamination of HM bars in tested beam specimens
compared to their LM counterparts.
Higher bond strength in ECC specimens could also be
attributed to the generation of high radial confinement due to
the presence of PVA fiber and transverse reinforcement. In
the beam specimens, the concrete surrounding the GFRP bar
was under tension, making the beam specimens more
susceptible to cracking. The strain hardening and
micro-cracking characteristics of ECC would provide more resistance to
such cracking compared to NC. The fiber-bridging tends to
stop the propagation of splitting cracks (initiated when the
tangential bond stress exceeds the tensile strength of ECC)
leading to comparatively ductile pullout failure of all ECC
specimens rather than splitting. The confinement provided
by the crack bridging combined with higher energy
absorption for crack propagation can be attributed to the
improved bond strength of ECC. Previous research studies
also confirmed such improved bond strength in steel fiber
reinforced concrete (
Ezeldin and Balaguru 1989
et al. 1991
; ACI 446.1R-91 (1991). For practical
construction point of view, the conclusions derived from the beam
tests are more practical (as it simulates actual concrete-rebar
interaction in structural elements) than traditional pullout
According to Table 4 and Fig. 9, the bond strength (LM
and HM bars), in general, decreased with the increase of
embedment length. This decrease was associated with the
calculation of bond strength assuming constant bond stress
distribution while higher non-linear distribution of bond
stresses normally happened with longer embedded length in
Achillides and Pilakoutas (2004)
reported that as
the embedment length increases, the stress is unevenly
distributed over a longer length leading to the decrease in
average bond strength. The reduction of bond strength (with
respect to 5D embedment bond strength) increased with the
increase of embedment length (Fig. 10) for ECC/NC or LM/
HM bar. For NC, maximum bond strength reductions of 25
and 30% were observed for LM and HM bar, respectively,
while reductions of 24 and 36% for ECC were observed
(Fig. 10). A reduction in bond strength with the increase in
embedment length was observed in fiber reinforced concrete
(made of hooked steel fiber and coarse stone aggregate) for
sand coated/ribbed GFRP bar in RILEM beam tests having
confining reinforcement in previous research studies
Mazaheripour et al. (2013)
—bond strength reductions of
about 28 and 26% were observed respectively for 12 mm
sand coated and ribbed bars showing pullout mode of failure
when embedment length was increased from 5D to 10D.
HM bars generally produced lower bond strength
compared to their LM counterparts. According to the Canadian
Bridge Design Code (CHBDC)
(CSA S6-06 2006)
modular ratio (ratio of modulus of elasticity of FRP bar to
that of steel bar, ‘Efrp/E’s) should yield higher bond strength.
This implies that HM bar should produce higher bond
strength than its LM counterpart. The lower bond strength of
HM bar in this study was attributed to the premature
detachment of the sand coating from the rebar core—for HM
bar, the entire sand coating was delaminated compared to
only small area of delamination for LM bar (Fig. 11a, b).
This finding can lead to the restricted use of HM bar.
However, this phenomenon was thought to be associated
with particular type of HM bar used in this study. It is
expected that HM bars should develop equal or higher bond
strength as LM bars and more investigations should be
conducted with the currently available HM bars in the
market. In general, the development length of steel bar
should be longer enough to guarantee the yielding of
reinforcement. This research indicates that the embedment
length of GFRP bars should not be longer than the certain
threshold as GFRP bars did not reach their failure load and
failure was governed by pullout (as shown in Fig. 11c, d)
due to delamination or detachment of sand coating.
4. Code Based Bond Analyses and Comparison
For comparative purposes, the bond strength provided by
the design codes are determined based on the specimen
configurations in terms of bar size, concrete strength,
concrete cover and reinforcement properties used in the RILEM
beam tests. The average bond strengths calculated as per
CSA S806-12 (2012), CSA S6-06 (2006) and ACI
440.1R15 (2015) are compared with those determined from beam
specimens using Eq. (1).
Bond strength as per CSA S806-12 (2012) can be
determined from Eq. (2):
s ¼ 1:15k1dkc2sk3kfc40 k5pdb ð2Þ
where k1 is bar location factor; k2 is concrete density factor;
k3 is bar size factor; k4 is bar fibre factor; k5 is bar surface
profile factor; dcs is the smaller of the distance from the
closest concrete surface to the centre of the bar being
developed and two-thirds of the centre-to-centre spacing of
the bars being developed (shall not be taken greater than 2.5
times bar diameter ‘db‘, mm); Efrp and fc0 is the concrete
Bond strength as per CSA S6-06 (2006) can be derived
from Eq. (3):
Embedment Length (mm) (x bar diameter D)
Embedment Length (mm) (x bar diameter D)
Fig. 11 a LM bar: partial delamination of sand coating, b HM bar: detachment of entire sand coating, c bar pullout from the left
hand part showing outward movement and d bar pullout showing inward movement of the bar.
dcs þ ktr EEfrsp fcr
dcs þ ktr Es
where k6 is bar surface factor; Efrp and Es are the modulus of
elasticity of FRP and steel in MPa, respectivelyp; ffiffficffir is the
flexural strength of concrete in MPa (usually 0.4 fc0); Atr is
the area of transverse reinforcement normal to the plane
through the anchored bars in mm2; fy is the yield stress of
steel in MPa; s is the spacing of transverse reinforcement in
mm; n is the number of bars being developed or spliced; ktr
is the transverse reinforcement index. The maximum
permissible value of f0c. is limited to 64 MPa by CSA S806-02
and CSA S6-06.
According to ACI 440.1R-15 (2015), bond strength can be
calculated from Eq. (4):
0:33 þ 0:025 db þ 8:3 dleb
where le is the embedment length, C is the lesser of distance
from the cover to the center of the bar, one-half of the
centeron-center spacing of the bars being developed (C/db should
not be taken larger than 3.5).
Table 6 compares the bond strengths obtained from beam
tests with those derived from Codes and other existing
equations. The concrete strength and the bar size do not
influence the bond strength as per CSA-S806-12 and CSA
S6-06. However, it can be noticed from the test results that
the higher tensile strength and toughness (due to fiber
presence and confinement) of ECC could affect the bond
strength—the prediction equations should include
compressive strength and some parameter that takes into account the
influence of enhanced tensile strength and confinement. As
the distance from the center of the bar to the closest concrete
surface was greater than two times the bar diameter, the
effect of confinement provided by the transverse
reinforcement (taken into account by the transverse reinforcement
index, ktr., as per CSA S6-06), was neglected.
Mean cylinder compressive strength of ECC and NC—57 and 63 MPa, respectively.
The ratios of experimental to predicted (Code based and
other existing equations) bond strength are summarized in
Table 7. Between 5D and 10D embedment length,
CSAS806-12 (with ratio of experimental to predicted bond
strength ‘‘r’’ ranging between 1.24 and 2.74 with a mean
value ranging between 2.08 and CSA S6-06 (with ‘r’
ranging between 1.77 and 4.06 with a mean value of 3.07
lead to conservative bond strength prediction of ECC
Okelo and Yuan (2005)
Eq. (5) also under predicted
the bond strength of both LM/HM bars in ECC (‘r’ ranging
between 1.63 and 3.58 with a mean value of 2.67—Eq. (5)
reflects the influence of bar diameter and the concrete
compressive strength of up to 60 MPa. ACI 440.1-15 (with
‘r’ ranging between 1.03 and 1.73 with a mean value of
1.37) provided a reasonably good bond strength prediction
though conservative. It should be noted that the ACI Eq. (4)
is developed based on concrete strength between 28 and
45 MPa. Equation 6 of
Lee et al. (2008)
comparatively better prediction of bond strength compared to other
equations except ACI Eq. (4) with ‘r’ ranging between 1.05
and 1.97 having mean value of 1.50.
For NC similar to ECC (Table 7), CSA-S806-12 (with ‘‘r’’
ranging between 1.43 and 2.37 with a mean value of 1.43,
CSA S6-06 (with ‘r’ ranging between 2.19 and 3.56 with a
mean value of 2.66 and Eq. (5) of
Okelo and Yuan (2005)
(‘r’ ranging between 1.69 and 2.76 with a mean value of
2.05) under predicted the bond strength. ACI 440.1-15 (with
‘r’ ranging between 0.55 and 1.34 with a mean value of
0.91) showed slight over prediction while Eq. (6) of
et al. (2008)
showed comparatively better prediction with ‘r’
ranging between 0.82 and 1.62 having mean value of 1.09.
Overall, bond strength (for both LM and HM bars) derived
from the beam tests are higher than those predicted by Code
based and other existing equations. It can be concluded that
these equations provide a conservative prediction of bond
strength. Hence the bond strength predicted by the design
Codes and other equations will be safe for both LM and HM
GFRP bars embedded in ECC confinement but needs to be
The bond behaviour of sand coated low/standard modulus
(LM) and high modulus (HM) GFRP bars embedded in
selfconsolidating engineered cementitious composite (ECC) and
normal vibrated concrete (NC) was investigated. Ninety
beam specimens were tested to study analyze the effects of
bar diameter, bar type, embedment length and concrete type
on bond strength and failure modes. Based on experimental
and Code based analyses, the following conclusions were
All specimens showed pullout mode of failure under
varying embedment length, bar type, bar diameter and
The peak load in GFRP bar increased with the increase
of the embedded length for both ECC and NC
specimens. ECC specimens developed higher peak load
compared to their NC counterparts. All beam specimens
failed due to bond failure and not due to rebar rupture as
the maximum peak load was only 74.2% of the ultimate
load of the bar.
Bond strength of both LM and HM bars in ECC was
higher (maximum 1.60 and 1.50 times higher for LM and
HM bars, respectively) than their NC counterparts
irrespective of embedded length and bar diameter.
Bond strengths of HM bars were consistently lower than
their LM counterparts for both ECC and NC irrespective
of embedded length and bar diameter—as evident from
the mean ratio of bond strength of HM bar to LM bar of
0.87. This was primarily attributed to the premature
detachment of the sand coating from the rebar core. This
phenomenon was thought to be associated with particular
type of HM bar used in this study and more
investigations are needed with HM bars currently available in the
The bond strength, in general, decreased with the
increase of embedment length from 5 to 10 time bar
diameter. For NC and ECC, maximum bond strength
reduction of 30 and 36%, respectively were observed.
Bond strengths (derived from beam tests) of both LM
and HM bars embedded in ECC were found to be higher
than those predicted by CSA S806-12, CSA S6-06, ACI
440.1R-15 and other existing equations. ACI 440.1-15
provided a better bond strength prediction compared to
other equations with ratio of predicted to experimental
values ranging from 1.03 to 1.33 for NC and from 1.18
to 1.49 for ECC. It is concluded that the Code based and
other existing equations provide a conservative
prediction of bond strength and can safely be used for the bond
strength prediction of GFRP bars embedded in ECC.
However, the mean ratio of predicted to experimental
values ranging between 1.37 and 3.07 suggests that
modifications to these Codes are necessary for the
prediction of bond strength of ECC.
Authors acknowledge the financial support provided by the
Natural Science and Engineering Research Council
(NSERC) Canada. The collaboration from V-ROD Canada
is also acknowledged.
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.
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