Fundamental Mechanisms of Bonding of Glass Fiber
Reinforced Polymer Reinforcement to Concrete
Wai How Soong1, J. Raghavan1,#, and Sami H. Rizkalla2
1
Department of Mechanical and Manufacturing Engineering
University of Manitoba, Winnipeg, MB R3T 5V6, Canada
2
Department of Civil, Construction, and Environmental Engineering
North Carolina State University, Raleigh, NC 27695, USA
#Author for Communication
Mechanical & Manufacturing Engineering Department
E2-327, EITC Building
University of Manitoba
75, Chancellor Circle
Winnipeg, MB R3T 6A6, Canada
Tel. 204-474-7430
Fax. 204-275-7507
E-mail:- rags@cc.umanitoba.ca
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ABSTRACT
Fundamental mechanisms of bonding between Glass Fiber Reinforced Polymer (GFRP) bar and
concrete are presented. Contributions from chemical bonding, bearing resistance, and frictional
resistance to bond were delineated by measuring the following: the load corresponding to
complete debonding of the bar, the load corresponding to onset of sliding and pullout of the bar
along the entire embedment length, and the frictional load corresponding to frictional resistance
to sliding. Research findings indicate that while chemical bonding was the main contributor to
the interfacial bond strength, the other two mechanisms contributed to the pullout strength of the
bar. Correlation between the bar’s surface geometry and the contributions from the three
mechanisms are discussed.
Keywords: Polymer Composite, Reinforced Concrete, Pullout, Interfacial strength, Bond
strength, Fiber Reinforced Polymer
1. Introduction
Recently, polymer composite materials are used in civil engineering applications in various
forms including rebars for concrete structures, sheets for flexural and shear strengthening, and
sheets to wrap concrete columns and bridge piers to increase the confinement. Fiber Reinforced
Polymer (FRP) bar is an excellent alternative to steel bars due to their relatively higher corrosion
resistance and high specific strength and modulus. Several successful applications can be found
in North America. While design codes are well established for the use of steel bars in concrete
members, they are slowly evolving for FRP bars. Development of these codes requires a good
understanding of bond between FRP bars and concrete and the relationship between the bond
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strength, and various material and test parameters. Past research has substantially advanced the
knowledge in this area. Nevertheless, a substantial variation is observed among the published
data on interfacial bond strength, which suggests that a comprehensive understanding is yet to
evolve.
Several variations of bar pullout and beam test have been used to study the bond
characteristics of FRP bars. Various parameters, whose effect on interfacial bond strength have
been studied in the past, are bar embedment length, bar diameter, compressive strength of the
concrete, confinement pressure exerted by the concrete on the bar, and surface geometry of the
bar.
Al-Zahrani [1] and Benmokrane et al. [2] have observed a decrease in interfacial bond
strength with increase in embedment length, as shown in Table 1 for different glass fiber
reinforced plastic (GFRP) bars with axi-symmetric lugs [1] and helical lugs [2]. It has been well
established that shear stress at the fiber – matrix interface, at any applied load during a single
fiber direct pullout test, is not constant. Al-Zharani et al. [3] and Benmokrane et al. [4] have
measured this shear stress distribution experimentally, for concrete reinforced with FRP bar.
The interfacial bond strength tabulated in Table 1 has been determined using average shear stress
and it varies with embedment length, since the shear stress distribution and the average shear
stress would change with embedment length.
In addition, reported bond strength values
correspond to bar pullout rather than onset or completion of debonding. This is also a reason for
the variation observed in Table 1 and will be discussed in subsequent sections.
Test results indicate that increasing bar diameter caused decrease in the bond strength as
given in Table 2 for smooth bars [1] and bars with lugs [5]. While variation of shear stress
distribution could be one reason, Tighiouart et al. [5] have indicated that bleeding of water could
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be another reason since an increase in bar diameter could increase the amount of water trapped
near the bar leading to higher amount of voids and lesser contact area.
Al-Zharani et al. [3] have varied the compressive strength of the concrete, reinforced with
bars wrapped with lugs, from 31.4 MPa to 66.1 MPa and have observed no change in the
interfacial bond strength. This is to be expected since the failure was interfacial. Another work of
Al-Zahrani [1] has shown that the induced lateral force could be influenced by the mismatch in
Coefficient of Thermal Expansion (CTE) between the concrete and the bar as observed in Table
3 for concrete reinforced with a smooth bar. A test temperature lower than the cure temperature
resulted in negligible bond strength probably due to lack of lateral pressure caused by the
contraction of the bar. However, a test temperature higher than the cure temperature resulted in a
bond strength higher than the reference case, for which the test and cure temperatures were same.
This is probably due to increase in lateral pressure due to expansion of the bar. These results are
similar to the results of Leung and Geng [6], who have shown that the interfacial bond for
concrete reinforced with steel bars increases with lateral pressure.
Researchers have also varied the surface geometry of the FRP bars to enhance the
resistance to sliding and pullout strength as tabulated in Table 4. Al-Zahrani [1] has observed
decrease of the bond strength by increasing the lug width from 3.8 mm to 8.9 mm. The latter also
changed the failure mode from interfacial failure to crushing failure of the concrete between the
lugs.
Test results of Benmokrane et al. [2] and Al-Zahrani [1] are compared in Table 5 to
illustrate the influence of loading rate and inconsistency in the definition of contact surface area
used in the calculation of interfacial bond strength. Since the dimensions and the properties of
the reinforcing bar and the concrete were same, the difference in the maximum pullout load is
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believed to be due to the difference in the loading rates used by the two groups of researchers. In
addition, the interfacial bond strength reported by Al-Zaharani [1] is higher than that reported by
Benmokrane et al. [2] despite the lower value for the measured pullout load and same bar
dimensions. This apparent discrepancy is believed to be due to the difference in the definition
and calculation of the contact surface area. While Benmokrane et al. [2] have calculated the
contact surface area using the average diameter of the reinforcing bar, Al-Zahrani [1] has
calculated the contact area using the actual diameter of the reinforcing bar and the dimensions of
the lugs.
The above discussion suggests that the interfacial bond strength could be very much
dependent on test parameters and surface geometry of the bar, whereas the interfacial bond
strength should be a unique value independent of test parameters and surface geometry of the
bar.
2. Fundamental behavior during pullout tests
A brief discussion on the pullout behavior of a FRP bar from concrete is included here to
assist in the comprehension of the scatter in published interfacial bond strength and to provide
the basis for the research approach used in this study.
This development length for civil engineers (lc), which is also known as critical length
among composite community, can be predicted using equation (1) based on assumption of
constant shear stress distribution along the embedded fiber length.
lc = fd / 4u
(1)
where f is the fiber strength, d is the bar diameter and u is the fiber-matrix interfacial bond
strength.
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Load applied to the concrete – FRP bar interface, during direct pullout test of a concrete
specimen reinforced with a FRP bar of embedment length (l)  lc, is shown schematically in Fig.
1 as a function of slip between the bar and the concrete. Fig. 1(a) represents the possible load –
slip curves for a concrete specimen reinforced with a smooth FRP bar. Curve A represents the
case when debonding of the smooth FRP bar occurs at a maximum load (Fd’). The applied load
may drop either to zero or to a finite frictional force value, Ff’. Pullout behavior of the FRP bar
may also follow curves B or C since the interfacial stress varies along the embedment length.
Curve C represents possible case for a stable debonding of the bar, which starts at Fi’ and
progresses to completion along the entire embedment length at Ff’. Curve B represents possible
partial debonding. In this case, the remaining bonded portion of the bar debonds in an unstable
manner at the peak load and the load drops suddenly to Ff’. Therefore, for cases B and C partial
sliding and pullout of the bar in the debonded region, and progressive debonding of the bar in the
bonded region would occur simultaneously during loading. Since frictional force supports the
sliding and pullout, the actual load - slip behavior is dependent on the frictional resistance, the
interfacial bond strength, and the length of bonded and debonded regions at any given load.
The interfacial bond strength calculated using Fd’, for case A, or Fi’, for cases B and C,
would be a measure of the chemical bonding between the FRP bar and the concrete. Owing to
weak chemical bonding between concrete and the FRP bar, researchers have modified the FRP
bar surface, using lugs and sand particles, to enhance its resistance to sliding and pullout. Fig
1(b) represents the load-slip plot for debonding and pullout of a bar with lugs or sand particles.
Due to shear stress distribution along the embedment length, during a pullout test, debonding of
the FRP bar would start at Fi and would complete at Fd, for all the three cases in Fig. 1(b). Curve
A’ represents the case when the applied load continues to increase even after the completion of
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debonding at Fd. This increase is attributed to the bearing resistance, caused by mechanical
interlocking of the surfaces of the concrete and the bar, and frictional resistance, caused by the
bar’s surface roughness. When the lugs or the sand particles along the entire embedment length
are sheared at the maximum load Fp, the bearing resistance due to them is eliminated and the
load drops suddenly to the frictional force, Ff. Alternatively, shearing of lugs and sand particles
can be progressive as observed in this study. Curve C’ represents the case when the shearing of
lugs or sand particles is complete before reaching the maximum load. Curve B’ represents the
case when the shearing is partial and hence, a load drop is registered at the maximum load due to
the sudden shearing of the remaining intact lugs or sand particles.
Between Fi and Fd, partial sliding and pullout of the bar occurs along the debonded
length. Assuming that the contribution from bearing and frictional resistances to Fd, during this
partial sliding and pullout, is negligible either Fi or Fd can be used in determining the interfacial
bond strength. However, most of the researchers have used the load at which sliding and pullout
of the bar occur along the entire embedment length to determine the interfacial bond strength.
The maximum pullout load, Fp, can be predicted using equation (2)
Fp = Fd + Fb + Ff
(2)
where Fd, Fb, and Ff are debond load, bearing load, and frictional load respectively. Fd, Fb, and Ff
represent contributions to pullout load from chemical bonding, bearing resistance, and frictional
resistance respectively.
The interfacial bond strength determined using the debond load would be a unique value
independent of the surface geometry of the bar and the test conditions. However, the interfacial
bond strength determined using the pullout load would not be a unique value because the pullout
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load is dependent on these three mechanisms, two of which are dependent on the surface
geometry of the bar and the test conditions. Unfortunately, most of the previous studies have
used the pullout load to determine the interfacial bond strength. Since the parameters that
influence the bearing and frictional loads were varied arbitrarily, the reported values for the
interfacial bond strength also vary by a wide margin, from researcher to researcher. In addition,
owing to lack of delineation of contributions from the three mechanisms, previous researchers
were unable to correlate the measured pullout load to test parameters and surface condition of the
bar.
Hence, the objectives of this research are to (a) delineate the contributions from the three
mechanisms (chemical bonding, bearing resistance, frictional resistance) to the pullout load,(b)
correlate these contributions to bar characteristics such as surface roughness, lug pitch, and
number of lugs per unit embedment length, and (c) study the effect of loading rate on pullout
load. Additionally, interpretation of the published results is also complicated due to the variation
in the definition of the contact surface area from one researcher to another and by the assumption
of constant shear stress distribution along the embedment length. The former issue is addressed
in this study through explicit definition of the contact surface area. The latter issue is deferred to
a future study though the impact of this could be observed in this study too.
Since the completion of this study, there has been a number of published experimental
research studies [7-11] focusing on the pullout behavior of FRP bars from concrete. However,
none of them have delineated the contributions to pullout load from various mechanisms, as done
in this study.
Finally, even though sliding and pullout of the bar after complete debonding is not focused
in this study, it is worth mentioning to understand the measured load-slip curve. If the bar is of
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infinite length and the interface is not altered during pullout, the bar will pullout at constant Ff’ or
Ff and curve D or D’ would be recorded respectively. If the bar is of finite length, load required
to overcome the friction will reduce with fiber pullout due to reduction in the embedment length
(l) and curve E or E’ would be recorded. If constant frictional resistance is assumed, then curves
E and E’ would have a constant slope. If frictional resistance is reduced due to abrasion between
the bar and the concrete, curves E and E’ would have non-linear slopes as shown in Fig.1(b). This
is termed as slip weakening. In many cases, a periodic increase and drop in the load is observed
during sliding and pullout as shown in curve E’. This is due to one or more of the following: (a)
resistance from the lugs in the free end of the bar while it moves against the concrete during
testing of a specimen type used in this study; all lugs in the free end of the composite reinforcing
bar were removed in this study to avoid this; (b) resistance caused by the debris of sheared lugs
or sand particles; (c) resistance due to mechanical interlocking of the surfaces of the bar and the
concrete due to surface roughness. Alternatively, if frictional resistance is increased due to debris
from interaction between the bar and the concrete during pullout, curve G or G’ would be
observed. Such a behavior is often termed as slip strengthening.
3. Experimental details
3.1 Material
Four types of GFRP bars used in this study are shown in Fig. 2. The diameters of Smooth
(S) and Sand-coated (SC) bars were 12.7 mm and 13.6 mm respectively. Their average tensile
strength and modulus, as per manufacturer’s data sheet, were 683 MPa and 40 GPa respectively.
The bars with lugs (RL), which were basically resin-impregnated fiber strands, had shape and
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dimensions as shown in Fig. 3. They had two longitudinal lugs positioned opposite to each other.
The helical lugs were wound on the bar between the longitudinal lugs such that it started at one
longitudinal lug and ended at the other. The number of lugs in the as-received RL bars was 12
per 88.9 mm, i.e. a pitch of 4.4 mm. Their average tensile strength and elastic modulus, as per
manufacturer’s data sheet, were 770 MPa, and 42 GPa respectively. These as-received RL bars
were machined in a lathe to remove the lugs. The pitch of lugs was altered by removing alternate
lugs or two consecutive lugs to result in bars with 6 (SL) and 3 (TL) lugs per 88.9 mm
respectively. The pitch of the lugs in SL and TL bars was 11.95 mm and 26.9 mm respectively.
Complete removal of lugs resulted in machined (M) bars. While S bars were used as the
reference, SC and M bars were used to study the influence surface roughness. RL, SL and TL
bars were used to study the influence of lugs and the pitch of lugs. The concrete mix of the airentrained concrete used in this study was a proper mixture of water, portland cement, sand and
gravel in the ratio 1:2.1:4.31:6, as suggested by Portland Cement Association [12].
3.2 Test specimen
The pullout specimen was a concrete cylinder with the FRP bar embedded at its center, as
shown in Fig.4. Diameter and height of these specimens were 152.4 mm and 304.8 mm
respectively. Bonding between the bar and the concrete was prevented at both the ends of the
cylinder using PVC tubes, to eliminate any end effect. The length of this tube at the load end was
maintained at 152.4 mm. The length of this tube at the free end was varied to accommodate bars
with various embedment lengths. Preliminary experiments were carried out using RL and M
specimens, with various embedment lengths given in Table 5, to determine the embedment
length that would yield interfacial failure during subsequent tests. Based on these tests, an
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embedment length of 88.9 mm was chosen for subsequent tests. A steel tube was bonded to the
load end of the bar using a room temperature curing epoxy and used as a gripping system for the
GFRP bars.
The specimens were cast using a plastic mold. The bars were held in position at the
center of the mold using a specially designed wooden fixture. The PVC tubes were bonded to the
bars using clay, which was subsequently removed after curing. Concrete was mixed in a concrete
mixture, poured in to the mold, and packed using a hand-held ram. The molds were covered with
a plastic sheet and the specimens were allowed to cure for 28 days. The specimen identification
codes are tabulated in Table 6. The compressive strength of the concrete was determined as per
ASTM C39-86 to be 49.77  2.26 MPa.
3.3 Direct pullout test
The direct pullout test was carried out using stroke-control mode using a MTS hydraulic
test frame with a maximum loading capacity of 1000 kN. The test setup is shown in Fig. 4.
During loading the concrete cylinder was pressed against a steel plate, bolted to the ground using
two 2.25” diameter bolts. Since the test specimen surface was not flat, a thin layer of plaster was
applied on the test specimen surface to ensure uniform contact with the steel plate. Two LVDT’s
were attached rigidly to the specimens using specially designed fixture as shown in Fig. 4.
Relative slip between the bar and the concrete was measured using these LVDTs at both ends of
the bar. The LVDT data was acquired using a data acquisition system and stored in a computer.
All tests were done at a displacement rate of 1.3 mm / min. Two additional loading rates of 0.26
mm / min and 6.5 mm / min were used to study the influence of loading rate on pullout load.
While most of the specimens were unloaded after recording the minimum load to which the
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applied load dropped beyond Fp, few specimens were loaded beyond this point to record the
pullout behavior. A minimum of three specimens was tested to obtain each data point.
All tested specimens were dissected to confirm the mode of failure. The specimens were
cut using a water-cooled 14” diameter diamond tipped blade to a distance of about 1” from the
bar. Subsequently all the specimens were pried open using a chisel and a hammer to reveal the
interface. In order to study the progression of debond, lug failure, and shearing of sand particles
during loading, RLC, SLC, TLC, SSC, MC, and SC specimens were loaded to pre-selected load
levels below Fp and then unloaded. Subsequently the specimens were dissected to examine the
interface.
While the nominal bar diameter, measured using a digital caliper, was used in calculating
the contact surface area for the S, M and SC specimens, equation (3) was used to calculate the
contact surface area for the RL, SL and TL type specimens.
A = Abar + (N * Ahlug) + Allug – (N* Aoverlap)
(3)
where
Abar = Total surface area without lug = d x l x 
Ahlug = Contact surface area for one helical lug = 2*(*(D2 – d2)/ (4*sin)) + (Da/sin))
N = Number of helical lug
Allug = Contact surface area for longitudinal lugs = 2 (2h + w) p
Aoverlap = Overlap area between bar and lug = da/sin
Values for D, d, a, h, w, P, are given in Fig. 3. It was assumed in this study that the area over
which debonding occured was the same area over which subsequent frictional sliding occured.
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Hence, the same contact surface area was used in the determination of both the interfacial bond
strength and the frictional stress. However, this assumption has to be examined in future studies
since progressive shearing of lugs and sand particles was observed in this study.
While the interfacial bond strength was determined using the debond load, no attempt
was made to determine the pullout strength since it would not be a unique value that can be used
in any design.
4. Results and Discussion
The experimental program undertaken in this study was designed to measure the debond
load (Fd), the pullout load (Fp), and the frictional load (Ff). Representative load – slip curves for
SC, MC, SCC, and SLC specimens are shown in Fig. 5. Elastic deformation of the bar has been
deducted from the LVDT measurements to obtain the relative slip between the bar and concrete.
Load versus slip relationship for the free end is nearly a mirror image of that for the load end.
The load at which the free-end slip started to increase was defined as Fd since the free end of the
bar could start to slip only after the debonding had progressed completely from the load end to
the free end. For all specimens, the load-end slip started to increase immediately after the start of
application of load while the free-end slip was zero until Fd. Since the free-end slip started to
increase beyond Fd, debonding was believed to be complete at Fd. All specimens loaded to a load
slightly higher than Fd were unloaded and dissected to confirm the mode of failure. Clean bar
surface without any adhering concrete particles confirmed that the failure mode was interfacial
[13]. This confirmed that debonding was complete at Fd.
The maximum measured load was defined as Fp and Ff was defined as the value to which
the applied load dropped beyond Fp. Subsequently, the bearing load was determined using
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equation (2) and experimentally measured values of Fd, Ff and Fp. Thus, the contributions from
chemical bonding, bearing resistance, and frictional resistance to the measured pullout load were
delineated in this study.
While the relative movement between the bar and the concrete is possible along the entire
embedment length at loads greater than Fd, this is possible only along the debonded length at
loads less than Fd. This relative motion would be resisted by the lugs and the sand particles
bonded to the surface of the bar and hence the observed increase in load between Fd and Fp is
due to frictional and bearing resistance. Beyond Fp, the applied load suddenly dropped to a
minimum value, Ff, in all specimens except SC. Unlike the schematic shown in Fig.1 this load
drop was accompanied with simultaneous pullout of the bar. Beyond Ff (Fp for SC), a stick-slip
type of pull out behavior was noticed for all specimens. This is believed to be due to surface
roughness and, in case of bars with lugs or sand particles, to be due to sheared lugs or sand
particles caught between the sliding bar and the concrete. It can be observed in Fig.5 that the
maximum and minimum loads for MC and SLC specimens, during the stick and slip behavior,
decreases with increase in slip. This indicates a change in surface characteristics of the bar and
the concrete, and hence, the frictional condition.
An examination of the dissected RLC and SCC specimens, loaded to Fp, revealed
complete shearing of lugs and sand particles as shown in Figures 6 and 7 respectively. Since lugs
and sand particles are the source of bearing resistance, it can be concluded that the bearing
resistance didn’t contribute to the load recorded beyond Fp. If there were no frictional resistance,
the load would drop to zero beyond Fp. Hence, the recorded Ff is a measure of contribution of
frictional resistance to measured Fp. Since SC samples didn’t have any surface features that
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would cause mechanical interlocking, bearing resistance was assumed to be zero and hence, the
increase in load between Fd and Fp was taken to be due to frictional resistance to slip.
The measured pullout load (Fp) is plotted in Fig. 8 as a sum of the debond load (Fd), the
frictional load (Ff), and the bearing load (Fb) for various types of bars. The pullout load increased
from SC to SCC and this is attributed to increase in surface roughness from SC to SCC. It also
increased from TLC to RLC and this is attributed to increase in the number of lugs from TLC to
RLC. It is interesting to note that the pullout load for SCC specimens is comparable to that for
RLC specimens.
Normally, relative sliding of two surfaces at a frictional interface would not occur until
the interfacial shear force exceeds the frictional resistance for sliding. This is referred to as static
friction and the maximum force (Fstatic) for initiation of the sliding is related to normal force (N)
acting on the interface through static friction coefficient, static (=Fstatic /N). The force (Fdynamic =
Ff) required to maintain the sliding of the interfaces may be equal to or less than (Fstatic). In case
of the latter,dynamic < static, which has been noted by numerous published literature. Whether
this would be applicable for a frictional interface without any chemical bonding and mechanical
interlocking, is still an unresolved issue.
Sliding of the bar with respect to concrete indicates the transition from static to dynamic
slip. The load corresponding to this transition may be lower (if this occurs in the partially
debonded region) or above (if this occurs only after complete debonding along the entire
embedment length) Fd. For all specimens, the slope of the load-slip curves above Fd was
significantly lower than the slop below Fd. This suggests that the slip beyond Fd is dynamic slip
and that the measured Ff corresponds to dynamic friction. Any contribution from static friction
may be included in Fd. The Fd and Fa values measured for SC specimens were 3.59 ± 0.84 kN
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and 4.26 ± 0.24 kN respectively. The difference of 0.67 kN between these two average values is
taken to be Ff since the surface of the bar was smooth without any surface features for
mechanical interlocking. Since this is within the scatter for Fd and Fa , it appears like Fstatic =
Fdynamic and dynamic = static. Further confirmation of this using varying confinement pressure is
required. It should be noted that dynamic would vary for different bars.
4.1 Contribution from chemical bonding
Since debonding is complete at Fd, it would be expected to be a measure of chemical
bonding and the interfacial bond strength (d) determined using this would be expected to be
same for all specimens. However, it varied from a minimum of 1.01 MPa for SC specimen to a
maximum of 5.68 MPa for SCC specimen as shown in Fig. 9. Possible reasons for this variation
are:
(a) underestimation of contact surface area, especially in SCC and MC samples, since calculation
using nominal bar diameter would have excluded the increase in surface area due to surface
roughness
(b) contribution from the bearing resistance, especially in SCC and LC samples, since measured
Fd corresponds to debond completion rather than debond on-set.
The increase in Fd with
increase in the number of lugs in LC (i.e. RLC/TLC/SLC) specimens (Fig. 8) appears to support
this reasoning.
(c) contribution from static friction as discussed above. However, contribution from the dynamic
friction to Fd is believed to be minimal because the magnitude of load-end slip recorded at Fd is
very low. This is supported by the observation of clean bar surface without any scoring marks
during post-mortem examination of the specimens loaded to Fd.
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Nevertheless, the value of 1.01 MPa obtained for SC specimens would be the maximum
limit to the interfacial strength since the bearing resistance is zero and the frictional resistance is
negligible. This value is comparable to the published values. Since lateral pressure was not
determined, its contribution to Fd is unknown.
4.2 Contribution from frictional resistance
Frictional load varied with type of bar as shown in Fig. 8. Lateral pressure is the same for
all specimens. In order to eliminate the effect of difference in contact area, these frictional loads
are normalized with contact surface area and plotted in Fig. 10 as frictional stresses (f) for
various specimens. Comments made in section 4.1 regarding the error in the calculation of the
contact surface area for MC and SCC specimens are also applicable here.
It can be inferred from Fig. 10 that the frictional stress is negligible for SC specimens.
This is to be expected since the surface of bar in SC specimens was smooth. However, the MC
specimens had relatively rougher bar surface after removal of lugs in a lathe. This resulted in a
higher dynamic and frictional stress in MC specimens when compared to SC specimens. SCC
specimens with sand-coated bars recorded the maximum frictional stress and this is believed to
be due to higherdynamic for the sand – concrete interface when compared to lowerdynamic for the
relatively softer polymer composite – concrete interface. Since the SCC specimens had higher
contact area due to higher surface roughness, when compared to SC and MC specimens, any
error in the contact area used in determining f could also be another contributing factor to the
observed difference. The frictional stress values for LC specimens were same. This is to be
expected since dynamic and the surface roughness in these specimens would not change with the
pitch of lugs. Contribution from frictional resistance (hencedynamic) for LC specimens was
17
higher than that for the SC specimens but lower than that of MC specimens. This is believed to
be due to (a) longitudinal lugs with relatively higher surface roughness, (b) rough bar surface in
locations where lugs were removed to alter the pitch, and (c) the resistance offered by the
sheared lugs caught between the bar and concrete.
Since complete shearing of sand particles from the bar’s surface is observed in Fig. 7, it
can be concluded that Ff recorded for SCC specimens is not representative of the frictional
resistance encountered during loading to Fp. In addition, examination of dissected SCC
specimens loaded to load levels between Fd and Fp revealed progressive shearing of the sand
particles with increase in load [13]. This suggests that the frictional stress changed continually
during loading and the values in Fig. 10 represent the state of friction after the load drop at Fp.
Future studies should focus on characterizing this change in frictional stress during loading.
Since the change in friction stress during sliding and pullout was not characterized, error in the
values of Ff and Fb is unknown. Similar conclusion is valid for LC specimens since progressive
shearing of lugs was observed with increase in load.
4.3 Contribution from bearing resistance
The bearing resistance is negligible in SC and MC specimens due to lack of geometrical
features on the bar surface that can cause mechanical interlocking of the bar with the concrete.
However, the bearing resistance is substantial in SCC and LC specimens. It is interesting to note
(Fig. 8) that the bearing resistance due to small sand particles in SCC specimens is comparable to
the bearing resistance due to the lugs in SLC and TLC samples.
The bearing resistance increased from SLC to RLC due to decrease in the pitch of lugs
(i.e. increase in the number of helical lugs). In order to quantify the influence of the pitch and the
18
number of helical lugs on the bearing resistance, a number of LC specimens were unloaded after
loading to specific load levels between Fd and Fp, and were dissected and examined to determine
the number of sheared lugs. It can be inferred from Fig. 6 that all the lugs in the LC specimens
sheared at Fp. However, at loads less than Fp fewer helical lugs were sheared as shown in Fig. 11
for a RLC specimen loaded to 0.85Fp. At this load, only two lugs near the load end were sheared.
The sheared lugs appear milky-white while the intact lugs have the same color as the bar. It is
also interesting to note that the very first lug at the load end was not sheared. This was noticed in
all specimens with bars with lugs. Experimental studies [1,4] have shown that the maximum
interfacial shear stress occurs just below the concrete surface. Hence, it is to be expected that the
first lug near the concrete surface would not shear.
The applied load versus the number of helical lugs sheared at that load is shown in Fig.
12, for all specimens with varying number of lugs and pitch. Since the number of sheared lugs
was determined by dissecting the specimens unloaded from a load level, the load values may or
may not correspond to the actual load values at which the lugs sheared. Hence, the trend
suggested by the data in Fig. 12 is emphasized here rather than the absolute values. A linear
increase in the applied load with increase in the number of sheared lugs confirms that the
increase in load between Fd and Fp is mainly dependent on the bearing resistance offered by each
lug and independent of the pitch of lugs. The average shearing load per lug, given by the slope of
the linear-fit line in Fig. 12, is 2.59 kN. Using this value, the average shear strength (i.e. shear
strength of the bond between the lug and the bar) of each lug is determined (2.59 kN /da) to be
24.55 MPa. This value is similar to that determined by Al-Zahrani [1]. Hence, the bearing
resistance can be determined using equation (4).
19
Fb = l Al Nl
(4)
where l is the shear strength of a helical lug, Al is the contact surface area of a lug (da), and Nl
is the number of lugs per embedment length of the bar, which depends on the pitch.
Theoretically, it is possible to increase the bearing resistance and Fp by increasing Al and / or Nl
provided the concrete between the lugs can withstand the compressive and shear stresses
introduced in it by the applied load. This is supported by the results of Al-Zahrani et al. [3] , who
observed a change in failure mode from shearing of lug to fracture of concrete between the lugs
when the width of the lug was increased. Hence, the maximum bearing stress contribution from
one lug can be determined from equations (5a), and either (5b) or 5c depending on the lowest
Fmax.
Fb per helical lug = Fmax in the concrete
(5a)
where
Fmax = c Ac / sin2 for compressive failure
(5b)
Fmax = c Ac / (sin cosfor shear failure
(5c)
c is the compressive strength of the concrete, c is the shear strength of the concrete, and Ac is
the area of the concrete between the lugs ((D2- d2)/4). For the lug shape and dimensions shown
in Fig. 3, Fmax would be minimum for compressive failure even if the shear strength is assumed
to be one-half of the concrete compressive strength of 49.77 MPa. The calculated Fmax of 2.09
kN is within the error band for the slope (2.59 kN) obtained using experimental results plotted in
Fig. 12. However, no failure of concrete was observed between the lugs in the LC samples tested
20
in this study. Further investigation on the stress distribution along the embedment length and
near the lugs is required to resolve this.
Finally, the extrapolated value for Fp, for the case of no lug, is 28 kN which is
comparable to the value of 23.56  3.21 kN obtained for MC specimens. The average of Fd for
RLC, SLC and TLC is 18.22  3.1, which is indicated as average debond load in Fig. 12. The
difference between Fp and Fd, for the case of no lug, should be due to the contribution from the
frictional resistance, and the bearing resistance due to surface roughness. However, the latter can
be neglected based on the results for MC specimens in Fig. 8. Hence, the difference of about 10
 3.1 kN should be due to frictional contribution and it is comparable to frictional values plotted
in Figures. 8 and 9 for LC specimens, a minimum of 6.3 kN for TLC to a maximum of 11 kN for
RLC. Thus, the data in Fig. 12 independently support the research approach used in this study.
The bearing resistance also increased from SC to SCC due to increased surface
roughness. When the sand particles were completely sheared from the bar surface at Fp, as shown
in Fig. 7, the nominal diameter of the bar decreased from 13.6 mm to 12.25 mm. Hence, the
maximum bearing resistance in SCC specimens is limited by the shear strength of the bond
between the sand particles and the bar. Assuming that the entire bar surface is covered with sand
particles, bond strength is estimated to be 3.95 MPa (13520 N / (3.14 *12.25 mm *88.9 mm).
Thus, if the contact surface area between the bar and the sand particles is known, the bearing
resistance due to the sand particles may be predicted.
4.4 Effect of loading rate
The pullout load for SCC specimens increased with increase in loading rate as shown in
Fig. 13. However, such a clear trend was not observed in RLD specimens. This is believed to be
21
due to higher contribution from frictional resistance in SCC specimens when compared to RLD
specimens. Further investigation is required to clearly understand the relationship between the
loading rate, and the frictional resistance and the bearing resistance. Nevertheless these
preliminary results highlight the need to control the loading rate while evaluating the interfacial
bond strength using the pull out test.
In summary, the approach used in this study has clearly delineated the contributions from
various mechanisms to the interfacial bond strength and the pullout load (i.e. the resistance to bar
pullout). The results of this study have also demonstrated the dependence of these contributions
on parameters such as bar surface geometry, test conditions, and concrete strength. Since these
parameters varied from researcher to researcher in the past, the reported strength corresponding
to the pullout load also varied widely. Hence, instead of using this strength corresponding to
pullout load for design as suggested in the reviewed literature, the approach used in this study is
suggested. Equation (2) can be modified, for bars with lugs, as
Fp = d Ad+ l Al Nl +dynamic 
(6)
where Ad is the contact surface area between the concrete and the bar. Other parameters have
been defined before. The maximum limit for Fp is the force required to fracture the bar and this
can be calculated using the tensile strength and the cross sectional area of the bar. The areas, Ad,
and Al are functions of bar embedment length and bar diameter. If the lug dimensions and shape,
bar diameter, shear strengths, N, and dynamic in equation (6) are known, the critical embedment
length (i.e. development length) that would yield maximum Fp can be predicted using equation
22
(6). Alternatively, the dimension and shape of lugs wrapped on a bar, which would result in
maximum pullout load, can be designed for a chosen embedment length.
5. Conclusions
Debonding load, corresponding to debond completion, is much lower than the pullout
load, corresponding to onset of sliding and pullout of the bar along the entire embedment length.
While the former is mainly a measure of chemical bonding, the latter is a sum of contributions
from chemical bonding, frictional resistance, and bearing resistance. The maximum value for the
interface bond strength for the concrete – bar specimen used in this study is 1.01 MPa. The
resistance from lugs to bar pullout is comparable to that from sand particles bonded to the bar.
Bearing resistance due to lugs is a function of shear strength of a lug, lug dimensions, pitch of
the lug, and number of lugs, and is limited by the concrete’s compressive or shear strength.
Bearing resistance due to sand particles is a function of surface roughness amplitude and is
limited by the strength of the bond between the sand particles and the bar. Frictional resistance is
a function of surface roughness and it may vary during loading due to progressive shearing of
lugs or sand particles. Loading rate influences the pullout load.
Acknowledegment
The authors would like to acknowledge the financial support for this project from
NSERC (Natural Sciences and Engineering Research Council of Canada), the University of
Manitoba, and material and testing support from ISIS (Canadian Network of Centers of
Excellence for Intelligent Sensing for Innovative Structures). In addition, the authors would like
to acknowledge the technical assistance of Mr. Moray McVey of the University of Manitoba.
23
References
[1]
Al-Zahrani MM. Bond behavior of fiber reinforced plastic reinforcements with concrete.
Ph.D. Thesis. Universiy Park, Pennsylvania, Pennsylvania State University, USA; 1995.
[2]
Benmokrane B, Maceachern M, That TT, Zhang B. Evaluation of bond characteristics of
GFRP rebars embedded in polymer and normal cement concrete. Technical Report No. 7,
University of Sherbrooke, 1999.
[3]
Al-Zharani M, Mesfer M, Al-Dulaijian SU, Nanni A, Bakis CE, Boothby TE. Evaluation
of bond using FRP bars with axisymmetric deformations. Construction and Building
Materials, 1999; 13 (6): 299-309.
[4]
Benmokrane B, Tighiouart B,
Chaallal O. Bond strength and load distributions of
composite GFRP reinforcing bars in concrete. ACI Materials Journal May – June 1996;
pp. 246 - 253.
[5]
Tighiouart B, Benmokrane B, Gao D. Investigation on the bond of fiber reinforced
polymer (FRP) rebars in concrete. In: Saadatmanesh H. and Eshani MR editors.
Proceedings of the second international conference on composites in infrastructure,
Tucson, Arizona, USA, 1998; pp. 102-112.
[6]
Leung CKY Geng Y. Effect of lateral stress on the debonding and pullout of steel fibers in
a cementitious matrix. In: Buyukozturk O and Wecharatana M editors. Proceedings of the
1993 ACI Spring Convention, ACI Special Publication, SP-156, 1995; pp. 153 - 172.
[7]
Zenon Achillides, Kypros Pilakoutas. Bond behavior of fiber reinforced polymer bars
under direct pullout conditions. Journal of Composites for Construction March/April 2004;
pp. 173-181.
24
[8]
Roman Okelo, Robert L Yuan. Bond strength of fiber reinforced polymer rebars in normal
strength concrete. Journal of Composites for Construction May/June 2005; pp. 203-213.
[9]
Tastani SP, Pantazopoulou SJ. Bond of GFRP bars in concrete: experimental study and
analytical interpretation. Journal of Composites for Construction Sep./Oct. 2006; pp. 381391.
[10] Marta Baena, Lluis Torres, Albert Turon, Cristina Barris. Experimental study of bond
behavior between concrete and FRP bars using a pullout test. Composites Part B 2009; 40:
784-97.
[11] Masoud Esfandeh, Ali R Sabet, Amir M Rezadoust, Mohammed B Alavi. Bond
performance of FRP rebars with various surface deformations in reinforced concrete.
Polymer Composites 2009; pp. 576-582.
[12] Design and Control of Concrete Mixtures, 11th Edition, Portland Cement Association, IL,
USA.
[13] Wai How Soong. Bonding between the concrete and fiber reinforced plastic (FRP) rods.
M.Sc. Thesis. Winnipeg, Manitoba, University of Manitoba, Canada; 2001
25
LIST OF TABLES
Table 1. Published results on effect of embedment length on bond strength
Table 2. Published results on effect of bar diameter on bond strength
Table 3. Published results on effect of lateral pressure on bond strength
Table 4. Published results on effect of lugs on bond strength
Table 5. A comparison of published results to illustrate the influence of test conditions and
contact surface area definition on the interfacial bond strength
Table 6. Test specimen identification codes
LIST OF FIGURES
Figure 1. A schematic of load versus slip behavior during single fiber direct pull-out test
Figure 2. Four types of polymer composite reinforcing bars used in the present study
Figure 3. Shape and dimensions of lugs
Figure 4. A schematic of specimen dimensions and test set-up used in the present study
Figure 5. Representative load-slip curve for SC, MC. SCC, and SLC specimens
Figure 6. Complete shearing of all the lugs along the embedment length of a RLC specimen
loaded to Fa
Figure 7. Complete shearing of sand particles along the embedment length of a SCC specimen
loaded to Fa
Figure 8. Measured pullout load plotted as a sum of measured debond load, measured frictional
load, and calculated bearing load for various specimens.
Figure 9. Interfacial bond strength for various specimens
Figure 10. Frictional stress for various specimens
Figure 11. Partial shearing of lugs along the embeddment length of a RLC specimen loaded to
0.7Fa
Figure 12. Progression of lug failure in LC specimens with increase in applied load
Figure 13. Influence of loading rate on pullout load
26
Table 1. Published results on effect of embedment length on bond strength
Reinforcement
Bar Diameter
Average Bond
Strength
MPa
Reference
mm
Embedment
Length
mm
C-BARTM (P1)
12.7
12.7
63.5
127.0
18.4
14.5
Benmokrane et
al. [2]
H. B. Rebar
12.5
12.5
63.5
127.0
15.1
12.7
Machined
GFRP
12.7
12.7
63.5
127.0
39.7
29.7
Al-Zahrani [1]
Table 2. Published results on effect of bar diameter on bond strength
Reinforcement
Type A
Type B
Smooth GFRP
Nominal Bar
Diameter
mm
12.7
15.9
19.1
25.4
12.7
15.9
19.1
25.4
6.35
12.7
Embedment
Length
mm
127
127
127
127
127
127
127
127
63.5
63.5
27
Average Bond
Strength
MPa
10.6
7.3
6.6
6.4
12.3
10.8
-7.4
1.37
0.97
Reference
Tighouart et al.
[3]
Al-Zahrani [1]
Table 3. Published results on effect of lateral pressure on bond strength [1]
Curing
Temperature
Tc : C
Testing
Temperature
Tt :C
Interfacial Bond
Strength
MPa
60
20
<0.04
20
-20
<0.04
20
60
1.79
20
20
0.831
Table 4. Published results on effect of lugs on bond strength
Reference
Reinforcement
Description
Lug width
Lug Depth
mm
mm
Smooth
Smooth
--GFRP
Machined
Lug
3.8
1.3
GFRP
Machined
Lug
8.9
1.3
GFRP
Benmokrane
C-BAR
Lug
2.8
1.3
et al. [4]
(P1)
Note: Embedment Length = 63.5 mm; Bar Diameter = 12.7 mm
Al-Zahrani
[1]
28
Bond
Strength
MPa
0.97
39.7
23.2
18.4
Table 5. A comparison of published results to illustrate the influence of test conditions and
contact surface area definition on the interfacial bond strength
Reference
Reinforcement
Nominal
Lug
Dimension
Depth x
Width
mm
Nominal
Bar
Diameter
Loading /
Displacement Rate
Max.
Pullout
Load
mm
kN
Interfacial
Bond
Strength
MPa
Benmokrane et al. [2]
C-BARTM
(P1)
Commercial
Available
GFRP Bar
1.3 x 2.8
12.7
22
kN/ min
46.0
18.4
Al-Zahrani 3
Machined
GFRP Bar
1.3 x 3.8
12.7
0.125
mm/min
24.2
39.7
Table 6. Test specimen identification codes
Bar Type
Bar
Diameter
mm
RL
SL
TL
13.94
SC
M
S
13.6
12.0
12.5
Pitch
of
Lugs
mm
4.40
11.95
26.90
----
Bar Embedment Length in Pullout Specimen
(mm) and Test Specimen Codes
127.00
95.25
88.90
63.50
31.75
RLA
---MA
--
RLB
---MB
--
RLC
SLC
TLC
SCC
MC
SC
RLD
---MD
--
RLE
---ME
--
29
Fp
(a)
(b)
Load
A’
F d’
Ff ’
F i’
0
A
G
B
D
C
E
Ff
Fd
B’
G’
C’
D’
E’
Fi
0
Slip
Figure 1. A schematic of load versus slip behavior during single fiber direct pull-out test
Smooth bar
Machined bar
As received bar
Sand coated bar
Figure 2. Four types of polymer composite reinforcing bars used in the present study
30
D =13.94 mm
d = 12 mm
Helical Lug
b =0.95 mm
p = 4.4 mm
a = 2.8 mm
w = 2.8 mm

 » 76°
h = 1.23 mm
Longitudinal Lug
Figure 3. Shape and dimensions of lugs
MTS Actuator
Load Cell
Grip
LVDT
( Top)
Metal
plate
Fixed Debonder (6”)
12”
Embeddment Length
Variable Length
Debonder
LVDT
( Bottom )
6” diameter
Specimen
Ground
Figure 4. A schematic of specimen dimensions and test set-up used in the present study
31
70
60
Free End
SCC
Load End
Load (kN)
50
40
F
30
F
d
p
F
f
20
SLC
10
0
-30
MC
SC
-20
-10
0
10
20
30
Slip (mm)
Figure 5. Representative load-slip curves for SC, MC. SCC, and SLC specimens
Figure 6. Complete shearing of all the lugs along the embedment length of a RLC specimen
loaded to Fa
32
Figure 7. Complete shearing of sand particles along the embedment length of a SCC specimen
loaded to Fa
60
50
F
Load (kN)
b
40
30
F
f
20
F
10
d
0
RLC
SLC
TLC
SCC
MC
SC
Figure 8. Measured pullout load plotted as a sum of measured debond load, measured frictional
load, and calculated bearing load for various specimens
33
Interfacial Strength (MPa)
6
5
4
3
2
1
0
RLC
SLC
TLC
SCC
MC
SC
Figure 9. Interfacial bond strength for various specimens
Frictional Stress (MPa)
6
5
4
3
2
1
0
RLC
SLC
TLC
SCC
MC
Figure 10. Frictional stress for various specimens
34
SC
Figure 11. Partial shearing of lugs along the embedment length of a RLC specimen loaded to
0.7Fa
80
Load = 2.59* (number of lugs) +28.04
70
Load (kN)
60
50
40
30
MC Specimen
20
Average Debond Load = 18.22 kN
10
0
0
2
4
6
8
Number of Lugs
10
12
Figure 12. Progression of lug failure in LC specimens with increase in applied load
35
70
SCC
Pullout Load (kN)
60
50
RLD
40
30
20
10
0
0.26
1.3
6.5
Loading Rate (mm/ min)
Figure 13. Influence of loading rate on pullout load
36