Post on 30-Apr-2018
Experimental study on FRP-to-concrete bonded joints
J. Yaoa,b, J.G. Tengb, J.F. Chenc,*
aDepartment of Civil Engineering, Zhejiang University, Hangzhou, 310027, P.R. ChinabDepartment of Civil and Structural Engineering, The Hong Kong Polytechnic University, Hong Kong, China
cInstitute for Infrastructure and Environment, School of Engineering and Electronics, Edinburgh University, Alexander Graham Bell Building,
The King’s Buildings, Edinburgh EH9 3JN, UK
Received 16 January 2004; revised 2 June 2004; accepted 20 June 2004
Available online 7 August 2004
Abstract
The behaviour of bond between FRP and concrete is a key factor controlling the behaviour of concrete structures strengthened with FRP
composites. This article presents an experimental study on the bond shear strength between FRP and concrete using a near-end supported
(NES) single-shear pull test. The test results are found to be in close agreement with the predictions of Chen and Teng’s [J. Struct. Eng.
127(2001) 784] bond strength model, which mutually verifies the reliability of both the test method and the Chen and Teng model in general.
The NES single-shear pull test, given its simplicity and reliability, is therefore a good candidate as a standard bond test. The test results also
showed that Chen and Teng’s [J. Struct. Eng. 127(2001) 784] bond strength model is slightly conservative when the FRP-to-concrete width
ratios are at the two extremes, but this small weakness can be easily removed when more test results of good quality become available.
q 2004 Elsevier Ltd. All rights reserved.
Keywords: A. Polymer-matrix composites (PMCS); B. Debonding; B. Interface; Strength
1. Introduction
External bonding of fibre reinforced polymer (FRP)
composites has become a popular technique for strength-
ening concrete structures all over the world [2]. An
important issue in the strengthening of concrete structures
using FRP composites is to design against various
debonding failure modes, some of which were first studied
for concrete beams bonded with a steel plate, including: (a)
cover separation [3–5]; (b) plate end interfacial debonding
[3,4,6]; (c) intermediate (flexural or flexural-shear) crack
(IC) induced interfacial debonding [7] and (d) critical diagonal
crack (CDC) induced interfacial debonding [8–10].
The bond strength between FRP and concrete is a key
factor controlling debonding failures of various forms in
FRP-strengthened structures. As a result, extensive research
on this topic has been carried out, in addition to earlier work
concerned with steel plates bonded to concrete which
1359-8368/$ - see front matter q 2004 Elsevier Ltd. All rights reserved.
doi:10.1016/j.compositesb.2004.06.001
* Corresponding author. Tel.: C44-131-650-6768; fax: C44-131-650-
6789.
E-mail address: jchen@staffmail.ed.ac.uk (J.F. Chen).
provided a useful initial basis. The existing work has
included experimental studies conducted using single shear
tests, e.g. [11–15], double shear tests, e.g. [16–23] and
modified beam tests, e.g. [23–25], theoretical studies using
fracture mechanics analysis [15,26–33] and finite element
analysis [34,35], and the development of empirical
models [1,23,36,37]. A review of these studies can be
found in Ref. [1].
Existing studies suggest that the main failure mode of
FRP-to-concrete joints in shear tests is cracking of concrete
under shear, occurring commonly at a few millimetres from
the adhesive-concrete interface [1]. The bond strength (i.e.
the maximum transferable load) of the joint therefore
depends significantly on concrete strength. In addition, the
FRP-to-concrete member width ratio has a significant effect.
A very important aspect of the behaviour of these bonded
joints is that there exists an effective bond length beyond
which an extension of the bond length cannot increase the
ultimate load. This is the fundamental difference between
externally bonded reinforcement and internal reinforcement
for which a sufficiently long anchorage length can always be
found that the full tensile strength of the reinforcement can
Composites: Part B 36 (2005) 99–113
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J. Yao et al. / Composites: Part B 36 (2005) 99–113100
be achieved. The majority of existing studies have been
concerned with the prediction of the ultimate load and the
effective bond length [1].
This article presents an experimental study on the bond
shear strength between FRP and concrete using a near-end
supported (NES) single-shear pull test in which the concrete
prism is supported at the end nearer the applied load. These
tests have been conducted with the following purposes: (a)
to examine the reliability and robustness of the NES single-
shear pull test as a candidate standard bond test; and (b) to
verify the accuracy of the bond strength model recently
developed by Chen and Teng [1].
2. Test program
2.1. Test methods
A recent survey [33] showed that many different
experimental set-ups have been used for determining the
FRP-to-concrete bond strength, but no consensus on a
standard test procedure has been reached. Chen et al. [33]
classified the existing test set-ups into the following five
types: (a) double-shear pull tests; (b) double-shear push
tests; (c) single-shear pull tests; (d) single-shear push tests;
and (e) beam (or bending) tests. For better clarity, the first
four test methods are renamed here as: (a) far end supported
(FES) double-shear tests; (b) near end supported (NES)
double-shear tests; (c) far end supported (FES) single-shear
tests; and (d) near end supported (NES) single-shear tests
(Fig. 1). Collectively, all these four tests may also be
referred to as pull tests, as the plate is always directly pulled
by a tensile force.
FES double-shear pull tests and NES single-shear pull
tests have been the most popular test methods so far due to
Fig. 1. Classification of bond
their simplicity [33]. Both numerical [33] and experimental
[38] studies have shown that different test set-ups can lead to
significantly different test results. Within each test method,
small variations in the test set-up such as the height of the
support block in a NES single- or double-shear test may also
have significant effects based on a recent stress analysis
[33].
An FRP-to-concrete bond strength model is the key to the
accurate prediction of debonding failures in FRP-strength-
ened RC beams, including shear crack-induced debonding
failures [8,39] as well as intermediate flexural or flexural-
shear crack-induced debonding failures [7].
In debonding failures in FRP shear-strengthened RC
beams with transverse plates, the bond strength model
developed from pull tests is directly applicable [39]. Such a
model is also important in understanding the mechanism of
debonding induced by a critical diagonal crack near the end
of a longitudinal tension face plate for flexural strengthening
[8,10], where the longitudinal plate increases the concrete
component of the shear capacity and where the bond
strength developed from pull tests is also directly
applicable.
Furthermore, in intermediate crack-induced debonding
failures, the stress state in the critical region of the beam is
also closely similar to that of the concrete prism in a NES
single-shear pull test. The NES single-shear pull test
therefore appears to be a promising candidate as a standard
set-up for determining the FRP-to-concrete bond strength
and was therefore adopted in the present study. One of the
aims of the present experimental study is to examine the
effect of a number of small variations in this test set-up on
the resulting bond strength to aid in fine-tuning this test
method as a standard bond test method. Results from
previous NES single-shear pull tests also formed part of the
database on which Chen and Teng’s [1] recent bond strength
tests (Chen et al. 2001).
Fig. 2. Test specimen.
J. Yao et al. / Composites: Part B 36 (2005) 99–113 101
model was based, so the present test results also provide an
appropriate independent check of the validity of this bond
strength model.
Fig. 3. Relative vertical displacement between two sides of a flexural-shear
crack.
2.2. Specimen design
The NES single-shear pull test specimens consisted of a
concrete prism bonded with an FRP strip (Fig. 2). The
factors considered in the present test program include the
bond length Lfrp, the width ratio between the FRP strip and
the concrete prism bfrp/bc, the height of the concrete free
edge hc (Zheight of concrete prism hKheight of the
support block hb) (Fig. 2) and the offset in the load position
d. The first two factors have been identified to have a
significant effect on the bond strength but there have been
insufficient test data to rigorously verify the proposed
relationships [1]. The height of the concrete free edge hc
(Fig. 2b) has been shown to have a significant effect on the
stress distribution in the specimen [33], but its effect on the
ultimate bond strength is yet unclear. In practical pull tests,
there may be a small unintended offset d in the position of
the load (Fig. 2b). This offset may alternatively be expressed
as the initial loading angle q. The effect of this loading angle
needs to be understood if standardisation of the test set-up is
to be considered in the future. Furthermore, in flexurally
strengthened concrete structures, when debonding is
induced by the opening up of a flexural-shear crack, there
exists a relative vertical displacement between the two sides
of the crack, e.g. [40–42], so the FRP strip (or plate or sheet)
is loaded at a small positive (peeling) inclination angle to
the longitudinal axis on one side and at the same but
negative angle on the other side of the crack (Fig. 3). This is
thus another reason why the effect of a small loading angle
is worthy of some attention.
A total of 72 specimens in seven series were prepared
to investigate the effects of the above factors on the bond
strength (Table 1). The variables considered in Series I
(Specimens I-1–16) include the bond length Lfrp and the
support height hb (or height of the free concrete edge on
the loading side hcZhKhb). Series II (Specimens II-1–6)
and III (Specimens III-1–8) were designed to investigate
the effects of the loading offset and the FRP-to-concrete
width ratio respectively. Series IV–VII (Specimens IV-1–
14, V-1–12, VI-1–8 and VII-1–8) were designed following
the completion of the first three series to further explore
the effects of Lfrp, bfrp/bc and hc. Key parameters and test
results of all specimens are listed in Table 1. Specimens
II-1 and II-4 had a loading offset of dZ4 mm (equivalent
to an initial loading angle of 1.78) whilst Specimens II-3
and II-6 had a loading offset of dZK4 mm (equivalent to
an initial loading angle of K1.78). All other specimens
had no loading offset.
Concrete prisms of two different sizes were used. Half of
the specimens in Series III and V used 100!150!350 mm
concrete prisms so that a desired range of bfrp/bc ratios could
be achieved. All other specimens used 150!150!350 mm
concrete prisms. Concrete cubes and cylinders were tested
according to BS 1881 [43] to determine the material
properties at the time when the series of specimens made
from the same batch of concrete were tested.
GFRP was used in Specimens III-7 and III-8 while CFRP
was used in all others. The nominal thicknesses for the
CFRP and GFRP strips were 0.165 and 1.27 mm respect-
ively, the former being roughly the fibre sheet thickness
before resin impregnation with the latter being similar to the
thickness of the cured FRP strip. The FRP strips were
bonded to the concrete prisms following the manufacturer’s
instructions. The mechanical properties of the FRP
composites are shown in Table 2. The tensile strengths of
FRPs were determined according to ASTM D3039/
D3039M-95a [44] on the basis of the nominal thicknesses.
The nominal thicknesses were also used in all other
calculations of the present study. FRP composites were
Table 1
Details of specimens and test results
Test
specimen
Concrete
cylinder
strength f 0c(MPa)
Width of
concrete
prism bc
(mm)
FRP width
bfrp (mm)
FRP bond
length Lfrp
(mm)
Height of
free concrete
edge hc
(mm)
Test failure
load Ptest
(kN)
Test failure
mode
Predicted
failure load
Ppred (kN)
Ptest/Ppred
I-1 23.0 150 25 75 5 4.75 DB-C 5.72 0.83
I-2 23.0 150 25 85 5 5.69 DB-C 5.96 0.96
I-3 23.0 150 25 95 5 5.76 DB-C 6.02 0.96
I-4 23.0 150 25 95 5 5.76 DB-C 6.02 0.96
I-5 23.0 150 25 95 5 6.17 DB-C 6.02 1.02
I-6 23.0 150 25 115 5 5.96 DB-C 6.02 0.99
I-7 23.0 150 25 145 5 5.95 DB-C 6.02 0.99
I-8 23.0 150 25 190 5 6.68 DB-C 6.02 1.10
I-9 23.0 150 25 190 5 6.35 DB-C 6.02 1.05
I-10 23.0 150 25 95 75 6.17 DB-C 6.02 1.02
I-11 23.0 150 25 75 120 5.72 DB-C 5.72 1.00
I-12 23.0 150 25 85 120 6 DB-C 5.96 1.01
I-13 23.0 150 25 95 120 6.14 DB-C 6.02 1.02
I-14 23.0 150 25 115 120 6.19 DB-C 6.02 1.03
I-15 23.0 150 25 145 120 6.27 DB-C 6.02 1.04
I-16 23.0 150 25 190 120 7.03 DB-C 6.02 1.17
II-1 22.9 150 25 95 120 5.2 DB-C 6.02 0.86
II-2 22.9 150 25 95 120 6.75 DB-C 6.02 1.12
II-3 22.9 150 25 95 120 5.51 DB-C 6.02 0.92
II-4 22.9 150 25 190 120 7.02 DB-C 6.02 1.17
II-5 22.9 150 25 190 120 7.07 DB-C 6.02 1.17
II-6 22.9 150 25 190 120 6.98 DB-C 6.02 1.16
III-1 27.1 150 25 100 120 5.94 DB-C 6.27 0.95
III-2 27.1 150 50 100 120 11.66 DB-C 11.19 1.04
III-3 27.1 150 75 100 120 14.63 DB-C 15.02 0.97
III-4 27.1 150 100 100 120 19.07 DB-C 17.91 1.06
III-5 27.1 100 85 100 120 15.08 CPF 13.42 1.12
III-6 27.1 100 100 100 120 15.75 CPF 14.16 1.11
III-7 27.1 100 25.3 100 120 4.78 DB-C 4.92 0.97
III-8 27.1 100 50.6 100 120 8.02 DB-C 8.30 0.97
IV-1 18.9 150 25 95 5 5.86 DB-C 5.72 1.02
IV-2 18.9 150 25 95 5 5.9 DB-C 5.72 1.03
IV-3 19.8 150 25 95 5 5.43 DB-C 5.80 0.94
IV-4 19.8 150 25 95 5 5.76 DB-C 5.80 0.99
IV-5 18.9 150 25 95 15 5 DB-C 5.72 0.87
IV-6 19.8 150 25 95 15 7.08 DB-C 5.80 1.22
IV-7 18.9 150 25 95 30 5.5 DB-C 5.72 0.96
IV-8 19.8 150 25 95 30 5.93 DB-C 5.80 1.02
IV-9 18.9 150 25 95 45 5.38 DB-C 5.72 0.94
IV-10 19.8 150 25 95 45 6.6 DB-C 5.80 1.14
IV-11 18.9 150 25 95 60 5.51 DB-C 5.72 0.96
IV-12 19.8 150 25 95 60 5.67 DB-C 5.80 0.98
IV-13 18.9 150 25 95 90 6.31 DB-C 5.72 1.10
IV-14 19.8 150 25 95 90 6.19 DB-C 5.80 1.07
V-1 21.1 150 15 95 60 3.81 DB-C 3.71 1.03
V-2 21.1 150 15 95 60 4.41 DB-C 3.71 1.19
V-3 21.1 150 25 95 60 6.26 DB-C 5.89 1.06
V-4 21.1 150 50 95 60 12.22 DB-C 10.51 1.16
V-5 21.1 150 75 95 60 14.29 DB-C 14.10 1.01
V-6 21.1 150 100 95 60 15.58 DB-C 16.82 0.93
V-7 21.1 100 80 95 60 14.27 CPF 12.28 1.16
V-8 21.1 100 80 95 60 13.78 CPF 12.28 1.12
V-9 21.1 100 90 95 30 13.56 CPF 12.88 1.05
V-10 21.1 100 90 95 5 15.66 CPF 12.88 1.22
V-11 21.1 100 100 95 30 15.57 CPF 13.30 1.17
V-12 21.1 100 100 95 5 17.43 CPF 13.30 1.31
VI-1 21.9 150 25 95 60 6.01 DB-I 5.95 1.01
VI-2 21.9 150 25 95 60 5.85 DB-I 5.95 0.98
VI-3 21.9 150 25 145 60 5.76 DB-I 5.95 0.97
VI-4 21.9 150 25 145 60 5.73 DB-I 5.95 0.96
(continued on next page)
J. Yao et al. / Composites: Part B 36 (2005) 99–113102
Table 1 (continued)
Test
specimen
Concrete
cylinder
strength f 0c(MPa)
Width of
concrete
prism bc
(mm)
FRP width
bfrp (mm)
FRP bond
length Lfrp
(mm)
Height of
free concrete
edge hc
(mm)
Test failure
load Ptest
(kN)
Test failure
mode
Predicted
failure load
Ppred (kN)
Ptest/Ppred
VI-5 21.9 150 25 190 60 5.56 DB-I 5.95 0.93
VI-6 21.9 150 25 190 60 5.58 DB-I 5.95 0.94
VI-7 21.9 150 25 240 60 5.91 DB-I 5.95 0.99
VI-8 21.9 150 25 240 60 5.05 DB-I 5.95 0.85
VII-1 24.9 150 25 95 60 6.8 DB-C 6.14 1.11
VII-2 24.9 150 25 95 60 6.62 DB-C 6.14 1.08
VII-3 24.9 150 25 145 60 7.33 DB-C 6.14 1.19
VII-4 24.9 150 25 145 60 6.49 DB-C 6.14 1.06
VII-5 24.9 150 25 190 60 7.07 DB-C 6.14 1.15
VII-6 24.9 150 25 190 60 7.44 DB-C 6.14 1.21
VII-7 24.9 150 25 240 60 7.16 DB-C 6.14 1.17
VII-8 24.9 150 25 240 60 6.24 DB-C 6.14 1.02
Average 1.04
CoV 9.6%
Note: (a) CFRP was used in all specimens except III-7 and III-8 in which GFRP was used; (b) all concrete prisms had a height of 150 mm; (c) concrete cylinder
strength determined from cube strength according to fcLZ0.79 fcu
L (d) DB-C, debonding in concrete; DB-I, debonding at adhesive-concrete interface; CPF,
Concrete prism failure.
Table 2
Properties of FRPs
Type Thickness
(mm)
Tensile
strength ffrp(MPa)
Elastic mod-
ulus Efrp
(GPa)
Ultimate
tensile
strain 3frp
(%)
CFRP 0.165 4114 256 1.61
GFRP 1.27 351 22.5 1.56
J. Yao et al. / Composites: Part B 36 (2005) 99–113 103
bonded to the concrete prisms with epoxy resins. More
details of the material properties and specimen preparation
procedures are available in Ref. [45].
2.3. Test set-up
A steel rig for NES single-shear pull tests (Fig. 4a) was
carefully fabricated to carry out all the tests reported in this
article. In this rig, the load could be accurately positioned
vertically by adjusting the height of the bearing plate.
Different support blocks could be used to achieve the
required support heights on the loaded end (i.e. the end
nearer the applied load or the near end) of the concrete
prism. A positioning frame was used to prevent the far end
of the concrete prism from uplifting. The concrete prism
was separated from the positioning frame by a thin layer of
rubber to allow horizontal sliding of the concrete prism.
2.4. Instrumentation and loading procedure
Strain gauges and LVDTs were used to measure strains
in the FRP and displacements at various positions. Details of
these measurements are not given here, but are available
elsewhere [45], as the main concern of the present paper is
with the bond strength. Loading was applied through a
hydraulic jack at increments of about 5% of the ultimate
load predicted by Chen and Teng’s model [1]. Fig. 4b shows
a specimen during the test.
Fig. 4. Test rig.
3. Chen and Teng’s bond strength model
As the specimens were designed based on Chen and
Teng’s bond strength model [1] and the results are
Fig. 5. Debonding in concrete.
J. Yao et al. / Composites: Part B 36 (2005) 99–113104
compared with its predictions later in the article, it is
necessary to introduce this model before the test results are
presented. The bond strength expressed as per unit width of
the FRP strip, qu, is
q ZPu
bfrp
Z abwblLe
ffiffiffiffif 0c
p(1)
where Pu is the ultimate load in N, bfrp is the width of the
FRP strip in mm, bw and bl are dimensionless coefficients
reflecting the effects of the FRP-to-concrete width ratio bfrp/
bc and the bond length Lfrp respectively, Le is the effective
bond length in mm and f 0c is the cylinder compressive
strength of concrete in MPa. Based on the regression of test
data collected from the literature, Chen and Teng [1]
obtained the best fit value of aZ0.427. It was proposed to
use the 95th percentile of aZ0.315 as the lower bound for
design. bw, bl and Le are given by
bw Z
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi2 Kbfrp=bc
1 Cbfrp=bc
s(2)
b1 Z
1:0 if LfrpRLe
sinp
2
Lfrp
Le
if Lfrp!Le
8><>: (3)
Le Z
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiEfrptfrpffiffiffiffi
f 0cp
s(4)
in which Efrpand f 0c are in MPa while tfrp and Le are in mm.
Fig. 6. Debonding at the adhe
It may be noted that this model was developed based on
a fracture mechanics solution [31] with rational simplifica-
tion, with its coefficient regressed from a set of either
single-shear or double-shear pull tests on FRP and steel plate-
to-concrete bonded joints. The model is thus semi-empirical
and generic, being applicable to both FRP (wet lay-up or
prefabricated) and steel plates, subject to the condition that
failure is not due to yielding of steel or rupture of FRP. The
model was developed for debonding failure in the concrete,
but may also be applicable to debonding failure at the
adhesive–concrete interface as shown later.
4. Test results and discussions
4.1. Failure modes
Fifty-six out of the 72 specimens failed due to debonding
in concrete adjacent to the adhesive-concrete interface in
which a thin layer of concrete is attached to the FRP strip
after failure (Fig. 5). It may be noted that this is not strictly
‘debonding’ because the failure actually occurs in concrete.
Nevertheless, the term is still adopted here because it has
been widely used by the research community as discussed in
Ref. [1]. Eight specimens failed by debonding at the
adhesive-concrete interface where much less concrete is
attached to the FRP strip after failure (Fig. 6). The
remaining eight specimens failed in the concrete prism by
the formation of a fracture plane that starts at the far end
of the FRP strip and extends to the top of the support
sive–concrete interface.
Fig. 7. Failure in the concrete prism.
J. Yao et al. / Composites: Part B 36 (2005) 99–113 105
block (Fig. 7). The failure mode of each specimen is
indicated in Table 1.
4.1.1. Debonding in concrete
For those specimens which failed by debonding in
concrete, the failure process started with visible concrete
cracking near the loaded end of the concrete prism. The
surface cracks observed on the sides of the FRP strip were
at about 458 to the longitudinal axis of the FRP strip. As
the load increased, visible cracking in the concrete initiated
debonding of the FRP from the concrete at the loaded end.
Debonding then propagated towards the far end of the FRP
strip and eventually led to the complete detachment of the
FRP strip from the concrete. The duration of this
debonding process depended on the bond length of the
FRP strip. It was very short or could not be noticed at all
for a small bond length but could be easily seen for a long
one. Debonding was due to failure in the concrete at a
small distance beneath the adhesive-concrete interface. A
lump of concrete from the loaded end was generally
attached to the debonded FRP strip, while a smaller
concrete lump was sometimes found at the far end of the
FRP strip (Fig. 5). The thickness of the concrete layer on
the debonded FRP strip elsewhere varied approximately
between 1 and 5 mm. The surface of the failure zone of the
concrete prism was uneven, with the aggregate being
clearly visible (Fig. 5). The phenomenon that more
concrete was usually attached to the FRP strip at both
ends of the interface than elsewhere may be related to the
stress concentration at the ends [34].
4.1.2. Debonding at the adhesive–concrete interface
The failure process was almost same as the above, but the
failure was mostly along the adhesive–concrete interface.
The FRP strip generally also pulled off a lump of concrete at
the loaded end. Much less or little concrete was attached to
the FRP strip elsewhere (Fig. 6).
This failure occurred only in eight Series VI speci-
mens. It may be noted that Series IV, V and VI
specimens were prepared by an assistant with limited
experience under the supervision of an experienced
researcher. There was some uncertainty with regard
to the surface preparation of the concrete prisms and
the mixing of the primer [45]. In light of this, the
concrete prisms used in Series VI were reused in a
following series of tests (series VII) with the FRP strip
bonded to the opposite side of the prism, while all other
parameters remained unchanged. All Series VII speci-
mens failed by debonding in concrete, confirming that
the results of Series VI had been influenced to some
extent by interfacial weakness introduced during
preparation.
4.1.3. Concrete prism failure
Specimens III-5 and III-6, and V-7–12 failed in the
concrete prism. The failure started by the initiation of a
crack in the concrete prism near the far end of the FRP strip.
Once the crack appeared, it propagated almost immediately
towards the upper end of the support block and the specimen
failed (Fig. 7). The FRP–concrete interface was intact after
failure and the failure process was catastrophic.
All specimens failed in the concrete prism had a concrete
width of 100 mm (compared with 150 mm for most of other
specimens) and an FRP strip which was quite wide (bfrp/bc
R0.8). This failure is obviously more likely when the FRP
strip and the concrete prism have similar widths, and is more
a consequence of the test set-up than any other factors. The
use of the positioning frame to prevent the concrete prism
from uplifting (Fig. 4a) introduces tensile bending stresses at
the upper surface of the concrete prism at the far end of the
FRP strip, while the use of a low support block allows
formation of a fracture plane at a relatively low load.
4.2. Load–displacement behaviour
Fig. 8 shows the load–displacement curves of Series VI
(which failed due to debonding at the adhesive–concrete
interface) and VII specimens (which failed due to debond-
ing in the concrete). The displacement was measured at the
right end of the grip (Fig. 4a) so it includes not only the
displacement due to interfacial slip, but also a number of
other components such as the elastic deformation of the un-
bonded part of the FRP strip and possible slip of the FRP
strip in the grip. Therefore, these curves shall only be treated
as qualitative information reflecting the global load–
displacement response.
Initially, the displacement increases almost linearly with
the load and the slopes of different curves are similar. Faster
increases in the displacement indicate the initiation of
micro-cracking at the loaded end. Substantial differences
between the curves are observed for the later stage of
loading as failure was approached and these differences are
attributable to different bond lengths and different failure
modes. For Series VI specimens with debonding at the
adhesive-concrete interface, all curves feature a plateau
before ultimate failure, and the length of the plateau
increases with the bond length. For Series VII specimens
with debonding in the concrete, only those for a long bond
length feature such a plateau and it is much short than that of
Fig. 8. Load–displacement curves: Series VI and VII specimens.
J. Yao et al. / Composites: Part B 36 (2005) 99–113106
Fig. 9. Strain distribution along the FRP strip for Specimen I-1: LfrpZ75 mm.
J. Yao et al. / Composites: Part B 36 (2005) 99–113 107
the corresponding specimen failed in FRP debonding at
concrete/adhesive interface.
4.3. Strain distributions in FRP
Figs. 9 and 10 show typical distributions of strains in
the FRP strip. These strains were found from strain
gauges mounted on the upper surface of the FRP strip,
except the strains at xZ0 which were deduced directly
from the applied load and the geometric and material
properties of the FRP strip, as readings from the strain
gauge at this location were found to be significantly
affected by local bending of the strip. When the applied
load P is smaller than about 60% of the ultimate load Pu,
the FRP strain is minimal beyond a small distance of
about 0.5Le from the loaded end (Figs. 9a and 10a),
indicating that almost all the applied load is resisted
within this small area. Here Le is the effective bond
length according to Chen and Teng’s model [1].
For Specimen I-1 with a small bond length (LfrpZ75 mm), the increase of FRP strain is gradual until P reaches
0.89Pu (PuZ4.75 kN) (Fig. 9b). Cracking at the loaded end
was first observed by naked eyes (i.e. visible cracking) at
PZ4.5 kN (P/PuZ0.95). This cracking led to an obvious
change of the strain distribution in the FRP strip indicating
the propagation of debonding, and the specimen failed soon
thereafter. The strain in the debonded part of the FRP strip is
seen to be almost constant.
For Specimen I-16 with a large bond length (LfrpZ190 mm), visible cracking occurred at a similar load (i.e.
PZ4.75 kN) but ultimate failure occurred at a higher load
(PuZ7.03 kN). The propagation of debonding is more
clearly reflected by the strain distribution as shown in
Fig. 10b. It may be noted that a large part of the FRP strip
near the far end still had minimal strain when the ultimate
load was reached, confirming the concept of effective bond
length implying that increasing the bond length beyond a
certain value does not further increase the bond strength.
Fig. 10. Strain distribution along the FRP strip for Specimen I-16: LfrpZ190 mm.
J. Yao et al. / Composites: Part B 36 (2005) 99–113108
However, a larger bond length can be expected to lead to a
longer deformation process as debonding propagates along
the interface.
Careful inspection of Figs. 9 and 10 reveals that local
debonding near the loaded end occurred much earlier than
was observed by naked eyes. Fig. 9a shows that there is a
significant change in the local strain distribution near the
loaded end (xZ0) when the applied load increases from
0.21Pu to 0.38Pu. When PZ0.21Pu, the deduced axial strain
at xZ0 is significantly larger than that measured on the
upper surface of FRP at xZ0.1Le. The strain decreases fast
away from the loaded end. When the load increases to over
0.38Pu, the deduced strain at xZ0 becomes slightly smaller
than that measured at xZ0.1Le and this pattern remains
unchanged until failure. This phenomenon is believed to be
due to very local debonding (not visible to naked eyes) that
occurred before the applied load reached 0.38Pu. This local
debonding moves the effective point of stress transfer from
the FRP to the concrete by a small distance (less than
0.1LeZ9.4 mm in this case) towards the free end of the FRP
strip. This phenomenon has also been noted by Yuan et al.
[33] and may be attributed to local stress concentration near
the loaded end [33,34]. The same phenomenon is evident
from the strain distributions shown in Fig. 10 for Specimen
I-16, where local debonding appears to have occurred at a
load P less than 0.31Pu (Fig. 10a).
4.4. Effect of height of free concrete edge
Test results from Series I and II for various heights of the
free zone at the near end of the concrete prism (i.e. height of
the free concrete edge hcZhKhb) are shown in Fig. 11a
and b. It is seen that the bond strengths of specimens with
hcZ120 mm are consistently larger than those with
hcZ5 mm, with the difference being of the order of 10%.
This indicates that the height of free concrete edge does
Fig. 11. Effect of height of free concrete edge.
J. Yao et al. / Composites: Part B 36 (2005) 99–113 109
have some effect on the bond strength, which is in
agreement with previous numerical observations [34].
This is because the local stiffness near the loaded end is
increased when the top of the support block is closer to the
FRP plate (smaller hc values). This increased local stiffness
attracts an increased local stress transfer from the FRP to the
concrete there, leading to early debonding and hence a
reduced bond strength.
Numerical results from linear elastic analysis [34] have
shown that there is a range in which the interfacial stress
distribution is insensitive to hc. The test data shown in
Fig. 11a and b do not allow the identification of such a range
for hc, as they only cover three values of hc. Series IV was
thus designed to further explore this issue, as this
information is useful for the development of a standard
bond test method. However, no definite conclusion can be
drawn from the results of Series IV (Fig. 11c) because they
show a relatively large scatter which may be attributed to
the less stringent specimen preparation procedure of these
specimens as discussed earlier in the article.
4.5. Effect of bond length
Fig. 12 shows the relationship between the FRP bond
strength and the bond length for all the specimens with the
same CFRP width of 25 mm and no loading offset from
Series I, II, VI and VII. The predictions of Chen and Teng’s
model [1] are also shown for comparison. These test results
clearly support the concept of an effective bond length and
the accuracy of the effective bond length formula of Chen
and Teng’s model.
It is seen that the test results from Series I with hcZ5 mm
and those from Series VI are slightly below Chen and
Teng’s predictions, whilst those from Series VII are above
the predictions. Overall, the experimental results are nicely
scatted around the predictions of Chen and Teng’s model.
This observation agrees with expectation because Chen and
Teng’s model was developed to provide best-fit predictions
of test data collected from the literature which are expected
to have been obtained from slightly different test set-ups and
by different researchers.
It should be noted that Series VI specimens failed due to
debonding at the adhesive–concrete interface as a result of a
less stringent specimen preparation procedure, while the
more carefully prepared specimens of Series VII failed due
to debonding in concrete. Clearly, Series VII results are
significantly higher than those of Series VI, further
confirming the importance of careful specimen preparation.
4.6. Effect of loading offset
Fig. 13 shows that both a positive and a negative loading
offset have a significant effect on the bond strength when the
bond length is small (LfrpZ95 mm). The loading offsets of
G4 mm (i.e. initial loading angles of G1.78) reduced the
bond strength significantly. This may reflect the effect of the
loading angle on the local stresses near the loaded end. It is
shown in Ref. [46] that both small positive and negative
angles increase the principal tensile stress locally and thus
may have a detrimental effect on the bond strength.
The effect of a small loading angle is insignificant for a
relatively long bond length of 190 mm (Fig. 13). A possible
explanation for this phenomenon may be as follows. For a
positive loading angle, as debonding propagates, the loading
angle and thus its effect reduces. When the loading angle is
negative, two factors may contribute to this phenomenon.
First, assuming that the bond length is sufficiently large, the
debonding crack appears first at the loaded end and then
Fig. 12. Effect of bond length.
J. Yao et al. / Composites: Part B 36 (2005) 99–113110
progresses towards the far end of the FRP strip, in contrast
to specimens with a short bond length in which complete
failure is reached immediately when debonding starts. Once
the debonding crack has progressed by a small distance, the
effect of a negative loading angle disappears as the
debonded portion of the FRP strip has to remain in contact
with the concrete. Second, a negative loading angle results
in compressive normal stresses on the debonded area, which
produce frictional forces to help resist the applied load.
Therefore, a small negative loading angle is expected to
have no detrimental effect on the bond strength if the bond
length is sufficiently large.
These test results illustrate the importance of a reliable
set-up for the determination of bond strength in a pull test.
Since a small loading offset is hard to avoid, the bond length
of the FRP strip in a bond test specimen should be
sufficiently long to minimise the effect of a loading offset.
They also imply that in the flexural strengthening of beams
and slabs, it is important to provide a sufficient bond
(anchorage) length so that the effect of relative vertical
displacements between the two sides of a flexural-shear
crack can be minimised.
Fig. 13. Effect of loading offset displacement.
4.7. Effect of FRP-to-concrete width ratio
Fig. 14 shows the effect of the FRP-to-concrete width
ratio bfrp/bc on the bond strength. It is seen that Chen and
Teng’s model [1] underestimates slightly the bond strength
both when bfrp/bc is close to 0 or 1. It may be noted that the
specimens with bfrp/bc close to 1 failed in concrete prism so
their actual bond strength should be even higher than the test
results. Whilst Eq. (2) may be modified to slightly better fit
the data points shown in Fig. 14, such a modifications is
statistically insignificant for a combined database contain-
ing test data presented in this article and those in Ref. [1]. As
there is still a lack of high quality test data at both extremes
of bfrp/bc (i.e. close to 0 and 1), such a modification is not
attempted here.
5. Comparison with Chen and Teng’s predictions
A comparison between the present test data and the
predictions of Chen and Teng’s model [1] is shown in
Fig. 15. Statistics of the test-to-predicted bond strength ratio
are given in Table 3. Here the effects of the height of free
concrete edge at the loaded end and the loading offset are
treated as factors contributing to the experimental scatter. It
is seen that Chen and Teng’s model [1] underestimates the
bond strength by 4% on average for failure by debonding in
the concrete, but overestimates the bond strength by the
same percentage for failure by debonding at the adhesive-
concrete interface, with the coefficient of variation being
less than 10% in both cases (Table 3). This comparison
confirms that Chen and Teng’s model [1] represents very
closely the bond strength overall.
The average test-to-predicted bond strength ratio and its
standard deviation for the complete data set containing both
Fig. 14. Effect of FRP-to-concrete width ratio.
J. Yao et al. / Composites: Part B 36 (2005) 99–113 111
specimens which failed by debonding in the concrete and
those which failed by debonding at the adhesive-concrete
interface are 1.03 and 0.093, respectively. The a value in
Eq. (1) for the 95th percentile can be easily found to be 0.37
(Z0.427!(1.03K1.64!0.093)) which is about 20% larger
than the value of 0.315 obtained from the database presented
in Ref. [1]. This is understandable because the standard
deviation of the test data presented here is (and should be)
smaller than that of the data presented in Ref. [1] which
were obtained by different researchers. As the actual quality
variations at practical construction sites with different
application personnel may be larger than those experienced
in laboratory tests, aZ0.315 is still recommended here as a
conservative value for design use. However, a more precise
value may be proposed when sufficient confidence in such a
value has been gained with more extensive research.
It may be noted that the bond strength model was not
developed for the concrete prism failure mode. This failure
mode can be prevented through the use of a higher support
block and a longer bond length. It is desirably to avoid this
and other failure modes which do not have a direct bearing
on the interfacial behaviour of FRP-to-concrete bonded
joints in a standard bond test procedure.
Fig. 15. Comparison with Chen and Teng’s (2001) model.
6. Conclusions
This article has presented an experimental study on the
bond shear strength between FRP and concrete using a near-
end supported (NES) single-shear pull test in which the
concrete prism is supported at the end nearer the applied
load. These tests have been conducted with the following
purposes: (a) to examine the reliability and robustness of the
NES single-shear pull test as a candidate standard bond test;
and (b) to verify the accuracy of the bond strength model
recently developed by Chen and Teng [1]. The results and
discussions presented in the present article allow the
following conclusions to be made.
(1)
Tabl
Stati
Test
(1) D
(56 s
(2) D
conc
(eigh
(1)C
(3) C
spec
All
Since the NES single-shear pull test as presented in this
article produced results which are in close agreement
with the predictions of Chen and Teng’s model [1], the
reliability of both the test method and the Chen and Teng
model are mutually verified in general. The NES single-
shear pull test, given its simplicity and reliability, is
therefore a good candidate as a standard bond test. This
test method is also robust provided a sufficiently long
bond length is employed to minimise the effect of
unintended loading offsets and a sufficiently high support
block is used to avoid non-interfacial failures. Based on
the present test results, it may be recommended that the
bond length in a standard test should be around two times
e 3
stics of test-to-predicted bond strength ratios
failure mode Average Standard
deviation
CoV
(%)
ebonding in concrete
pecimens)
1.04 0.093 8.9
ebonding at the adhesive–
rete interface
t specimens)
0.96 0.050 5.2
(2) 1.03 0.093 9.0
oncrete prism failure (eight
imens)
1.16 0.078 6.7
1.04 0.100 9.6
J. Yao et al. / Composites: Part B 36 (2005) 99–113112
the effective bond length specified by Chen and Teng’s
model and the height of the free concrete edge should be
around 50 mm for a concrete prism of 150 mm in height.
In addition, the distance between the positioning frame
preventing the uplifting of the concrete prism and the far
end of the FRP strip should be appropriate to avoid high
flexural tensile stresses near the far end of the FRP strip
as well as interference with interfacial behaviour.
(2)
The test results showed that Chen and Teng’s bondstrength model [1] is slightly conservative when the
FRP-to-concrete width ratios are at the two extremes of
0 and 1. When more reliable test results become
available, this small weakness can be easily removed.
(3)
The test results highlighted the importance of carefulspecimen preparation as the results can be significantly
affected. In the development of a standard bond test
procedure, measures should be included to minimise
this effect, while in the development of design methods,
due allowance should be made for the expected quality
variations at sites.
Acknowledgements
The authors are grateful for the financial support from
The Hong Kong Polytechnic University (G-V784) and from
the Research Grants Council of the Hong Kong SAR (PolyU
5151/03E).
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