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 bond

strength 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 careful

specimen 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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