Debonding resistance of FRP-to-clay brick masonry joints

Debonding resistance of FRP-to-clay brick masonry joints

J. Kashyap a, C.R. Willis a, M.C. Griffith a,⇑, J.M. Ingham b, M.J. Masia c

a School of Civil, Environmental and Mining Engineering, The University of Adelaide, Adelaide, SA 5005, AustraliabDept. of Civil and Environmental Engineering, The University of Auckland, Auckland, Private Bag 92019, New ZealandcCentre for Infrastructure Performance and Reliability, The University of Newcastle, Faculty of Engineering, Callaghan, NSW 2308, Australia

a r t i c l e i n f o

Article history:

Received 21 May 2011

Revised 15 March 2012

Accepted 16 March 2012

Keywords:

Masonry

FRP

Strengthening

Debonding

Bond pull-tests

Bond strength model

a b s t r a c t

Debonding at the FRP-to-masonry interface has been identified as the preferred failure mechanism in

fibre-reinforced polymer (FRP) retrofitted masonry as it allows for some redistribution of forces. The

results of 14 FRP-to-masonry bond tests are presented, where the FRP was near surface mounted

(NSM) to stack-bonded clay brick masonry. These tests were conducted to investigate the effect that vari-

ables such as cyclic loading and FRP strip dimensions have on the debonding resistance of a NSM FRP-to-

masonry joint. These results were then incorporated into a large database of FRP retrofitted masonry pull

test results by various researchers over the past 10 years. The database includes results for both exter-

nally bonded (EB) and NSM retrofitting techniques. From this database, local bond–slip parameters such

as the maximum interface shear stress, smax, and the maximum slip, dmax, were investigated to determine

correlations between these values and masonry material properties. Further, 15 existing concrete and

masonry bond strength (maximum load at the FRP-to-substrate interface) models in the literature were

assessed for their use with masonry by comparing these models against the results in the pull test data-

base. Based on the comparative statistics of the test-to-predicted bond strength it is concluded that a new

FRP-to-masonry bond model is required which gives more accurate predictions. Results include a discus-

sion on the global load–slip response and FRP-to-masonry interface behaviour.

Ó 2012 Elsevier Ltd. All rights reserved.

1. Introduction and background

Unreinforced masonry (URM) structures constitute both a sig-

nificant portion of the world’s heritage buildings and a significant

component of the modern residential building stock, and are par-

ticularly susceptible to damage from out-of-plane loads such as

those generated by earthquakes [1]. Over the past decade or so,

the use of near surface mounted (NSM) fibre reinforced polymers

(FRPs) has gained much attention as a promising new seismic

strengthening technique [2–6]. The NSM FRP retrofitting technique

(i.e. inserting FRP strips into grooves cut into the surface of a wall)

provides significant advantages over externally bonded (EB) FRP

such as improved aesthetics, reduced surface preparation and bet-

ter protection from UV exposure and vandalism. Importantly, NSM

FRP debonds at higher strains than EB FRP and thus leads to more

efficient use of the FRP material [7]. When using the NSM tech-

nique, the most efficient cross-section is a thin rectangular strip

as its benefits, in terms of efficiency and construction time, are

superior to those of other shapes such as bars [8].

Some of the common out-of-plane failure mechanisms of FRP

strengthened masonry walls include sliding of the masonry units,

flexural-shear cracking, FRP rupture, FRP debonding, punching

shear and crushing of brick in compression [9–14]. Among these

debonding mechanisms, intermediate crack (IC) debonding gov-

erns the increase in moment capacity and sectional ductility. IC

debonding involves progressive detachment of the NSM FRP strip

which initiates at the location of intermediate flexural or flex-

ural-shear cracks when the strip is subjected to large tension

stress. If the FRP is perfectly attached to the masonry then theoret-

ically the FRP strip requires infinite strain capacity in order to

bridge the intercepting cracks. As such strain levels are not possi-

ble, debonding cracks will occur at the FRP-to-substrate interface

and gradually propagate towards the strip ends [15,16] as long

as the debonding strain is lower than the tensile rupture strain

for the FRP.

IC debonding may be further divided into two types: (i) associ-

atedwith the presence of a single crack in amember; and, (ii) where

multiple cracks are distributed along the length of a member

[17,18]. The stress state at the interface for type (i) (refer Fig. 1a)

can be idealised using a pull test where a tensile force (P) is applied

to the FRP strip causing a slip (D) at the crack face as shown in

Fig. 1(b). The resulting P–D response is referred to as the global

load–slip response which depends on the interfacial bond charac-

teristics between the FRP and the surrounding masonry substrate

such as the interface shear stress, s and the local interface slip, d.

The local s–d response is known as the local bond–slip behaviour.

Pull tests are a useful and relatively inexpensive testing procedure

0141-0296/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved.

http://dx.doi.org/10.1016/j.engstruct.2012.03.032

⇑ Corresponding author. Tel.: +61 8 8303 5451; fax: +61 8 8303 4359.

E-mail address: [email protected] (M.C. Griffith).

Engineering Structures 41 (2012) 186–198

Contents lists available at SciVerse ScienceDirect

Engineering Structures

journal homepage: www.elsevier .com/locate /engstruct

that can be used to investigate the interface s–d behaviour aswell as

to determine the critical bond length, Leff, the strain in the FRP strip

at the onset of debonding, edb, and other factors impacting the bond

strength. Furthermore, the debonding resistance obtained from

FRP-plated pull tests is a lower bound to the IC debonding resistance

in beams due to effects of moment and crack distribution [17–20].

Hence, to understand the behaviour of a FRP-strengthened URM

wall subjected to out-of-plane loading, the FRP-to-masonry joint

behaviour must be understood first [18,19,6,5].

Due to material similarity between concrete and masonry, such

as low tensile strength and brittleness, the debonding mechanisms

for retrofitted masonry have been shown to be similar to those of

retrofitted reinforced concrete (RC) members (Fig. 2). Moreover,

factors affecting the FRP-to-concrete bonded joint behaviour simi-

larly influence FRP-to-masonry bonded joint behaviour [19,6]. The

use of FRP to improve the flexural resistance of RC members is now

well established (e.g. [21,15]). Hence, the research on the debond-

ing mechanisms in plated RC structures has been used as a starting

point for research on FRP retrofitted masonry structures.

A large database of pull test results is presented, reporting pre-

viously published data for masonry retrofitted with FRP strips. The

derivation of analytical models to determine the IC debonding

resistance of bonded joints is also presented. Empirical relation-

ships for key parameters such as the fracture energy, Gf, and shear

stress–slip (denoted as local bond–slip) relationship, s–d, are

investigated. Whilst the application of NSM FRP strips appears to

be a particularly viable retrofitting technique (Dizhur et al. [22]),

limited research has previously been conducted on the application

of this retrofitting technique to masonry structures. For example,

of the 223 FRP-to-masonry pull tests available in the open litera-

ture, NSM tests account for only about 25% of the total. Hence, to

add to the relatively small existing NSM database, 14 new pull

tests on clay brick masonry prisms retrofitted with NSM carbon

FRP strips were conducted. These experimental tests are reported

below.

2. New pull test data

As part of ongoing research at the University of Adelaide 14

NSM carbon FRP-to-masonry pull tests were conducted. Different

variables that affect the bond behaviour of FRP-to-masonry joints

were considered such as FRP strip cross-section and loading type.

2.1. Material properties

Clay brick unitswith nominal dimensions of 230 � 110 � 76 mm

were used. The mortar used to construct the pull-test specimens

consisted of Portland cement, hydrated lime and sand, combined

in a ratio of 1:1:6 by volume.Material testswere conducted in accor-

dance with Standards Australia [23,24] to determine the quality of

the masonry and to quantify any material parameters that are re-

quired for development of the empirical model. The mean value of

the flexural tensile strength of the brick unit, fut and the modulus

of elasticity of themasonry, Em, were 3.41 MPa and10.7 GPa, respec-

tively. The ultimate tensile strength of the FRP, frupt, and Young’s

modulus of the FRP, Ep, provided by the manufacturer were

2700 MPa and 165 GPa, respectively.

2.2. Test setup and specimen details

Fig. 3 shows the test setup and instrumentation details. Each

pull test prism consisted of a five brick stack, with 10 mm mortar

bed joints and an FRP-to-masonry bonded length of 420 mm. The

groove for the NSM strip was cut using a diamond blade saw and

then filled with an epoxy adhesive. The strip was cleaned with ace-

tone to remove any foreign substances before being inserted into

to the epoxy-filled groove and allowed to cure for 7 days. The

FRP strip was positioned flush with masonry surface as shown in

Fig. 3 for all specimens. A layer of quick drying paste was applied

to the top and bottom surfaces of the masonry prism to ensure that

the load was transferred evenly using a manually controlled

hydraulic ram. A solid steel plate with a small gap for the FRP to

pass through was placed onto the top surface (loaded end) of the

specimen. This restraining plate was used to apply approximately

1kN of pre-compression to settle the specimen at the early stages

of loading. Once the specimen was in place, a monotonic tensile

load was applied in displacement control using an Avery testing

machine at a constant rate of approximately 2.5 kN per minute un-

til failure, with 0.5 mm per minute as an upper bound for displace-

ment rate. For cyclic loading, one cycle of load consisted of

increasing the load monotonically until the target load (displace-

ment) was reached, then reducing the load to zero. Thus, the

NSM strips of FRP were only loaded in tension. Each cyclic speci-

men was subjected to three cycles of load for each target

displacement.

Each test specimen had two strain gauges glued on the FRP

strips and positioned 25 mm away from the top brick unit at the

Fig. 1. Pull test simulating IC debonding.

Fig. 2. Debonding failure in pull test specimen.

J. Kashyap et al. / Engineering Structures 41 (2012) 186–198 187

loaded end to help align the FRP strip during testing and to confirm

the manufacturer’s value for the modulus of elasticity of the FRP.

Ten of the specimens were strain gauged along the bonded length

of the strip. These strain gauges were located 20 mm from the top

and bottom edges of each brick unit to measure the strain profile

along the FRP strip, as shown in Fig. 3. The FRP strips were made

up of multiple layers of 1.2 mm or 1.4 mm thick carbon FRP strips

with the strain gauges sandwiched between the strips to minimise

the negative effect that the presence of the strain gauges had on

the bond strength [5]. Two linear variable differential transformers

(LVDT) placed at the top of the masonry prism measured the slip

(D) of the FRP strip at the loaded end.

The specimens were labelled using the notation shown in Ta-

ble 1. The first term refers to the loading type (monotonic or cyc-

lic); the second term denotes whether strain gauges were

attached along the bonded length (NSG stands for no strain gauges

and SG stands for strain gauges); the third and the fourth terms re-

fer to the width, tp and depth, bp of the FRP strip, respectively and

the last term signifies the test number for cases where more than

one specimen was tested with the same FRP strip configuration

and loading.

2.3. Pull test results

2.3.1. Failure mode

The pull test results are summarised in Table 1, where tp and bprefer to the width and depth of the FRP strip, respectively, and Pexpis the maximum experimental load. All specimens failed by IC deb-

onding, which is characterised by the formation of wedges (‘her-

ringbone cracks’) at the face of the prism along the bonded

length (Fig. 2b). The propagation of diagonal cracks in the masonry

commenced from the loaded end and propagated to the unloaded

end, indicating a gradual redistribution of the shear bond stress

along the strip at the FRP-to-masonry interface. From the results

it can be noted that the development of the crack pattern and deb-

onding failure plane were similar to that of FRP-to-concrete pull

tests (Fig. 2a), confirming the similarity between the load transfer

mechanism of FRP-to-masonry joints and FRP bonded to concrete.

The load–slip response at the loaded end is shown in Fig. 4 for

the four monotonically loaded specimens that contained a

3.6 � 10 mm NSM strip. For all four specimens, progressive

Fig. 3. Pull test specimen details.

Table 1

Summary of pull test results.

Specimen No tp (mm) bp (mm) Pexp (kN) Group averages

M-SG-3.6-10-1 3.6 10.0 64.8 65.0

M-SG-3.6-10-2 3.6 10.0 70.0

M-SG-3.6-10-3 3.6 10.0 61.0

M-SG-3.6-10-4 3.6 10.0 64.0

C-SG-3.6-10-5 3.6 10.0 58.8 58.9

C-SG-3.6-10-6 3.6 10.0 59.0

M-NSG-4.2-10-1 4.2 10.0 59.0 56.2

M-NSG-4.2-10-2 4.2 10.0 53.3

M-NSG-7.2-10-1 7.2 10.0 59.0 67.2

M-NSG-7.2-10-2 7.2 10.0 75.3

M-NSG-4.8-7.5-1 4.8 7.5 60.5 59.4

M-NSG-4.8-7.5-2 4.8 7.5 58.3

M-NSG-4.8-5-1 4.8 5.0 41.0 44.8

M-NSG-4.8-5-2 4.8 5.0 48.5

Slip at loaded end, ∆ (mm)

0

20

40

60

80

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5

Lo

ad

, P

(k

N)

M-SG-3.6-10-2

M-SG-3.6-10-1 M-SG-3.6-10-4

M-SG-3.6-10-3

P

Loaded

end∆

Fig. 4. Load–slip response for monotonic specimens with 3.6 mm wide FRP strip.

188 J. Kashyap et al. / Engineering Structures 41 (2012) 186–198

damage in the FRP-to-masonry joints was indicated through re-

duced stiffness of the load–slip response with increasing load. Prior

to failure, significant diagonal cracking had occurred in the ma-

sonry (Fig. 2b) (step-wise through each brick unit, interrupted by

the mortar bed joints), demonstrating significant stress transfer

into the masonry substrate and ‘good’ bond at the FRP-to-masonry

interface. The ultimate failure plane was mostly within the ma-

sonry, near the masonry-to-adhesive interface as evidenced by a

very thin layer of masonry being detached from the masonry prism

(Fig. 5b). This observation suggests that the bonding characteristics

of FRP-to-masonry are closely related to the tensile strength of the

masonry unit and highlights the importance of fut in theoretical

models. From Fig. 4 it is evident that there was reasonable consis-

tency between the overall behaviour of these four pull test

specimens.

2.3.2. Effect of variables on test results

2.3.2.1. Effect of cyclic loading. The global load–slip (P–D) response

of two specimens subjected to cyclic loading (Fig. 6) closely traces

the average P–D curve for the four monotonically loaded speci-

mens shown previously in Fig. 4. This observation suggests that

cyclic loading does not significantly affect the overall behaviour

of the FRP-to-masonry bond. The small reduction in strength that

was observed after each loading–unloading cycle indicates that

strength degradation of the bond was cumulative. Similar behav-

iour was observed in the study by Konthesingha et al. [25]. How-

ever, it was noted that these relatively small differences in

strength and deformation capacity for the pull test specimens

may simply be due to material variability (30% is typical for ma-

sonry tests, e.g. Drysdale et al. [26]). Therefore, further testing is

recommended to confirm the effects of cyclic loading on FRP-

to-masonry bond.

2.3.2.2. Effect of FRP strip dimensions. As seen in Table 1 and Fig. 7, no

significant improvement in load–slip behaviourwas observedwhen

the width of the FRP strip, tp (measured parallel to the

masonry face), was varied from 3.6 mm to 7.2 mm. In contrast,

increasing the depth of the FRP strip, bp (measured into the ma-

sonry), was found to be effective in enhancing the debonding load

(Pexp) (Fig. 8). Considering the average value of Pexp (Table 1), increas-

ing the depth from5 mm to 7.5 mm resulted in about a 33% increase

in the maximum load (Fig. 8). Most of this strength increase can be

attributed simply to the 27% increase in the perimeter of the deb-

onding failure plane, Lper (Fig. 5b), for specimenswith a 7.5 mmdeep

strip (Lper = (4.8 + 2) mm + 2 � (7.5 + 1) mm = 23.8 mm) compared

to specimens with a 5 mm deep strip (Lper = (4.8 + 2) mm + 2

Fig. 5. Failure plane in masonry.

0

20

40

60

80

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5

Lo

ad

, P

(k

N)


C-SG-3.6-10-5

C-SG-3.6-10-6

M-SG-3.6-10-Avg

Loaded

end

P

∆

Fig. 6. Effect of cyclic loading on load–slip response.


� (5 + 1) mm = 18.8 mm). The remainder of the strength increase is

probably due to more efficient confinement of the surrounding

masonry for the deeper cut needed for the 7.5 mm strip.

2.3.3. Interface behaviour

A typical example of the axial strain distribution along an FRP

strip with increasing load is shown in Fig. 9. As seen in Fig. 9, for

specimen M-SG-3.6-10-1 the strain at the loaded end increased

with increasing levels of applied load and eventually began to pla-

teau out to a value of about 10,500 le at the maximum load Pexp of

64.8 kN. The strain plateau represents debonding at that point of

the FRP strip, as there is no further significant change in strain,

indicating little transfer of shear across the FRP-to-masonry

interface. The position along the bonded length where the strain

has reduced to near zero corresponds to the location at which

the entire applied load has been transferred from the FRP to the

masonry through interface bond. This behaviour is evident in

Fig. 9 for loads up to 60 kN. At a load of 64.8 kN, it is evident that

the strain is no longer zero at the last gauge and therefore, the force

had not been fully transferred to the masonry over the bonded

length between this point and the loaded end, i.e. the debonding

crack had propagated along the entire length of the FRP strip.

The strain gauge data was used to estimate the shear stress

transferred from the FRP to the masonry, for all the specimens with

strain gauges, using Eq. (1)

savg ¼ðDeÞEpbptp

ðDLÞð2bp þ tpÞð1Þ

where savg is the average shear stress over the length DL; De is the

change in strain over lengthDL; Ep is the modulus of elasticity of the

FRP strip; bp is the depth of strip; tp = width of strip; and DL is the

incremental length along the FRP (equal to strain gauge spacing).

Fig. 10 shows the typical shear stress distribution along the

bonded length for increasing load. The shear stress at the loaded

end initially increases with applied load until the maximum shear

stress is reached. At this point there is a reduction in the magni-

tude of the shear stress as micro-cracking develops at the loaded

end and begins to propagate along the FRP strip towards the un-

loaded end (Fig. 10). With further loading, the maximum shear

stress location shifts away from the loaded end, to the right in this

case, as observed in FRP-to-concrete pull tests [8]. The shear stress

distribution also provides an experimental measure of the effective

bond length, Leff, (Fig. 10) which is the length of bonded FRP re-

quired to develop the maximum IC debonding load (PIC). The effec-

tive bond length ranged between 250 mm to 300 mm for all the

specimens in this study.

Fig. 11 shows the local bond–slip response at various strain

gage locations (measured from the loaded end). The local bond–

slip response (Fig. 11) varied along the bonded length for each

specimen, which is likely attributed to: (1) the varying spacing be-

tween strain gauges, (2) the location of cracking relative to the

strain gauges, and/or (3) the variation in the tensile strength of

the bricks (fut). As shown in Fig. 11, the load transfer mechanism

at the FRP-to-masonry interface is similar to that of FRP-to-con-

crete [8] and can be represented in three stages, i.e. elastic, mi-

cro-cracking (softening) and macro-cracking (debonding). In the

elastic stage, the shear-stress in the FRP-to-masonry interface in-

creases until the peak shear stress, smax, is reached at local slip

d1. This elastic stage is followed by a softening region as micro-

cracking develops and the shear stress decreases to zero at a slip


0

20

40

60

80

0.0 0.5 1.0 1.5 2.0 2.5 3.0

Lo

ad

, P

(k

N)

M-NSG-7.2-10-2M-SG-3.6-10-Avg

M-NSG-4.2-10-1

M-NSG-4.2-10-2

M-NSG-7.2-10-1

P

Loaded

end∆

Fig. 7. Effect of width of FRP strip on load–slip response.

0

20

40

60

80

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Lo

ad

, P

(k

N)


M-NSG-4.8-7.5-1

M-NSG-4.8-7.5-2

M-NSG-4.8-5-1

M-NSG-4.8-5-2

∆Loaded end

P

Fig. 8. Effect of depth of FRP strip on load–slip response.

Fig. 9. Typical FRP strain distribution (Specimen M-SG-3.6-10-1).

Loaded

end

P

τ

0

5

10

15

20

0 50 100 150 200 250 300 350 400

Sh

ear

Str

ess,

τ (

MP

a)

Distance from loaded end (mm)

Leff

Fig. 10. Typical interface shear-stress distribution (Specimen M-SG-3.6-10-1).


of dmax beyond which macro-cracking occurs and the debonding

interface is no longer capable of transferring force. An idealised

bilinear model was fitted to the experimental data as shown by

the dashed line in Fig. 11.

3. Existing pull test data

A database was developed containing the results of 223 FRP-to-

masonry pull tests reported in the open literature for both EB and

NSM FRP techniques. It should be noted that other tests reported in

the literature (e.g. [27–29]) had insufficient information (e.g. mate-

rial and/or geometric properties) provided to allow meaningful

analysis and comparison. The database includes results from ten

different experimental studies [30–33,19,6,25,34,5,35]. Further, re-

sults with failure modes other than IC debonding and bond lengths

less than Leff were excluded where Leff was calculated using Eq. (4).

Consequently, a dataset of 109 pull test results was available as

shown in Table 2 (where Pexp is the experimental value of the IC

debonding resistance). The range of experimental parameters in-

cluded in the database (are: 22.3 GPa < Ep < 230 GPa; 1.3 MPa < -

fut < 3.57 MPa and 0.02 < uf < 6.56, where uf = df/bf is the IC

debonding failure plane aspect ratio and df and bf are the lengths

of the failure plane perpendicular and parallel to the surface of

the substrate as shown in Fig. 5. For some of the results, the flex-

ural tensile strength of the brick unit, fut, was not reported and

so, it was calculated assuming the following relation for concrete

between the tensile and compressive ðf 0cÞ strength is also applicable

to masonry units [36].

ffiffiffiffi

f 0c

q

¼fut

0:53ð2Þ

4. Existing bond strength models

Many theoretical models have been previously developed to

predict the bond strength of FRP-to-concrete joints where the bond

strength refers to the shear strength capacity of the FRP-to-sub-

strate interface. A review of the existing models applicable to EB

FRP-to-concrete bonded joints can be found in Lu et al. [37], Kharb-

hari et al. [38] and Sayed-Ahmed et al. [39]. A total of 24 FRP-to-

concrete bond strength models from the literature were investi-

gated. These models are based on either empirical relations

calibrated with experimental data or on fracture mechanics based

theories or combinations of the two. Some of these models involve

parameters that are not applicable to masonry while for others,

insufficient information was available to include the models in

the reported analysis. Further, for the models proposed by some

researchers [40–43] adhesive properties are needed. However,

the properties of adhesives used in the specimens of the present

database were not always reported and hence, these models are

not further discussed in this paper. Consequently, only 12 EB mod-

els were considered in this study (Tanaka (from Sayed-Ahmed et al.

[39]), Hiroyuki and Wu [44], Maeda et al. (from Chen and Teng

[45]), Neubauer and Rostasy (from Sayed-Ahmed et al. [39]), Kha-

lifa et al. [46], Chen and Teng [45], Yang et al. (from Sayed-Ahmed

et al. [39]), Monti et al. [47], Iso (from Sayed-Ahmed et al. [39]),

Sato (from Sayed-Ahmed et al. [39]), Lu et al. [37] and Kharbhari

et al. [38] (extended from Nakaba et al. [48])). Analytical models

specific to the NSM technique have been published by Blaschko

[49] and Seracino et al. [8]. As Blaschko’s model requires adhesives

property data, it also was not included in this analysis.

All the above models are specific to a particular retrofitting

technique, i.e. either EB or NSM. The first generic analytical model

applicable to any adhesively bonded plate cross-section and mate-

rial was developed by Seracino et al. [50] for predicting the IC

debonding resistance, PIC, using Eq. (3):

PIC ¼ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi

smaxdmax

p

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi

LperðEAÞp

q

ð3Þ

where the maximum interface shear stress,

smax ¼ ð0:802þ 0:078uf Þf0:6c ; fc is the cylinder compressive strength

of the concrete; the maximum local slip, dmax ¼0:976u0:526

f

0:802þ0:078uf; uf and

Lper are as defined previously and (EA)p is the axial rigidity of the

FRP strip.

The effective bond length, Leff, is given by Eq. (4)

Leff ¼p2k

ð4Þ

where k ¼ffiffiffiffiffiffiffiffiffiffiffiffiffiffi

smaxLperdmaxðEAÞp

q

.

From the current literature, only three theoretical models

[33,5,51] are available for masonry to predict the debonding resis-

tance of FRP-to-masonry joints. Moreover, the models proposed by

Camli and Binici [32] and Milani et al. [51] are specific to the EB

retrofitting technique. The generic model by Willis et al. [5] was

adapted from the generic model of Seracino et al. [50] described

earlier in this section using Eq. (2) and can be applied to both EB

and NSM techniques.

5. Comparison of existing bond strength models

Table 3 presents results that were used to evaluate the reliabil-

ity and accuracy of 15 existing models to predict the bond strength

of the FRP-to-masonry pull tests given in the present test database

(Tables 1 and 2), based on the test-to-predicted bond strength ratio

(Pexp/PIC). As can be observed in Table 3a, the generic model by Wil-

lis et al. [5] under-predicts the bond strength on average by about

1/3rd when both EB and NSM specimens are considered. Table 3b

indicates that seven of the fifteen existing models applicable spe-

cifically for EB strips (Tanaka (from Sayed-Ahmed et al. [39]), Hiro-

yuki and Wu [44], Neubauer and Rostasy (from Sayed-Ahmed et al.

[39]), Khalifa et al. [46], Iso (from Sayed-Ahmed et al. [39]), Sato

(from Sayed-Ahmed et al. [39]), Monti et al. [47] and Camli and

Binici [33]) do not agree well with the experimental data, having

either a mean value greater than or equal to 2 or a COV greater

than 35%. It is worth noting that some of these models have also

been identified by Chen and Teng [45] and Lu et al. [37] as being

poorly performing models for FRP-to-concrete joints. Further,

although the models by Chen and Teng [45], Willis et al. [5] and

Yang et al. (from Sayed-Ahmed et al. [39]), have a COV within

the typical limit for masonry (i.e. 30%) and a relatively good corre-

lation coefficient, these models significantly underestimate the

bond strength. Based on the data analysed in this study, only four

0

5

10

15

20

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2

Sh

ear

stre

ss, τ

(MP

a)

Local slip, δ (mm)

δ1 δmax

τmax

Bilinear model

τ

δ

Loaded

end

P

τ

Softening

Elastic

Debonding

Fig. 11. Local bond–slip relationship (Specimen M-SG-3.6-10-1).


Table 2

FRP-to-masonry pull tests database.

References Specimen EB/NSM tp (mm) bp (mm) Ep (GPa) Lb (mm) fut (MPa) bm (mm) Pexp (kN)

Turco et al. [31] E-10-SS NSM 6.35 6.35 40.8 254 1.93 230 19.17

E-15-SS NSM 6.35 6.35 40.8 381 1.93 230 18.55

Liu et al. [32] CM50(A) EB 1.00 25.00 22.3 50 2.05 400 4.88

CM50(B) EB 1.00 25.00 22.3 50 2.05 400 5.63

CM50(C) EB 1.00 25.00 22.3 50 2.05 400 4.25

CM50(D) EB 1.00 25.00 22.3 50 2.05 400 3.75

CM50(E) EB 1.00 25.00 22.3 50 2.05 400 5.13

CM75(A) EB 1.00 25.00 22.3 75 2.05 400 5.81

CM75(B) EB 1.00 25.00 22.3 75 2.05 400 5.44

CM75(C) EB 1.00 25.00 22.3 75 2.05 400 6.38

CM75(D) EB 1.00 25.00 22.3 75 2.05 400 3.94

CM75(E) EB 1.00 25.00 22.3 75 2.05 400 7.13

CM100(A) EB 1.00 25.00 22.3 100 2.05 400 4.75

CM100(B) EB 1.00 25.00 22.3 100 2.05 400 5.00

CM100(C) EB 1.00 25.00 22.3 100 2.05 400 6.50

CM100(D) EB 1.00 25.00 22.3 100 2.05 400 7.25

CM100(E) EB 1.00 25.00 22.3 100 2.05 400 7.25

CM100(F) EB 1.00 25.00 22.3 100 2.05 400 8.50

BM50(A) EB 1.00 25.00 22.3 50 2.73 200 9.25

BM50(B) EB 1.00 25.00 22.3 50 2.73 200 7.38

BM50(C) EB 1.00 25.00 22.3 50 2.73 200 8.63

BM50(D) EB 1.00 25.00 22.3 50 2.73 200 6.88

BM75(A) EB 1.00 25.00 22.3 75 2.73 200 10.69

BM75(B) EB 1.00 25.00 22.3 75 2.73 200 8.44

BM75(C) EB 1.00 25.00 22.3 75 2.73 200 9.38

BM75(D) EB 1.00 25.00 22.3 75 2.73 200 9.56

BM75(E) EB 1.00 25.00 22.3 75 2.73 200 8.25

BM100(A) EB 1.00 25.00 22.3 100 2.73 200 8.50

BM100(B) EB 1.00 25.00 22.3 100 2.73 200 10.00

BM100(C) EB 1.00 25.00 22.3 100 2.73 200 10.00

BM100(D) EB 1.00 25.00 22.3 100 2.73 200 9.00

BM100(E) EB 1.00 25.00 22.3 100 2.73 200 10.00

CSG100 EB 1.00 25.00 22.3 100 2.05 400 5.58

BSG100 EB 1.00 25.00 22.3 100 2.73 200 9.40

Camli and Binici [33] TS-4 EB 1.00 25.00 61.0 125 1.30 280 4.06

TS-7 EB 1.00 50.00 61.0 100 1.30 280 5.90

TS-8 EB 1.00 50.00 61.0 125 1.30 280 5.14

Xia and Oehlers [19] Pull 3 EB 1.20 50.00 165 210 2.75 230 25.25

Pull 4 EB 1.20 50.00 165 280 2.75 230 28.40

Konthesingha et al. [25] 1A NSM 2.80 15.00 207 355 3.57 230 61.60

2A NSM 2.80 15.00 207 355 3.57 230 65.24

2B NSM 2.80 15.00 207 355 3.57 230 63.53

2C NSM 2.80 15.00 207 355 3.57 230 66.52

Lam [34] C1 EB 0.17 50.00 230 200 3.35 250 15.94

C2 EB 0.17 50.00 230 200 3.35 250 17.12

C3 EB 0.17 50.00 230 200 3.35 250 17.66

C4 EB 0.17 50.00 230 200 3.35 250 19.61

C5 EB 0.17 50.00 230 200 3.35 250 20.15

G1 EB 0.23 50.00 65.0 200 3.35 250 11.69

G2 EB 0.23 50.00 65.0 200 3.35 250 13.97

G3 EB 0.23 50.00 65.0 200 3.35 250 13.65

G4 EB 0.23 50.00 65.0 200 3.35 250 13.20

G5 EB 0.23 50.00 65.0 200 3.35 250 14.18

Petersen et al. [6] S1-A-NG-1 NSM 2.80 15.00 207 336 3.57 230 83.45

S1-A-NG-2 NSM 2.80 15.00 207 336 3.57 230 71.09

S1-A-SG NSM 2.80 15.00 207 336 3.57 230 81.48

S1-B-NG-1 NSM 2.80 15.00 207 336 3.57 230 70.36

S1-B-NG-2 NSM 2.80 15.00 207 336 3.57 230 59.41

S1-C-NG-1 NSM 2.80 15.00 207 336 3.57 230 63.88

S1-C-NG-2 NSM 2.80 15.00 207 336 3.57 230 69.41

S1-C-SG NSM 2.80 15.00 207 336 3.57 230 84.50

Willis et al. [5] St1.0-3-15-1/2-NSG NSM 1.20 15.00 162 241 3.55 230 46.80

HO1.5-4-15-1/4 NSM 1.20 15.00 162 328 3.55 230 44.00

HO1.0-4-15-0 NSM 1.20 15.00 162 328 3.55 230 38.30

HO1.0-4-15-0-NSG NSM 1.20 15.00 162 334 3.55 230 46.70

St1.0-4-20-AC NSM 1.20 20.00 162 328 3.55 230 50.00

St1.0-4-20-BC NSM 1.20 20.00 162 328 3.55 230 51.20

GP-5-Ne-M-1 EB 2.00 50.00 65.0 420 3.55 230 22.10

GP-5-Ne-Q-1 EB 2.00 50.00 65.0 395 3.55 230 21.50

GP-5-Ne-M-2 EB 2.00 50.00 65.0 419 3.55 230 21.90

GP-5-Ne-Q-2 EB 2.00 50.00 65.0 396 3.55 230 18.10


models (Maeda et al. (from Chen and Teng [45]), Milani et al. [51],

Lu et al. [37] and Kharbhari et al.[38]) out of the 15 models appli-

cable for EB can be considered to provide satisfactory predictions

of the debonding strength (Table 3b). From Table 3b it can be ob-

served that the models by Lu et al. [37] and Kharbhari et al. [38]

provide particularly good estimates of the bond strength when

compared to the other models. It is interesting to note that the gen-

eric model by Willis et al. [5] correlates better with the NSM test

data (Table 3c) than with the EB test data (Table 3b). Further, the

NSM specific bond strength model by Seracino et al. [8] is also

not reliable, with a large coefficient of variation (COV) and low cor-

relation coefficient as shown in Table 3c. From these results it can

Table 2 (continued)

References Specimen EB/NSM tp (mm) bp (mm) Ep (GPa) Lb (mm) fut (MPa) bm (mm) Pexp (kN)

GP-5-Sa-Q EB 2.00 50.00 65.0 394 3.55 230 24.70

GP-5-Gr-Q EB 2.00 50.00 65.0 393 3.55 230 24.30

GD-5-Sa-Q-1 EB 0.62 50.00 73.0 386 3.55 230 19.90

GD-5-Sa-Q-2 EB 0.62 50.00 73.0 386 3.55 230 18.60

CP-2-Ne-M EB 1.20 50.00 162 140 3.55 230 26.80

CP-3-Ne-M EB 1.20 50.00 162 210 3.55 230 24.90

CP-4-Ne-M EB 1.20 50.00 162 280 3.55 230 28.40

Oliveria et al. [35] G150RS-1 EB 0.15 25.00 80.2 150 1.57 235 3.48

G150RS-2 EB 0.15 25.00 80.2 150 1.57 235 4.81

G150RS-3 EB 0.15 25.00 80.2 150 1.57 235 4.69

G150RS-5 EB 0.15 25.00 80.2 150 1.57 235 4.64

G100RS-1 EB 0.15 25.00 80.2 100 1.57 235 3.66

G100RS-2 EB 0.15 25.00 80.2 100 1.57 235 3.17

G100RS-3 EB 0.15 25.00 80.2 100 1.57 235 2.85

G100RS-4 EB 0.15 25.00 80.2 100 1.57 235 3.68

G100RS-5 EB 0.15 25.00 80.2 100 1.57 235 3.79

G200RS-2 EB 0.15 25.00 80.2 200 1.57 235 4.48

G200RS-3 EB 0.15 25.00 80.2 200 1.57 235 5.06

G150RI-5 EB 0.15 25.00 80.2 150 1.57 235 5.27

G150RT-1 EB 0.15 25.00 80.2 150 1.57 235 4.20

G150RT-2 EB 0.15 25.00 80.2 150 1.57 235 4.89

G150RT-3 EB 0.15 25.00 80.2 150 1.57 235 5.60

G150RT-4 EB 0.15 25.00 80.2 150 1.57 235 4.34

G150RT-5 EB 0.15 25.00 80.2 150 1.57 235 5.49

G150ES-1 EB 0.15 25.00 80.2 150 1.57 235 3.52

G150ES-2 EB 0.15 25.00 80.2 150 1.57 235 4.83

G150ES-4 EB 0.15 25.00 80.2 150 1.57 235 4.53

G150ES-5 EB 0.15 25.00 80.2 150 1.57 235 5.46

G150XS-1 EB 0.15 25.00 80.2 150 1.57 235 4.55

G150XS-2 EB 0.15 25.00 80.2 150 1.57 235 3.73

G150XS-3 EB 0.15 25.00 80.2 150 1.57 235 3.82

G150XS-4 EB 0.15 25.00 80.2 150 1.57 235 4.54

G150XS-5 EB 0.15 25.00 80.2 150 1.57 235 4.06

C150RS-1 EB 0.12 25.00 216 150 1.57 235 4.78

C150RS-2 EB 0.12 25.00 216 150 1.57 235 4.29

C150RS-3 EB 0.12 25.00 216 150 1.57 235 4.02

C150RS-4 EB 0.12 25.00 216 150 1.57 235 4.33

C150RS-5 EB 0.12 25.00 216 150 1.57 235 4.26

Table 3

Pexp/PIC ratios for existing models.

Model Mean Median Max Min SD COV Correlation

(a) For all specimens

Willis et al. [5] 1.507 1.514 2.388 0.668 0.458 0.304 0.919

(b) For EB specimens only

Kharbhari et al. [38] 1.111 1.125 1.848 0.309 0.327 0.294 0.850

Lu et al. [37] 1.204 1.049 1.854 0.792 0.294 0.244 0.890

Milani et al. [51] 1.288 1.281 1.989 0.434 0.367 0.285 0.885

Maeda et al. (from [45]) 1.332 1.260 2.283 0.360 0.398 0.299 0.831

Chen and Teng [45] 1.486 1.508 2.289 0.529 0.408 0.274 0.886

Willis et al. [5] 1.642 1.646 2.388 0.668 0.426 0.259 0.884

Yang et al. (from [39]) 1.702 1.612 2.562 0.709 0.406 0.239 0.924

Camli and Binici [33] 1.844 1.585 3.813 0.496 0.772 0.419 0.837

Iso (from [39]) 2.046 2.008 3.508 0.710 0.724 0.354 0.879

Monti et al. [47] 2.099 1.843 4.225 0.878 0.899 0.428 0.926

Nebauer and Rostasy (from [39]) 2.256 1.955 4.434 0.907 0.953 0.422 0.922

Sato (from [39]) 2.341 2.681 4.961 0.180 1.280 0.547 0.710

Khalifa et al. [46] 2.459 2.468 4.405 0.882 0.890 0.362 0.915

Hiroyuki and Wu [44] 9.174 8.350 17.759 3.536 3.731 0.407 0.735

Tanaka (from [39]) 2.294 1.680 11.726 0.632 2.071 0.903 0.174

(c) For NSM specimens only

Willis et al. [5] 1.148 1.074 1.919 0.742 0.333 0.290 0.565

Seracino et al. [8] 1.405 1.329 3.198 0.643 0.574 0.409 0.260


be seen that the existing bond strength models are inadequate for

predicting the FRP-to-masonry joint test data considered in this

study. Additionally, numerical investigations carried out by Milani

[52] and Milani and Bucchi [53] on reinforced masonry walls and

curved structures, respectively, suggests that application of exist-

ing FRP-to-concrete models to masonry is questionable. This

further highlights the need for development of new models that

can more accurately predict the bond strength of FRP retrofitted

masonry and its application to complex realistic geometries.

6. New model for FRP-to-masonry joints

The generic model of Seracino et al. [50] for defining PIC was

considered as the starting point for development of a new model.

Seracino et al.’s [50] model for concrete is a function of the cylinder

compressive strength, fc, which was originally incorporated to rep-

resent the principal tensile stress that resulted in the interface deb-

onding cracks. The corresponding material property for masonry is

the flexural tensile strength of the masonry unit, fut which is a va-

lue used widely in design and so is readily available for use in an

expression of PIC.

It is known that the debonding strain is higher for a NSM strip

than for an EB strip due to better confinement. Therefore, to ac-

count for the type of retrofitting technique (i.e. EB or NSM) it

was assumed that smaxdmax, which is twice the fracture energy, Gf

(area under the local bond–slip curve), is a function of the geomet-

ric property,uf. Also, since the debonding cracks are initiated with-

in the masonry, smaxdmax was assumed to be a function of the

masonry material property, fut [50]. Hence, the following expres-

sion can be written:

smaxdmax ¼ Cumf f

nut ð5Þ

6.1. New generic model

In order to derive values for C, m and n in Eq. (5), first a value of

smaxdmax was calculated for each test specimen in the pull-test

database by substituting Pexp into Eq. (3) and solving for smaxdmax.

A statistical analysis was then undertaken to determine optimal

values for the constants C, m and n using the data for 109 speci-

mens from the database (Table 2) and the 14 new pull test results

(Table 1). The parameters C and m were determined from linear

regression, with n varied incrementally until the lowest coefficient

of variation was identified. Eq. (6) was obtained from this process

and is represented graphically in Fig. 12. Data points clustered on

the left end of the graph represent results for EB test specimens

and those on the right end represent NSM test results.

ðsmaxdmaxÞGeneric ¼ 3:94u0:38f f 0:94ut ð6Þ

By substituting Eq. (6) into Eq. (3), a generic expression was de-

rived for the IC debonding resistance, PIC, given by Eq. (7) in terms

of the conventional design parameters, fut, for masonry and the

other standard parameters for an FRP section.

ðPICÞGeneric ¼ 1:99u0:19f f 0:47ut


LperðEAÞp

q

ð7Þ

The predictions of the new generic model (Eq. (7)) compared

with the 123 test results of the present database (Tables 1 and 2)

are shown in Fig. 13 and summarised in Table 4a. A value greater

than one for the test-to predicted bond strength ratio (Pexp/PIC)

(i.e. above the line in Fig. 13) signifies a conservative estimate of

bond strength. As shown in Table 4a and Fig. 14a, the experimental

and the predicted bond strengths agreed reasonably well with

mean, COV, and linear regression correlation coefficient of the

test-to-predicted strength ratios being 1.036, 25.6% and 0.922,

respectively.

Although Eq. (7) is valid for both EB and NSM pull tests, it was

considered important to further investigate the EB and NSM spec-

imens separately to identify the sensitivity of the new generic PICexpression (Eq. (7)) to consideration of the database in subsets,

and whether further improvements of the new generic model

could be achieved.

6.2. New EB specific model

When considering a model specific to EB FRP strips, 89 tests re-

sults were available out of the total of 109 tests listed in Table 2.

This subset of the data was used to determine smaxdmax giving Eq.

(9) for calculating the bond strength of EB FRP strips in masonry.

The same method used for generic equation was applied to cal-

ibrate the constants C,m and n (Eq. (5)) for 89 EB test results giving

Eqs. (8) and (9).

ðsmaxdmaxÞEB ¼ 187:40u1:69f f 1:80ut ð8Þ

ðPICÞEB ¼ 13:69u0:84f f 0:90ut


LperðEAÞp

q

ð9Þ

The statistics of the Pexp/PIC ratios for the EB specific model is gi-

ven in Table 4b. The correlation coefficient for the EB model is

0.884 and the mean value is close to one (Table 4b and Fig. 15a)

which demonstrates that the ‘EB specific’ bond strength model

-1

0

1

2

-2 -1 0 1

log( )

log(τmaxδmax/fut0.94)

Fig. 12. Statistical analysis for key bond–slip parameters.

-

20

40

60

80

100

- 20 40 60 80 100

Pex

p(k

N)

PIC (kN)

Fig. 13. Comparison of experimental and predicted (using new generic model) IC

debonding resistance.


Table 4

Bond strength ratios Pexp/PIC for new models.

Model Mean Median Max Min SD COV Correlation

(a) All specimens

New generic (Eq. (7)) 1.036 1.027 1.629 0.379 0.265 0.256 0.922

(b) EB specimens only

New EB specific (Eq. (9)) 1.030 1.014 1.544 0.570 0.248 0.2406 0.884

New generic (Eq. (7)) 1.031 1.052 1.559 0.377 0.273 0.265 0.876

(c) NSM specimens only

New NSM specific (Eq. (11)) 1.009 1.006 1.266 0.668 0.135 0.133 0.727

New generic (Eq. (7)) 1.023 0.983 1.625 0.730 0.234 0.229 0.613

(a) New generic model (b) Willis et al. [5]

0

10

20

30

40

No

. o

f sp

ecim

ens

Ratio of Pexp vs PIC

Mean = 1.036

COV = 0.256

0

10

20

30

40

0.0 0.6 1.2 1.8 2.4 3.0 3.6 4.2 0 0.6 1.2 1.8 2.4 3 3.6 4.2

No

. o

f sp

ecim

ens

Ratio of Pexp vs PIC

Mean = 1.507

COV =0.303

Fig. 14. Effectiveness of models for bond strength prediction for all specimens.

Fig. 15. Effectiveness of models for bond strength prediction for EB specimens.


predicts the trends of the test data reasonably well. Further, it was

noted that the values of the constants C,m and n in the new generic

expression (Eq. (7)) and the EB expression (Eq. (9)) are significantly

different in spite of the commonly assumed form for the relation-

ship between fracture energy and aspect ratio and masonry unit

tensile strength. Moreover, when the new generic model (Eq. (7))

was used to predict the bond strength for only EB specimens from

the database, the correlation coefficient was only slightly lower

and the COV was only slightly higher than for the ‘EB specific’

model as shown in Table 4b. Hence, while the EB specific model

provides reasonable estimate of the bond strength, it does not

appear to be any more accurate or reliable than the new generic

model.

6.3. New NSM specific model

In a similar fashion, the 34 NSM specimens from the database

(Tables 1 and 2) were used to calibrate the constants C, m and n

in Eq. (5) in order to define a bond strength model specific to the

NSM technique. The final expression is given by Eq. (11).

ðsmaxdmaxÞNSM ¼ 6:94uÿ0:24f f 0:94ut ð10Þ

ðPICÞNSM ¼ 2:63uÿ0:12f f 0:47ut


LperðEAÞp

q

ð11Þ

From Table 4c it can be observed that the average value of the Pexp/

PIC ratio for the NSM model (Eq. (11)) is close to one and the COV is

about 0.133. Further, by comparing the generic (Eq. (7)), EB (Eq. (9))

and NSM (Eq. (11)) models it can be noted that the constants C, m

and nwere significantly different for each model, possibly reflecting

the difference between the EB and NSM applications. Moreover, the

results for the new generic model, when applied to only NSM

specimens, were only slightly less accurate than those obtained

using the ‘NSM specific’ model’ (Table 4c). This was also clear from

Fig. 16 which shows the differences between these two models

(Fig. 16a versus b). Given the versatility of the new generic expres-

sion over the NSM specific expression, the new generic expression is

preferred. In any case, given the relatively small size of the data set

available, there seems to be scope for further NSM bond tests on

masonry.

6.4. Local bond–slip model

The parameters, smax and dmax, from the pull test data are the

principal parameters for the numerical investigation of FRP

strengthened structures and are needed to calculate the critical

Fig. 16. Effectiveness of models for bond strength prediction for NSM Specimens.

0

2

4

6

8

10

12

14

0 1 2 3 4 5 6 7

τm

ax/f

ut0

.2

Fig. 17. Statistical analysis for smax.


bond length, Leff ((Eq. (4)). Hence, the statistical analysis was ex-

tended to determine an empirical expression for smax. Reported val-

ues of smax (for 20 tests) were plotted against uf and the data was

analysed using power regression. The resulting expression for smax

is given by Eq. (12) and is graphically represented in Fig. 17. There

is good agreement between the experimental and the predicted

values of smax with mean of 1.009 and a correlation coefficient of

0.897.

smax ¼ 8:83u0:15f f 0:2ut ð12Þ

Eq. (6) can then be divided by Eq. (12) to generate the following

expression for dmax:

dmax ¼ 0:45u0:23f f 0:74ut ð13Þ

Note, that smax and dmax are functions of both the masonry unit

tensile strength and the aspect ratio of the FRP strip. This conclu-

sion differs from that reported by Seracino et al. [50] for concrete

where dmax was expressed as a function of the aspect ratio only.

It should be noted that the above expressions for smax (Eq. (12))

and dmax (Eq. (13)) are based on small data set available.

7. Comparison of the proposed new bond strength models with

existing models

Fig. 14 compares the model by Willis et al. [5] with the new

generic model (Eq. (7)), considering results from both NSM and

EB specimens given in the complete test database of 123 speci-

mens (Tables 1 and 2). It is clear that the new generic model is

more accurate than the model by Willis et al. [5] suggesting that

equations developed for concrete (e.g. Eq. (3)) cannot be applied

directly to masonry by simply substituting f 0c with the correspond-

ing tensile masonry strength (fut).

For the EB specimens the new generic model gives good esti-

mates of the bond strength, with a mean value for the Pexp/PIC ratio

being close to unity and a COV value of approximately 25%

(Table 4b). From Fig. 15 it can be identified that the new generic

model (Fig. 15b) is more accurate as compared to the four best per-

forming existing models (Fig. 15c–f) mentioned previously. While

the model by Lu et al. [37] gives results that are close to the new

generic model it is suggested that the Lu et al. model [37] is too

complicated for practical application. Fig. 16 gives similar compar-

isons between the new generic model (Fig. 16b) and pre-existing

expressions (Fig. 16c and d) applicable to NSM specimens. It is

concluded that the new generic model is more accurate than the

existing models with reasonably low COV values (of about 25%).

8. Conclusions and recommendations

The results of an experimental and analytical study into the

FRP-to-masonry bond behaviour are presented. In the experimen-

tal study, 14 new NSM FRP-to-masonry pull tests were conducted

to assess the effect of different test variables such as cyclic loading

and FRP strip dimensions to establish their influence on FRP-to-

masonry joint behaviour. All specimens failed by IC debonding.

The pull test results also indicated that increasing the depth of

the NSM FRP strips (measured perpendicular to the face of the ma-

sonry) was more effective at enhancing the debonding resistance

of FRP-to-masonry joint than was increasing the width of the FRP

strip, possibly because of the increase in Lper for specimens with

a deeper strip and hence providing more efficient confinement of

the surrounding masonry. The experimental investigation into

the effects of cyclic loading on load–slip behaviour of FRP-to-ma-

sonry bonded joints was somewhat inconclusive. Therefore, it is

recommended that further research into the effects of cyclic load-

ing on FRP-to-masonry bond be conducted.

An analytical model was derived for the IC debonding resistance

of FRP-to-masonry joints using a large set of test data from the

open literature. A new generic model that is applicable to both

EB and NSM techniques (generic) was derived. This model is a

function of the usual geometric and material design variables such

as the FRP strip aspect ratio, the axial rigidity of the FRP strip and

the masonry unit tensile strength. For the range of experimental

parameters investigated as part of this study, the predicted values

from the new model correlate reasonably well with the experi-

mental values. Further, empirical expressions were developed for

the key local bond–slip parameters such as peak interface shear-

stress and maximum slip. These bond characteristics are necessary

to determine the critical bond length of FRP-to-masonry joints and

for analysing the behaviour of FRP strengthened structures. Impor-

tantly, it should be noted that the proposed models give the lower

bound debonding resistance in FRP strengthened structures due to

the presence of the multiple cracks in flexural members.

Although the model proposed in this study is reasonably accu-

rate there is further scope for improvement by adding more FRP-

to-masonry pull test results to the present database. In addition,

a significant number of test results were excluded from the test

database due to insufficient information available for analysis.

The exclusion of some data points to the need for standardised

reporting of test results as well as standardised testing procedures.

References

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Debonding resistance of FRP-to-clay brick masonry joints