experimental bond behavior of frp sheets glued on brick masonry

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Experimental Bond Behavior of FRP Sheets Gluedon Brick Masonry

Daniel V. Oliveira, Ph.D.1; Ismael Basilio, Ph.D.2; and Paulo B. Lourenço, Ph.D.3

Abstract: This paper deals with the experimental characterization of the mechanical tensile and shear bond behavior of fiber-reinforcedpolymer �FRP� sheets externally glued on masonry prisms, in terms of load capacity and stress distribution along the bonded length. Thebrick masonry adopted tries to replicate ancient brick masonry, by using handmade low-strength solids bricks and low-strength lime-basedmortar. Key parameters relative to the FRP-masonry interface response, particularly bonded length, FRP materials, anchor schemeadopted, and shape of masonry substrate, were studied. Finally, an analytical bond stress-slip formulation was developed, allowingdeducing local bond stress-slip curves directly from the experiments.

DOI: 10.1061/�ASCE�CC.1943-5614.0000147

CE Database subject headings: Bricks; Masonry; Fiber reinforced polymer; Bonding; Experimentation; Sheets.

Author keywords: Brick masonry; Fiber-reinforced polymers; Bond; Bond-slip behavior; Pull-off.

Introduction

The preservation of cultural heritage buildings is considered amajor issue in the cultural life of modern societies. However, thistask requires the establishment of proper methodologies, able totackle simultaneously their considerable architectural and culturalvalue and their structural safety �ICOMOS 2001�. An importanttopic in current research is the development of efficient and cost-effective strengthening techniques, able to reestablish the perfor-mance of cultural heritage buildings and to prevent their brittlefailure, particularly under earthquake loading.

In the last years, growing attention has been devoted to thedevelopment of innovative materials and advanced technologiesto be applied in the preservation of cultural heritage buildings.Strengthening solutions based on the external bonding of fiber-reinforced polymer �FRP� composites have become a popular op-tion �e.g., fib 2001; Bakis et al. 2002; Yuan et al. 2004; Shrive2006�. FRP presents several advantages, such as low specificweight, corrosion immunity, and high tensile strength. Its flexibil-ity and easy application contribute to a wide range of interventionscenarios. An important research effort has been directed to thecharacterization of FRP strengthened structures but the main em-phasis has been directed to concrete structures, whereas only fewworks have dealt with masonry supports, particularly ancientones. As a consequence, masonry guidelines and recommenda-tions on FRP strengthening reflect this unbalanced situation,

1Assistant Professor, ISISE, Dept. of Civil Engineering, Univ. ofMinho, Azurém, P-4800-058 Guimarães, Portugal �corresponding au-thor�. E-mail: [email protected]

2Formerly, Ph.D. Student, Dept. of Civil Engineering, Univ. of Minho,Azurém, P-4800-058 Guimarães, Portugal. E-mail: [email protected]

3Professor, ISISE, Dept. of Civil Engineering, Univ. of Minho,Azurém, P-4800-058 Guimarães, Portugal. E-mail: [email protected]

Note. This manuscript was submitted on October 24, 2009; approvedon June 4, 2010; published online on June 9, 2010. Discussion periodopen until July 1, 2011; separate discussions must be submitted for indi-vidual papers. This paper is part of the Journal of Composites for Con-struction, Vol. 15, No. 1, February 1, 2011. ©ASCE, ISSN 1090-0268/
2011/1-32–41/$25.00.
32 / JOURNAL OF COMPOSITES FOR CONSTRUCTION © ASCE / JANUAR

J. Compos. Constr. 2

being the theoretical framework usually derived from researchon concrete, as testifies the recent Italian guidelines for FRPstrengthening of masonry �National Research Council of Italy2004�. Further research encompassing the specific features of ma-sonry is therefore needed.

The success of externally bonded strengthening strategiesdepends to a great extent on the capacity of the FRP-masonryinterface in developing an efficient load transfer mechanism.The occurrence of debonding along an interface is the directcause of failure in a great number of masonry supports. Thisfeature demands a comprehensive characterization of the me-chanical behavior of FRP-masonry interfaces. Only recently, thecharacterization of externally bonded masonry substrates withcomposites has been investigated �Cosenza et al. 2000; Accardiand La Mendola 2004; Aiello and Sciolti 2006; Grande et al.2008; Panizza et al. 2008�, involving the analysis of different keyparameters as substrate materials, composite systems, bondedlength, test setup, among others. In spite of the differences amongresearchers, two aspects seem to be common: use of single unitsand adoption of double-shear test configurations. In addition,writers tend to use carbon fibers.

Different test setups have been adopted for the characterizationof bond behavior between FRP composites and concrete-basedsubstrates, but a unique well-established setup has not been de-fined yet. Existing test setups are usually classified as: single-shear tests �Täljsten 1997�, double-shear tests �Lee et al. 1999�,and beam tests �De Lorenzis et al. 2001�. A comprehensive re-view of available test setups is given in Chen et al. �2001�. Inspite of the simplicity associated to double-shear test configura-tions, the geometric symmetry imposed to the FRP strengtheningsimultaneously in two substrates cannot assure an equally distrib-uted load is applied throughout the test, due to the different in-trinsic nonlinear behavior of the substrate. This might lead todifficulties in interpreting the results. From the three possible con-figurations given above, the single-shear test setup is closer to realapplications, thus being considering the most reliable setup to beused in research �Mazzotti et al. 2009�.

This paper is focused on the experimental bond behavior of
FRP-masonry interfaces, considering both tensile and shear bond.
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Monotonic pull-off and single-shear tests were carried out underdisplacement control. Pull-off tests provide valuable informationabout failure under Mode I, whereas an existing single-shear testsetup �Täljsten 1997� was adapted to be used for the experimentalbond behavior characterization �Mode II�. Within the frameworkof thin externally bonded plates to rigid substrates, Mode I behav-ior is associated to the plate-substrate interface deformation undernormal tensile stresses, while Mode II behavior corresponds tothe interface deformation under shear stresses.

Two types of handmade clay bricks were used for pull off.Different bond lengths, composite materials, anchor schemes, andsupport shapes were considered in the single-shear experiments.Masonry prisms were used instead of single units to include theeffect of mortar joints and also to have specimens representativeof real applications. Glass FRP composites �GFRP� were chosenas the main strengthening material. When compared with carbonfibers, the lower axial stiffness of GFRP seem more appropriate touse for low strength and low stiffness masonry structures, reduc-ing stress concentrations along the bonded length.

The main objective of the present paper is the characterizationof the ultimate load and failure patterns of masonry prismsstrengthened with FRP materials under Mode I and Mode II. Inaddition, the stress profile along the bonded length was character-ized for the different arrangements and the local bond stress-slipcurves were analytically derived from experiments and discussed.

Materials Characterization

FRP sheets were externally bonded to masonry prisms followingthe typical wet layup system, namely the application of an epoxyprimer on the masonry substrate followed by the application ofputty to level the masonry surface and finally the application ofepoxy resin and a single layer fiber sheet oriented perpendicularlyto mortar joints. The mechanic characterization of all componentsis briefly described next, with further details provided in Basilio�2007�.

Bonding Materials

The computation of Young’s modulus, tensile strength and strainat failure of both primer and resin was based on molded dog-boneshaped specimens, prepared according to guideline ISO 527-2�International Organization for Standardization 1993b�. Puttyspecimens were also prepared according to the same guideline.However, due to its relatively high viscosity, the preparation ofputty specimens required a more delicate procedure. Followingspecimen molding and drying, a strain gauge was glued at themiddle of the gauge length and aligned with the longest speci-men’s symmetrical axis. The tensile testing of all specimens wascarried out by means of a Universal Zwick machine with a loadcell capacity of 5 kN, in accordance with guidelines ISO 527-1and ISO 527-2 �International Organization for Standardization1993a,b�. All primer and putty specimens failed at the midgaugelength, the ideal rupture mode. In turn, resin specimens failedclose to the grip, but all failure sections were located inside thegauge length.

Table 1 presents the tensile strength and strain at failure aswell as the Young’s modulus values of the three bonding materi-als. Each value was obtained as the average value of five experi-mental results. The same results are graphically illustrated in
Fig. 1�a� in terms of stress-strain diagrams. The results obtained
JOURNAL OF COMPOSITE


exhibit low coefficient of variation �COV�, which seems to indi-cate satisfactory molding and testing procedures.

The three bonding materials exhibited the typical very fragilebehavior and, therefore, no postpeak behavior was captured.While primer and putty specimens present a clear linear elasticbehavior, resin specimens seem to possess an incipient nonlinearbehavior. Similar experimental data range values of tensilestrength and Young’s modulus were reported for primer and resinby Casareto et al. �2002� and for the putty by Bonaldo et al.�2005�.

Table 1. Tensile Parameters and Young’s Modulus of the Bonding Ma-terials and of the FRPs Employed

Material

Tensile parametersYoung’smodulus

�kN /mm2�Stress

�N /mm2�Strain�%�

Primer 5.9 �3%� 4.76 �12%� 0.32 �10%�

Putty 32.2 �3%� 0.44 �7%� 7.44 �2%�

Resin 48.6 �9%� 1.91 �16%� 3.67 �8%�

Glass fibers 1,473 �15%� 3.66 �12%� 80.2 �4%�

Carbon fibers 2,535 �5%� 2.42 �9%� 215.6 �2%�

Note: Average value of five specimens; the COV is provided inside theparentheses.

Fig. 1. Tensile stress-strain diagrams for: �a� bonding materials �PR,PU, and RE stands for primer, putty, and resin, respectively�; �b�glass and carbon fibers

S FOR CONSTRUCTION © ASCE / JANUARY/FEBRUARY 2011 / 33

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Fibers

The mechanical characterization of the bond behavior throughsingle-shear tests, to be described latter on, was performed usinga unidirectional carbon fiber fabric �reference CF-120, mass den-sity of 0.20 kg /m2� and a bidirectional glass fiber fabric �refer-ence G-AR 90/10, mass density of 0.44 kg /m2�, with 90% of thefibers oriented in the main direction. The purpose of using bothfibers was to understand the potential influence of the FRP mate-rial on the bond behavior.

All specimens were prepared and tested according to theguideline ISO 527-5 �International Organization for Standardiza-tion 1997�. The fibers were wrapped with a tiny layer of resin anduniformly brushed on both sides. After cutting, specimens weremounted and tested in a universal Instron testing machine with aload cell capacity of 50 kN. Recommendations of good practiceprovided by the manufacturer were followed during the entire testprocedure. Results of tensile strength, strain at failure andYoung’s modulus are displayed in Table 1 �average values fromfive specimens�, while the tensile stress-strain curves are depictedin Fig. 1�b�, for both glass and carbon fibers.

The tensile behavior is characterized by an almost linearstress-strain prepeak relationship. Even though both materialsfailed to exhibit a postpeak branch, carbon fibers had a suddenand explosive failure when compared with the smoother mis-aligned thread of glass fibers at collapse. Fig. 1�b� shows clearlythat the lower stiffness and larger failure strain exhibited by glassfibers makes them a better option from the mechanical point ofview to be used in low strength and low stiffness masonry struc-tures.

Masonry

Aiming at replicating ancient brick masonry structures, two dif-ferent Portuguese traditional solid clay bricks and a commercialpremixed hydraulic lime-based mortar �Mape-antique MC� wereselected to construct the masonry prisms to be strengthened withcomposite materials. The two types of bricks, here termed asBrick Type 1 and Brick Type 2, have different geographical ori-gins and sizes. The resulting masonry prisms were prepared withthe same mortar and are termed here as Masonry Type 1 �MT1�and Masonry Type 2 �MT2�. A detailed description on the geo-graphical origin can be found in Basilio �2007�.

Tests performed on eight cylindrical mortar specimens withdimensions �50�110 mm2 provided an average compressivestrength of 7.3 N /mm2 and a COV of 8%. The compressivestrength and the Young’s modulus of masonry were assessedthrough the testing of 10 representative prisms of each masonrytype, after two weeks of curing. The dimensions of the prismsreached 235�130�90 mm3 �four bricks of 45 to 50-mm thick-ness and three mortar joints of approximately 10- to 15-mm thick-ness� for MT1, and 163�100�50 mm3 �five bricks of 25-mmthickness and four mortar joints of 10-mm thickness� for MT2.All tests were performed under monotonic displacement control,allowing the characterization of the softening behavior. For MT1,an average compressive strength of 8.8 N /mm2 �COV=27%�and an average elastic modulus of 1,300 N /mm2 �COV=39%�were obtained, whereas for MT2, an average compressivestrength of 9.1 N /mm2 �COV=16%� and an average elasticmodulus of 2,040 N /mm2 �COV=34%� were computed. Despitethe moderate to high COV’s found and the different geographicorigins of bricks, the compressive strength of both masonry types
can be considered similar.


Pull-Off Tests

The failure of externally bonded quasi-brittle substrates �as ma-sonry� with FRP sheets is generally controlled by the interfacialshear and normal stress distributions �De Lorenzis and Zavarise2008�. It is here assumed that the tensile failure of the FRP sheetsis reached for loads much higher than of the interface. For jointssubjected to predominant shear stresses, debonding takes place�Mode II failure�. However, in the proximity of inclined cracks orcurved substrates, a mixed-mode condition occurs, where the in-terfacial normal strength may play an important role in the defi-nition of the collapse load �De Lorenzis and Zavarise 2008�. Bymeans of the standard pull-off test applied to FRP-masonry inter-faces, it is possible to obtain a good estimation of its tensilestrength, under Mode I failure.

Test Setup and Test Procedure

For the pull-off tests, two types of masonry �MT1 and MT2� andtwo types of FRP fibers �carbon and glass� were used. For eachcombination of masonry and FRP types, five specimens weretested. Following the cleaning of the masonry substrate, the com-posite system was applied according to the wet layup system.After primer application, a FRP strip of 70-mm width was gluedon masonry. Afterwards, a partial-depth core with 49-mm diam-eter and 10-mm depth was drilled, see Fig. 2, leading to the cre-ation of circular strengthened areas where rigid steel plates wereglued. The drillings were centered with the masonry joints. Thearea of the masonry cylinder to be tested under direct tensile loadwas composed, on average terms, of 2/3 of brick and 1/3 of mor-tar. The adopted location of the cylinders is justified by the needof involving the mortar on the interface load capacity assessment.

The tests were carried out using a Proceq DYNA Z15 portablemachine. A constant velocity rate of 20 kPa/s was applied so theultimate load was attained in about 1 min. These tests were per-formed following the guideline ASTM D4541-02 �ASTM 2002�.

Normal Tensile Strength

Table 2 illustrates the average pull-off test results for the fourcombinations of masonry and FRP types. The results indicate that

Fig. 2. Typical setup for pull-off tests: �a� plan view; �b� elevationview �units are in millimeters�

the pull-off strength is practically independent of the FRP mate-

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rial and masonry type, despite the moderate COV. This featurecan be explained based on the failure pattern registered. For allspecimens tested, failure was characterized by the ripping of athin layer of brick and mortar �peeling�. The rigid plate surfacewas fully covered with brick and mortar, which exhibited an ir-regular thickness variation. The results obtained show that thetensile strength of the interfaces tested depends on the tensilestrength of the substrate, which is the weakest element of theFRP-resin-masonry assemblage under direct tensile loading�Mode I�, and not on the tensile strength of the bonding agents.These results seem to indicate that for low strength substratesstrengthened as described above, the tensile strength of the sub-strate is much lower than of the bonding agents, see also Table 1.Within this context, the use of a cheaper and low strength bondingagent seems to be advantageous. This result also shows that thetwo types of bricks used possess a similar tensile strength, despitethe differences in terms of geographical origin and productionprocedures, which corroborate the conclusions about the com-pressive strength given above.

Table 2. Pull-Off Test Results

Masonrytype

Fibertype

Tensile strength�N /mm2�

MT1 Glass 1.1 �18%�

Carbon 1.2 �27%�

MT2 Glass 1.2 �20%�

Carbon 1.3 �22%�

Note: Average value of five specimens; the CoV is provided inside theparentheses.

Bon

C150RS

FRP Material

G150RS

Subst

Fig. 3. Different arrangements deriv



Experimental Bond Behavior Characterization

Test Setup and Test Procedure

The masonry prisms with dimensions 235�130�90 mm3 wereprepared and strengthened according to the techniques describedabove. Given the mechanical resemblance of the two masonrytypes used for the pull-off tests, only masonry prisms Type 1 wereused here.

The prisms were strengthened with a glass FRP strip of 25-mmwidth and 150-mm length. This specimen was taken as a refer-ence from which several variations were adopted. To evaluate theinfluence of the major parameters that dominate the bond behav-ior of the FRP-masonry interface, the following variations withrespect to the reference specimen were considered, see alsoFig. 3:• Bonded length;• FRP material;• Use of anchor schemes; and• Shape of substrate.

Therefore, eight different types of arrangements were tested,as illustrated in Fig. 3, being termed according to the followingdescription:• G150RS: reference or standard specimen, with a flat surface

and strengthened with a 25-mm width and 150-mm lengthglass FRP strip. No anchor devices were used.

• G100RS: the difference between this specimen and the previ-ous one is only related to the bonded length, being now 100mm.

• G200RS: a 200-mm bond length was considered in this ar-

th

G100RS G200RS

chorheme

G150RI G150RT

ape

G150ES G150XS

m the reference G150RS specimen

d leng

Ansc

rate sh

ed fro


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rangement with respect to the reference specimen.• C150RS: this specimen is equal to the reference one except for

the composite material, made now of carbon fibers instead.• G150RI: in this specimen, a nonconventional anchor scheme

based on the use of flexible GFRP-based spike anchors wasimplemented �Casareto et al. 2002�. Glass fibers impregnatedwith resin were inserted into a predrilled 6-mm diameter hole,close to the free unloaded end, and fanned out over the strip.Only one spike anchor was considered per specimen.

• G150RT: another nonconventional anchor scheme was testedon this specimen. A FRP bar of 45-mm length and 6-mm di-ameter was placed transversally over the strip, close to the freeend, and glued to the strip and to the substrate. A groove ofabout 8-mm diameter was previously made on the brick inorder to accommodate the bar, which was intended to me-chanically anchor the strip as to increase the ultimate load.

• G150ES: in this specimen, a concave surface shape �r=750 mm� was adopted instead, aiming at representing theintrados surface of masonry arches and vaults, keeping con-stant the remaining parameters with regard the reference speci-men G150RS.

• G150XS: For this specimen, a convex surface shape �r=750 mm� was considered for the application of the GFRPstrip, to replicate the extrados conditions of masonry archesand vaults. All the remaining parameters were kept unchanged

Fig. 4. Experimental test setup developed: �a� general view of aprism externally strengthened with a FRP strip; �b� elevation view ofthe rigid steel device and specimen with location of LVDTs and straingauges

with respect the reference specimen.



The steel device used to test the prisms was composed of an Hprofile welded to a rigid plate and stiffened by means of twowelded lateral bars, see Fig. 4. The prisms were positioned on thesteel device and firmly clamped to it as to avoid any type of rigidbody movement, while the device was mounted on the UniversalInstron testing machine aforementioned, able to work under axialdisplacement or load control. All masonry surfaces in contactwith the steel device were carefully smoothed to minimize anynon-uniform stresses distribution, thus leading to an enhancedclamping solution. The tests were carried out under displacementcontrol, by imposing a constant displacement rate of 0.1 mm/minat the loaded end of the FRP strip, while the resulting load wasmeasured by means of a load cell. All tests were carried out up tofailure. In total, five specimens of each series were tested, result-ing in 40 specimens successfully tested.

The relative displacements between the FRP and the substratewere measured at the loaded �LVDT1� and free ends �LVDT2�,see Fig. 4�b�. Any possible rigid body movement of the steeldevice was also monitored �LVDT3�. For each specimen, twostrain gauges were used at the thirds of the bonded length, seealso Fig. 4. By assuming that strain is constant in the compositecross section and that the FRP performs in linear elastic regime, itis possible to know the FRP tensile stress at four points along thebonded length.

Ultimate Load

Table 3 presents the tensile stress applied to FRP that causedfailure of the strengthening, in terms of average and COV values.The results obtained show that by decreasing the bond lengthfrom 150 �G150RS� to 100 mm �G100RS� a lower ultimate loadis reached, as expected, resulting in an average decrease of 28%.In turn, when shifting from 150 to 200 mm �G200RS�, the incre-ment in load capacity is relatively small, about just 3%. Theseresults are due to the change of failure mode. For the bond lengthof 100 mm, all specimens failed due to debonding of the FRPwith the detachment of a thin layer of brick and mortar, as men-tioned in previous researches �e.g., Aiello and Sciolti 2006�. Inthe case of 150-mm bond length, one of the five specimens faileddue to tensile rupture of the FRP strip. Furthermore, for the caseof 200-mm bond length, three specimens out of five failed due totensile rupture of FRP strip, whereas the average failure stress ofG200RS series �Table 3� represents 90% of the average tensilestrength of glass fibers �Table 1�. Most probably, further increasesof the bond length would lead to marginal increments of the load

Table 3. Normal Strength of the Eight Different Bond Arrangements

Parameter Specimen series

Normal strength

Average�N /mm2�

CoV�%�

Bond length G100RS 921 12

G150RSa 1,276 20

G200RS 1,318 11

FRP material C150RS 1,482 6

Anchor scheme G150RI 1,614 16

G150RT 1,316 13

Substrate shape G150ES 1,221 15

G150XS 1,112 9aReference or standard series.

capacity. In conclusion, the results seem to show that the 150-mm

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bond length is close to the optimal bond length, meaning thatfurther increases of this length do not produce noticeable incre-ments of the ultimate load.

The results described above suggest that the use of a strongerFRP material would allow an increase of the ultimate load. In-deed, the experimental results confirm this hypothesis. The use ofcarbon fibers �C150RS� instead of glass ones �G150RS� allowedan increase of the ultimate load of about 16%, as shown inTable 3. Due to the higher tensile strength of the carbon fibers, thetensile failure was prevented and thus all specimens failed bydebonding, making clear that the optimal bond length increases ifstronger fibers are used.

As for the anchor schemes tested, the effect of a handmadeflexible spike placed close to the free end �G150RI� was notice-able. The anchorage effect promoted by the spike shifted the fail-ure mode from debonding to tensile failure. As a result, theaverage ultimate load registered an increase of about 26%. Thesecond anchor scheme �G150RT�, composed of a FRP bar placedtransversally over the strip and glued to it and to the substrate, didnot seem to be effective as only a marginal increase of 3% of theaverage ultimate load was registered. This means that the adoptedarrangement was not able to provide an enhanced anchoring ef-fect. In line with this result, failure was due to debonding of theFRP strip.

The consideration of curved substrates introduces normalstresses along the FRP-masonry interface due to equilibrium, ei-ther of tensile �concave substrates� or compressive �convex sub-strates� nature. In particular, it was expected that the interactionbetween the normal tensile and shear stresses �at the concaveinterfaces� leaded to a reduction of the ultimate load, as numeri-cally foreseen �De Lorenzis and Zavarise 2009�. The failure of thespecimens with curved substrates was due to debonding, with novisual differences with respect to flat substrates. The experimentsshowed a 4% decrease of the average strength of concave speci-mens �G150ES�, see Table 3. This reduced variation in strength

Fig. 5. Typical failure modes due to debonding of the FRP strips withdetachment of thin layers of brick and mortar

might suggest a marginal influence of normal stresses in failure.



However, it is possible that the influence of normal stresses in thedebonding process might increase for smaller radii of curvature,thus implying larger reductions in strength. As for the convexspecimens, results showed a decrease of 13% in comparison withthe reference series. This was an unexpected result, as a decreasein load was not foreseen. Furthermore, similar experiments areplanned to better understand the influence of curvature.

The COV range from 6 and 20%, which seems to be slightlyhigh considering the laboratory conditions used, even if moderateto high COV values are usually associated with masonry. How-ever, other writers have reported even higher COV values �e.g.,Aiello et al. 2007�. It is possible that the presence of mortar jointsmight constitute an additional source of scattering with regard tothe use of single units.

In general, the specimen’s failure was due to debonding of theFRP strip, see Fig. 5, except for the cases where tensile failure ofthe FRP took place �as addressed above�. In both failure modes abrittle behavior was registered. In the case of tensile failure, thiswas due to the well-known brittleness of the FRP, which has nopostpeak, see also Fig. 1�b�. For the general case, the brittle be-havior of the specimens was due to the brittleness of the brick�Basilio 2007� combined with the inevitable release of elasticstrain energy stored in the specimen during the loading process.Preliminary tests using displacement control options were at-tempted, but it was not possible to avoid uncontrolled failure ofspecimens due to debonding.

Load Transfer

The load transfer phenomenon between the tensile FRP strip andthe brick masonry can be followed in Figs. 6–9, where the aver-age stress profiles at the FRP strip are shown for 25, 50, 75, and

Fig. 6. Average stress profiles in the FRP strip for 25, 50, 75, and100% of the ultimate load: �a� G150RS; �b� C150RS

100% of the ultimate load. The stress was computed based on the


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strain measurements and assuming linear elastic behavior for theFRP. The depicted lines represent linear interpolations betweenmean values of five specimens tested within each series. As ex-plained previously, it became impossible to register the stress pro-file during the softening regime.

The normal stress increases from the free to the loaded endand with the applied load, in every cross section �except for thefree end, where it is assumed to be equal to zero�, but its variationdepends on specimen’s series. In the same way, the shape of thestress profile depends on the series. For low load levels, the stressprofile is essentially exponential, for all specimens, indicating thatthe load transfer occurs along a short length close to the loadedend. With the increase of load level a longer load transfer lengthis mobilized. In most of the specimens, the initial exponentialstress profile was progressively substituted by a quasi-triangularone, see Figs. 6–9. Analog results with the same trend have beenreported by other writers for concrete �e.g., Bizindavyi and Neale1999� and masonry �e.g., Aiello and Sciolti 2006�. The exceptionsare the C150RS series, where a kind of extended S-shape profilewas observed, and the G200RS series, where the stress profile iskept exponential along the entire load procedure. Most probably,the “premature” tensile failure of the FRP strip in the G200RSseries �in the sense that the interface was able to sustain furtherload� did not allow the evolution of the stress profile. Apparently,the anchor schemes used did not have a visible influence on thestress profile, inducing almost triangular stress profiles at the peakload. As for the concave and convex surface shapes, it seems thatthe stress profile tends to exhibit a slightly convex shape at theultimate load �see also Fig. 9�.

Analytical Approach

The equilibrium of a piece of FRP strip of length dx bonded to

Fig. 9. Average stress profiles in the FRP strip for 25, 50, 75, and100% of the ultimate load: �a� G150ES; �b� G150XS

Fig. 7. Average stress profiles in the FRP strip for 25, 50, 75, and100% of the ultimate load: �a� G100RS; �b� G200RS

Fig. 8. Average stress profiles in the FRP strip for 25, 50, 75, and100% of the ultimate load: �a� G150RI; �b� G150RT
masonry is expressed according to
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d�FRP�x�wFRPtFRP = ��x�wFRPdx �1�

where �FRP=normal stress at the FRP strip; �=bond stress at theFRP-masonry interface; wFRP and tFRP=width and the thickness ofthe FRP strip, respectively; and x=longitudinal coordinate, withorigin at the free end. Assuming linear elastic behavior of thecomposite, Eq. �1� can be rewritten as

��x� = EFRPtFRPd�FRP�x�

dx�2�

where EFRP and �FRP=Young’s modulus and the normal strain ofthe FRP strip, respectively. By assuming that the deformability ofthe epoxy adhesive and masonry are negligible when comparedwith the FRP deformability, �FRP can be expressed simply as

�FRP�x� =ds�x�

dx�3�

where s=slip between masonry and the reinforcement. IntegratingEq. �3� and neglecting the slip value at the free end �confirmed byLVDT2 readings�, the slip at a distance x from the free end reads

s�x� =�0

x

�FRP�x�dx �4�

By considering the available discrete strain readings �FRP�xi�along the bonded length, s�x� and ��x� can be approximated nu-merically, by integration of Eq. �4� and differentiation of Eq. �2�,as follows:

s�xi� =1

2�i

��FRP�xk� + �FRP�xk−1��xk − xk−1� �5�

Fig. 10. Envelope of five analytical bond stress profiles for G150RSseries: �a� 50% of the ultimate load; �b� 100% of the ultimate load

k=1



��xi� =1

2EFRPtFRP��FRP�xi� − �FRP�xi−1�

xi − xi−1+

�FRP�xi+1� − �FRP�xi�xi+1 − xi

��6�

Analytical curves might show a certain degree of irregularity, asthey are based on discrete and limited strain readings. Whereasnumerical integration is a smoothing process, numerical differen-tiation tends to magnify irregularities. In order to smooth the ana-lytical curves, additional discrete “readings” were consideredbetween each two consecutive real strain records, as average val-ues of adjacent readings. As an example, Fig. 10 shows envelopesof analytical bond stress profiles for G150RS series, considering50 and 100% of the ultimate load.

Local Bond Stress-Slip Relationship

The local bond stress-slip relationship ��s� can be generated bycombining the individual s�x� and ��x� curves, obtained numeri-cally, and drawn for each x coordinate along the bonded length.Diagrams from Figs. 11–14 represent the interfacial local bond-slip curves derived for all the series, computed at the loaded end,so a larger extension can be observed. In spite of the differencesamong series and among specimens within the same series, aswell as errors due to numerical differentiation, a general trendmight be observed. For low loads, bond-slip curves show an ini-tial linear elastic stage characterized by similar stiffness withineach series. With increasing loading, a hardening stage is origi-nated at around 60–80% of the ultimate load. The extension ofthis hardening branch is variable, ranging from an incipient andshort nonlinear behavior to a progressive nonlinear behavior. Al-though differences exist among series, a trend is visible within

Fig. 11. Local bond stress-slip curves at the loaded end: �a� G150RS;�b� C150RS

each one.


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In some specimens it is possible to identify a ductile phase inthe sense that specimens can sustain a constant bond stress withincreasing slip. The next stage is characterized by a marked soft-ening behavior, although this stage was not registered in all speci-

Fig. 13. Local bond stress-slip curves at the loaded end: �a� G150RI;�b� G150RT

Fig. 12. Local bond stress-slip curves at the loaded end: �a� G100RS;�b� G200RS



mens. Available results are not totally conclusive about theexistence of residual bond strength after debonding, due to fric-tion over the debonded length. While some specimens indicate theabsence of residual bond strength, others suggest its existence.

Concluding Remarks

The experimental performance of FRP strengthened clay masonryprisms under tensile and shear loading was presented and dis-cussed in the paper. The results from pull-off tests show thatfailure was due to the ripping of a thin layer of brick and mortar,making the substrate the weakest element �lower tensile strength�of the FRP-resin-masonry assemblage under Mode I. Therefore,the tensile strength of the brick and mortar controls the tensilestrength of the FRP-masonry interface. Within this context, itseems reasonable to use low strength bonding agent in thestrengthening of �low strength� masonry constructions.

The increase of the bonded length allowed shifting from deb-onding to tensile failure, with location of the optimal bond length.It was observed that the value of this parameter depends on theFRP tensile strength. Debonding of the FRP involved always thedetachment of a thin layer of brick and mortar. The use of asimple handmade spike placed at the free end was very effective,allowing an increase of the average ultimate load of about 26%.This result shows that suitable anchorage schemes might be ef-fective in preventing or delaying debonding phenomena in ma-sonry structures. The use of concave substrates caused a slightdecrease of the ultimate load. This was an expected result, how-ever the experimental behavior of curved substrates �including thevariation of the radius of curvature� requires more attention fromresearchers as available experimental information is almost absent

Fig. 14. Local bond stress-slip curves at the loaded end: �a� G150ES;�b� G150XS

from literature.

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The analysis of the stress profiles in bond specimens allowedunderstanding how stress varies along the bonded length and be-tween load increments. For low load levels, the stress profile isexponential, showing that load transfer occurs along a shortlength close to the loaded end. With the increase of the load level,a longer load transfer length is mobilized and the initial exponen-tial stress profile was progressively substituted by a quasi-triangular one.

The development of an analytical approach supported on theexperiments allowed drawing the local bond-slip curves. Thesecurves show an initial linear elastic stage followed by a hardeningstage at around 60–80% of the ultimate load. In some specimensit was possible to identify a ductile stage. The next stage wascharacterized by a noticeable softening behavior, although thisstage was not registered in all specimens. Available results are notconclusive about the existence of residual bond strength after deb-onding.

The local bond stress-slip curves seem somehow to be differ-ent from the ones derived from externally reinforced concretespecimens �e.g., see Yuan et al. 2004�, so further research shouldbe developed on this issue, as bond stress-slip laws suitable formasonry are necessary. Another relevant feature is the need for asuitable test setup configuration for masonry bond tests, probablybased on a single-shear test setup, able to provide an efficientcharacterization of the softening regime.

Acknowledgments

The financial support provided by the Portuguese Science andTechnology Foundation through Project No. POCTI/ECM/38071/2001 is acknowledged. FRP composites and premixed mortarused in the laboratory tests were gently offered by BETTORMBT and MAPEI, respectively, to whom the writers are grateful.The second writer would like to express his gratitude to CONA-CYT for the scholarship granted.

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experimental bond behavior of frp sheets glued on brick masonry

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