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Shear behaviour of masonry walls strengthened by external bonded FRP and TRC Thi-Loan Bui a , A. Si Larbi b,, N. Reboul c , E. Ferrier c a Institute of Construction Engineering, University of Transport and Communications, 3 Cau Giay, Lang Thuong, Dong Da, Ha Noi, Viet Nam b Université de Lyon, Ecole Nationale d’Ingénieurs de Saint-Etienne (ENISE), Laboratoire de Tribologie et de Dynamique des Systèmes (LTDS), UMR 5513, 58 rue Jean Parot, 42023 Saint-Etienne Cedex 2, France c LGCIE, Université Claude Bernard LYON 1, 82 bd Neils Bohr, 69622 Villeurbanne, France article info Article history: Available online 2 July 2015 Keywords: Masonry wall Strengthening FRP TRC Shear behaviour abstract This experimental study focuses on the behaviour of hollow concrete brick masonry walls, especially walls reinforced with composite materials under in-plane loading conditions. This work is a step towards defining reliable seismic strengthening solutions. Indeed, in France, more stringent seismic design requirements for building structures have been considered with the replacement of old design codes. Thus, an experimental program has been performed at the laboratory scale. Six walls have been submit- ted for shear–compression tests – five walls are reinforced by (1) – fibre-reinforced polymer (FRP) strips using E-glass and carbon fabrics and/or (2) a textile-reinforced concrete (TRC), and the last wall acts as a reference. It is noted that the composite strips are mechanically anchored into the foundations of the walls to improve their efficiency. All of the walls share the same boundary and compressive loading conditions, which are representative of a seismic solicitation. Nevertheless, masonry wall performances and anchor efficiency are only evaluated under monotonic lateral loadings. A comparative study on global behaviour and on local mechanisms is performed and, in particular, highlights that the mechanical anchor systems play an important role in improving the behaviour of reinforced walls (by FRP and TRC) and that the solutions for strengthening by TRC permit the upgrade of the walls’ ductility with a lower strength compared with the solutions with FRP. Ó 2015 Elsevier Ltd. All rights reserved. 1. Introduction Masonry has a long history as a building technique. Even if rein- forced concrete and steel prevail in the modern structures, masonry units are also used. In France, a significant part of build- ings is erected with hollow concrete blocks. However, a relatively important manufacturing tolerance and a design with large holes give these blocks – and even more to hollow concrete block struc- tures – a complex behaviour. Therefore, it is obvious that we should pay attention to these structures in a seismic context, par- ticularly when a seismic hazard assessment has been revised, lead- ing to a tightening of the safety rules in France. Indeed, past earthquakes have revealed that unreinforced masonry structures can suffer extensive damage. Their vulnerabil- ity often lays in the weakness of mortar joints in tension and shear, which are adversely and highly subjected to shear stresses during earthquakes [1,2]. In brief, due to seismic actions, walls in a building can be subjected to shear forces both in the in-plane and out-of-plane directions. The in-plane structural walls (i.e., shear walls, subjected to lateral load along their longitudinal axis) are the primary force resisting elements [3]. Out-of-plane walls (i.e., flexural walls, subjected to lateral load transverse to their longitudinal axis) are in turn excited and if they are not resistant enough, their collapse may disrupt the stability of the building and can result in a major loss of life and property. Although these out-of-plane failures should not be overlooked, practitioners (in a broad sense, including the scientific community) tend to make the in-plane seismic response of shear walls their first priority; they indeed appear as key vertical components to bear seismic loading. Solutions for repairing or strengthening masonry structures are many and are varied. Nevertheless, externally bonded fibre- reinforced polymer (FRP) composites are often preferentially cho- sen by prime contractors [4], mostly because of their lightweight and their ease of use. However, the reinforcing efficiency of FRP is rarely fully valued when they are only externally bonded to structural elements. FRP mechanical properties are limited because http://dx.doi.org/10.1016/j.compstruct.2015.06.057 0263-8223/Ó 2015 Elsevier Ltd. All rights reserved. Corresponding author. Tel.: +33 4 77 43 75 38; fax: +33 4 78 43 33 83. E-mail address: [email protected] (A. Si Larbi). Composite Structures 132 (2015) 923–932 Contents lists available at ScienceDirect Composite Structures journal homepage: www.elsevier.com/locate/compstruct

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paper

Transcript of shear wall 8

Page 1: shear wall 8

Composite Structures 132 (2015) 923–932

Contents lists available at ScienceDirect

Composite Structures

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

Shear behaviour of masonry walls strengthened by external bonded FRPand TRC

http://dx.doi.org/10.1016/j.compstruct.2015.06.0570263-8223/� 2015 Elsevier Ltd. All rights reserved.

⇑ Corresponding author. Tel.: +33 4 77 43 75 38; fax: +33 4 78 43 33 83.E-mail address: [email protected] (A. Si Larbi).

Thi-Loan Bui a, A. Si Larbi b,⇑, N. Reboul c, E. Ferrier c

a Institute of Construction Engineering, University of Transport and Communications, 3 Cau Giay, Lang Thuong, Dong Da, Ha Noi, Viet Namb Université de Lyon, Ecole Nationale d’Ingénieurs de Saint-Etienne (ENISE), Laboratoire de Tribologie et de Dynamique des Systèmes (LTDS), UMR 5513, 58 rue Jean Parot, 42023Saint-Etienne Cedex 2, Francec LGCIE, Université Claude Bernard LYON 1, 82 bd Neils Bohr, 69622 Villeurbanne, France

a r t i c l e i n f o

Article history:Available online 2 July 2015

Keywords:Masonry wallStrengtheningFRPTRCShear behaviour

a b s t r a c t

This experimental study focuses on the behaviour of hollow concrete brick masonry walls, especiallywalls reinforced with composite materials under in-plane loading conditions. This work is a step towardsdefining reliable seismic strengthening solutions. Indeed, in France, more stringent seismic designrequirements for building structures have been considered with the replacement of old design codes.Thus, an experimental program has been performed at the laboratory scale. Six walls have been submit-ted for shear–compression tests – five walls are reinforced by (1) – fibre-reinforced polymer (FRP) stripsusing E-glass and carbon fabrics and/or (2) a textile-reinforced concrete (TRC), and the last wall acts as areference. It is noted that the composite strips are mechanically anchored into the foundations of thewalls to improve their efficiency. All of the walls share the same boundary and compressive loadingconditions, which are representative of a seismic solicitation. Nevertheless, masonry wall performancesand anchor efficiency are only evaluated under monotonic lateral loadings. A comparative study on globalbehaviour and on local mechanisms is performed and, in particular, highlights that the mechanicalanchor systems play an important role in improving the behaviour of reinforced walls (by FRP andTRC) and that the solutions for strengthening by TRC permit the upgrade of the walls’ ductility with alower strength compared with the solutions with FRP.

� 2015 Elsevier Ltd. All rights reserved.

1. Introduction

Masonry has a long history as a building technique. Even if rein-forced concrete and steel prevail in the modern structures,masonry units are also used. In France, a significant part of build-ings is erected with hollow concrete blocks. However, a relativelyimportant manufacturing tolerance and a design with large holesgive these blocks – and even more to hollow concrete block struc-tures – a complex behaviour. Therefore, it is obvious that weshould pay attention to these structures in a seismic context, par-ticularly when a seismic hazard assessment has been revised, lead-ing to a tightening of the safety rules in France.

Indeed, past earthquakes have revealed that unreinforcedmasonry structures can suffer extensive damage. Their vulnerabil-ity often lays in the weakness of mortar joints in tension and shear,which are adversely and highly subjected to shear stresses duringearthquakes [1,2].

In brief, due to seismic actions, walls in a building can besubjected to shear forces both in the in-plane and out-of-planedirections. The in-plane structural walls (i.e., shear walls, subjectedto lateral load along their longitudinal axis) are the primary forceresisting elements [3]. Out-of-plane walls (i.e., flexural walls,subjected to lateral load transverse to their longitudinal axis) arein turn excited and if they are not resistant enough, their collapsemay disrupt the stability of the building and can result in a majorloss of life and property. Although these out-of-plane failuresshould not be overlooked, practitioners (in a broad sense, includingthe scientific community) tend to make the in-plane seismicresponse of shear walls their first priority; they indeed appear askey vertical components to bear seismic loading.

Solutions for repairing or strengthening masonry structures aremany and are varied. Nevertheless, externally bonded fibre-reinforced polymer (FRP) composites are often preferentially cho-sen by prime contractors [4], mostly because of their lightweightand their ease of use. However, the reinforcing efficiency of FRPis rarely fully valued when they are only externally bonded tostructural elements. FRP mechanical properties are limited because

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Table 1Mechanical characteristics of composites.

Compositestrengtheningsystem

Nominalthickness(mm)

Youngmodulus(GPa)

Tensilestrength(Mpa)

Ultimatestrain (lm/m)

CFRP 0.48 105 1700 16000GFRP 1.7 7.2 100 13.800

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of the debonding of the composite sheets. To address this issue, anadequate mechanical anchorage system needs to be set up toenhance the bond (between a masonry structure and its foundation)performance. The benefits of this solution – in terms of the FRPefficiency and lateral load resistance of a masonry wall – havenow been widely acknowledged in the case of out-of-planeactions [5].

In addition, in the context of sustainable development andhealth and safety conditions for workers, consideration shouldbe given to an alternative material to FRP, which is oftenmanufactured with highly toxic epoxy resins. The idea is tosubstitute these resins with cementitious materials while pre-serving or even improving the dissipative capacity of reinforcedstructures. From this perspective, textile-reinforced concrete(TRC) composites, which combine a suitable fine-grained mortarwith the latest generation of textile fabrics, would benefit frompromotion.

The efficiency of TRC for strengthening masonry structures hasrecently been investigated [6–11]. Compared with FRP, TRC com-posites show a nonlinear tensile behaviour with multiple matricescracking, giving them a greater deformation capacity, a priori moresuitable for seismic reinforcement [8].

Although instructive, these studies lack diversity for studiedmaterials, reinforcement configurations, applied normal loadsand slenderness ratio of walls. Sometimes, a small amount of infor-mation is known regarding damage and failure mechanisms orregarding the interaction between masonry material andreinforcements.

On the one hand, this work is aimed at further developing theexisting experimental database, with special emphasis onidentifying the performances of anchorage devices, particularlyin the framework of a comparative study between FRP andTRC composites. This comparison will cover criteria at the globalscale and, to a lesser extent, at the local scale. On the otherhand, this paper tries to help identify and clarify damagedissipative mechanisms and their impact on the failure modesof the masonry walls.

To attain the aforementioned objectives, an experimental cam-paign has been performed, based on static monotonic shear tests,which are a simplified way to simulate stress states resulting fromearthquakes. Certainly, inertial effects and the inherent cyclicalnature of seismic actions are not addressed in the present study.However, this work can be regarded as a first step towards thedefinition of efficient reinforcement solutions. The approach is totest some strengthening configurations to have relevant andvaluable information and to offer prospects that would beappropriate to assess with more realistic loadings in terms ofearthquake hazards.

« Concrete loading beam »

« Reinforced concrete foundation »

Fig. 1. Description of unreinforce

2. Experimental program

2.1. Masonry walls

A series of six walls has been built with the same dimensiongiven in the Fig. 1. It should be mentioned that all of the specimenswere built by a professional mason and must be considered to be incompliance with the practices. The hollow concrete block units,whose dimensions are 500 mm long, 200 mm high and 75 mmthick, belong to Group 2 according to Eurocode 6, with a strengthclass B40 (characteristic compression strength of 4 MPa).However, these blocks have been halved lengthwise before beingassembled to make walls dimensions compatible with the limitedmeans of the laboratory in terms of space and actuator capacity(Block work size at reduced scale: 250 � 200 � 75 mm3).

The compressive strength of the individual masonry blocks hasbeen determined and ranges from 4 to 10 MPa (6.5 MPa on averagewith a standard deviation of 2.33). These blocks are assembledwith a mortar composed of Portland cement (CEM I 52.5) and sandin the proportion 1:3 with a water/cement ratio equal to 0.5.Mortar test prisms of 40 � 40 � 160 mm3 were tested for compres-sive and flexural strengths. At 31 days, these strengths are 48 MPaand 10 MPa, respectively.

2.2. Reinforcement

2.2.1. Strengthening materialsTwo types of composites have been used: the first composite is

a fibre-reinforced polymer (FRP) while the latter composite is atextile-reinforced cementitious composite (TRC).

2.2.1.1. FRP composite. The fibre-reinforced composite materialsconsist of a two-component epoxy matrix and bi-directional fab-rics made of either carbon (CFRP) or glass (GFRP). Their mechanicalcharacteristics have been measured on six specimens according toISO 527-1. The obtained results are listed in Table 1.

2.2.1.2. TRC composite. Knowledge on TRC composites is notablyless significant than knowledge relating to FRP. However, it is

d masonry wall (reference).

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Table 2Mix design for TRC composite.

Textile reinforcement Micro-mortar**

Nature of fibres Glass-AR Grain size <2 mm*

TEX 1200* Silica-fume Yes*

Fibre diameter 19 lm* Thixotropy Yes*

Number offilament/yarn

1600* Shrinkage �0*

Knitted grid size 5 � 5 mm* Tensile strength 5 MPa**

Tensile strength (yarn) 1102 MPa* Compressivestrength

40 MPa*

Young modulus 74,000 MPa* Young modulus 1700 MPa*

* Provided properties.** Laboratory characterisation.

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currently well established that various levers (roving diameters,yarn number, impregnation, etc.) can be mobilised to optimiseTRC – with occasionally opposite consequences – and many ofthem have already been identified [11].

Therefore, in the course of works performed by Contamine et al.[12], it has been possible to select a composite, whose componentsare listed in Table 2, that results from a compromise between all ofthe aforementioned parameters, including workability and thixo-tropy (to apply strengthening materials more easily).

The composite material contains the 4.36% AR-glass fibre vol-ume fraction, that is, 2.18% in each principal direction. The textileused for reinforcement is a bidirectional warp-knitted grid fabric.

Direct quasi-static tensile tests, whose protocol has been vali-dated [13], have been performed to mechanically characterisethe TRC reinforcements. Stress–strain curves are given in Fig. 2.

Two different behaviour laws appear in Fig. 2. Indeed, withoutthe impregnating resin (latex), TRC exhibits a bilinear behaviourwhereas by using latex, the evolution law is nearly linear becauseof a more homogeneous yarn contribution.

2.2.2. AnchoragesTo reduce or ideally remove the overturning effects due to

lateral loads on walls and to best maximise the potential of eachreinforcement solution by a priori improving its efficiency, aconnector, in the form of an anchorage (see Fig. 3), has been intro-duced between walls (on their lower part, on both sides) and theirfoundation footings.

Fig. 3. Description of a MAPEI anchor.

Fig. 2. Tensile behaviour of TRC materials during uniaxial tensile tests.

Given the existing solutions and their well-established perfor-mances (easy application and high strength), only connectionsolutions based on carbon fibres and an epoxy resin have been con-sidered. The anchorage system (MAPEI) is an anchor made frommonodirectional carbon fibres with at least 36 yarns, each includ-ing 12,000 fibres. The anchorage strength given by the manufac-turer is 30 kN at the ultimate limit state.

2.2.3. Strengthening configurationsThe definition of strengthening patterns must be part of a

strategy aimed at finding a «balance» between lateral strengthand energy dissipation capacity [14]. Thus, the objective in rein-forcing a structure is to improve its strength capacity, to enhanceits ductility, or both.

According to this strategy, different strengthening configura-tions with TRC and FRP composites have been proposed. Thereinforcing material is always applied symmetrically on both wallsurfaces.

The experimental program consists of testing six masonry wallsto failure, including an unreinforced wall (as the reference speci-men) and five TRC or FRP-reinforced specimens (see Fig. 4). WithFRP composites, three strengthening patterns have been proposed.The first wall, referenced CGRW, has been reinforced with both car-bon fibre-reinforced polymer (CFRP) and glass fibre-reinforcedpolymer (GFRP) to significantly improve strength capacity. Eachside of both faces is reinforced by a continuous sheet, is comprisedof two glass layers over a width of 400 mm along the entire heightof the masonry and is combined with two discontinuous carbonsheets, which are 1410 mm long and 60 mm wide; the horizontaldistance between these two carbon sheets is 100 mm. The tworemaining walls, for their part, have been reinforced by eitherCFRP sheets (CRW wall) or GFRP sheets (GRW wall) to combinestrength capacity and ductility.

Concerning TRC solutions, embedding glass fabrics (rather thancarbon fabrics) in an epoxy matrix tends to produce compositeswith better energy dissipation capacities. Thus, it is appropriateto take advantage of this property in the context of reinforcingmasonry walls. This choice to use only glass fibres is more appro-priate because it adds value to TRC materials (compared withcarbon–epoxy composites), which achieves a smaller ecologicalfootprint and has fewer problems with the hygiene and safety con-ditions for workers.

As a consequence, it is clear that without overlooking theloadbearing capacity, it is desirable to increase the ductility of rein-forced walls, which is the main motivation of our choices. In thefirst pattern (TRCW1 wall), only one TRC layer, 1410 mm highand 200 mm wide, is applied on each side of both faces.However, in the second selected pattern, the main objective is toimprove the dissipation capacities, so a vertical strip is applied inthe middle of the wall. Each strip is made of three TRC-layers,which are 1410 mm high and 200 mm wide.

2.2.4. Application of strengthening systems and anchor placementFRP/TRC reinforcements have been laid up. First of all, wall

surfaces have been cleaned to remove dust and loose materials,which could disturb the bond between the masonry wall and itsreinforcements. The application of a strengthening system firstinvolves covering the wall with a layer of epoxy resin (FRP) or mor-tar (TRC). Then, the fabric (FRP) or textile (TRC) is placed along thewall and is pressed against the resin or mortar. A second layer ofresin or mortar is eventually applied to ensure fabric impregnation.For walls reinforced with several layers, the last two steps arerepeated as necessary.

These reinforcements are anchored to the foundation because ofthe CFRP structural connections. An anchorage anchor is composedof two parts: the anchorage strictly speaking and the «whip». The

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anchorage that is first impregnated with epoxy-resin is insertedinto holes that have been drilled before in the reinforced concretefooting and is filled with resin. The remaining part of the anchor(the whip) is deployed similar to a fan on the wall, and resin isagain applied so that the adhesive completely penetrates into thefibres. This splayed part ensures the connection of strengtheningmaterials with the masonry wall.

2.3. Testing procedure

As mentioned above, monotonic in-plane combined compres-sion–shear tests have been selected.

The vertical axial load (N) is centred and applied by means of ahydraulic actuator with 200 kN capacity. This actuator is combinedwith a force sensor that is located on a stiff steel plate (100 mmwide � 100 mm long � 20 mm thick) and is placed on the concreteloading beam. These devices are held by a steel profile connectedto the strong floor through steel rods, which allows us to controlthe applied vertical load (Fig. 5).

The vertical load is should represent the weight of the upperfloors. In literature, this value varies greatly and depends on themasonry compression strength. Tomazevic [15] has proposed toadopt a constant value equal to 20% of the compression strengthfor confined brick masonry. Papanicolaou et al. [9] have workedwith large ranges from 2.5% to 10% for rectangular walls and from10% to 25% for slender walls. This value is between 11% and 16% inF.da Porto’s study [10] regarding steel-reinforced masonry. For agiven wall geometry, the axial compression load magnitude influ-ences the overall performances, in particular, the efficiency ofstrengthening solutions. Papanicolaou et al. [9] have noted thatreinforcement efficiency is reduced as vertical compression loadrises. Because this study is motivated by assessing the shear contri-bution of the reinforcements, a relatively low value of approxi-mately 6% of the masonry compression strength has beenadopted. It corresponds to a mean vertical stress of 0.2 MPa(15 kN).

In practical terms, the vertical load is very slowly applied up tothe target value equal to 15 kN. It is kept constant during the testbecause of a force control of the vertical actuator. At this time, hor-izontal load is imposed under quasi-static monotonic conditions.This loading step is performed under displacement control at a rateof 0.015 mm/s to better capture post-pic behaviour. Lateral loadingis stopped when masonry walls have obviously failed, that is, whenlateral force drops significantly. According to Tomazevic [15], fail-ure is ascertained when horizontal load falls by 20%. This criterionwill be adopted herein.

In addition to the force sensors that have been presented beforein the description of the test set-up, for the sake of clarity, severaldisplacement transducers and strain gauges complete the instru-mentation. Their locations and purposes will now be discussed.

The displacement transducer C1 (LVDT ± 100 mm) measuresthe lateral displacement at the top of the wall and will enable usto characterise the overall behaviour of the wall through load–dis-placement curves. The assumption that the foundation footing hasno slips is controlled because of the displacement transducer C3(LVDT ± 100 mm). To have information at the local scale and toassess the contributions of different reinforcements, strain gauges(120 X), numbered from J1 to J9, are bonded on only one face ofthe walls (Fig. 4).

2.4. Experimental results

2.4.1. Global behaviourThe number of steel rods, their high axial rigidity and their suit-

able tightening, without overlooking the high flexural rigidity oftransverse steel girders, are many arguments for assuming that

walls do not heave. Concerning measures from LVDT C3, theyemphasise that the concrete footing relative to the rigid floor doesnot slide for all six walls. As a consequence, the obtained relation-ships of lateral loads versus horizontal displacements at the top ofthe wall can be considered without caution towards boundaryconditions.

All walls show a nonlinear behaviour after an initial linearelastic branch. A significant deviation is experienced in thelength of the linear zone although stiffnesses are close. Untilthe end of this first phase, wall integrity is preserved at the glo-bal scale. Next, a nonlinear phase starts, which differs dependingon the nature of the walls and reinforcements. This nonlinearzone can be related to the damage (in tension or in shear) ofone or more masonry components, the strengthening materialsor even the reinforcement/block (or mortar joint) interface(interphase).

A dramatic increase in the ultimate strength and in the secondphase stiffness is clearly observed, even if it is conditioned by rein-forcing materials and strengthening configurations. Similarly, dis-sipation capacities increase overall in varying degrees that mustbe assessed and quantified. It must be underlined that only thewall CGRW behaves as brittle (as a result of a sudden and prema-ture failure) (Fig. 6).

To evaluate the performances of these walls in terms of appro-priately chosen and unbiased indicators, the experimentalload–displacement curves will be idealised according to the trilin-ear model proposed by Tomazevic [15] (Fig. 7). Given the experi-mentally observed behaviours, this approach seems to be moreconsistent than the idealised bilinear relationship proposed byMagenes and Calvi [16].

The three phases of the conventional trilinear diagram arebounded by three characteristic points, which facilitate discussionand comparison on wall behaviours. Thus, energy dissipation, cal-culated from the idealised diagrams, will be useful to assess thesuitability of strengthening solutions with dissipation needs inthe context of earthquakes. The first zone, called elastic, ends atlateral load Vcr and displacement dcr. They mark the formation ofthe first significant cracks in the wall, which entail a change instiffness [3]. As masonry walls exhibit highly nonlinear behaviours,a conventional Vcr value is generally adopted. According toTomazevic’s study [15], Vcr is equal to 70% of the maximum resis-tance Vmax. The second zone extends to the maximum lateral loadVmax and displacement dVmax. At last, the ultimate zone is charac-terised by the ultimate load Vdu, corresponding to 80% of the max-imum load and the maximum displacement du on the softeningbranch.

Kel, lu and Ediss are the initial stiffness (the slope of the firstphase, considered as elastic), ductility coefficient and dissipationenergy, respectively. This last parameter provides information onresistance and deformation capability. Indeed, it is determined asthe area below the idealised load–displacement diagram and isgiven by the following equation:

Ediss ¼12½dcr :Vcr þ ðdVmax � dcrÞðVmax þ VcrÞ þ ðdu � dVmaxÞðVmax

þ VduÞ� ð1ÞThe ductility coefficient, obtained by dividing the ultimate horizon-tal displacement (du) by the elastic displacement (dcr), reflects thedeformation capacity of shear walls in the post-elastic zone. Thisparameter is a decisive criterion for paraseismic construction.

2.5. Strength capacity

From these results, it appears that reinforced walls achieve sub-stantially higher ultimate loads than the reference unreinforcedwall, regardless of reinforcement type. Shear strength increases

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CGRW CRW GRW

URW TRCRW1(One layer without latex)

TRCRW2(3 layers with latex)

Fig. 4. Unreinforced wall and FRP/TRC-reinforced walls.

Wall

Threaded rod

Beams

Hydraulic jack

Post-tension rod

Spreader beam

Vertical load actuatorLoad cell

Reaction wall

Load cell

Fig. 5. Test set-up.

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from 110% (TRCW1) to 450% (CGRW). TRC reinforcements areslightly less efficient (175% on average) than FRP reinforcementsin terms of strength capacity. A more detailed analysis allows aquantitative comparison of the different configurations with TRCand FRP materials and highlights that strength capacity gains growproportionally to reinforcement axial stiffness (qv :ER) (The case ofthe CGRW wall has been eluded because it collapsed when theanchorages prematurely failed) – Fig. 9. This relationship is consis-tent with Mahmood et al. [17].

It must be noted that the TRCW1 wall appears as an exceptionbecause its reinforcement has a slight effect on the ultimate load. Itcan be explained by the excessively low stiffness of this reinforce-ment, only comprising one TRC layer.

2.6. Initial stiffness

Initial stiffness is marginally affected by reinforcements. Thiscan stem from the low thickness of strengthening systems, in

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Fig. 6. Curves of load versus horizontal displacement at the top of the wall.

Fig. 7. Idealised tri-linear diagram.

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particular with FRP materials and to a lesser extent with TRC mate-rials. By contrast, the size of the linear zone is strongly subordi-nated to the adopted reinforcements’ nature and configuration.Indeed, according to experimental results (Vcr indicator seemsirrelevant or at least unsuitable for results obtained in the presentcase and presented as load–displacement curves), overall, wallmacroscopic damage is delayed, except for the TRCW1 wall whosereinforcement ratio is low.

In the first stage, the discernible increase in initial stiffness maybe attributed to the ability of the reinforcement (mainly when itsaxial stiffness is sufficient) to bridge emerging microcracks andcracks because of a suitable load transfer.

2.7. Ultimate displacement and ductility

Furthermore, it is important to note that the ultimate displace-ment is a minimum (except for the CGRW wall) maintained whencompared with the reference wall. This emphasises the interest ofchosen reinforcement solutions and a priori attests to the useful-ness of anchorages between footings and walls.

It is also worth noting that two reinforced walls (TRCW2 andGRW) see a tangible increase, of approximately 36% in their ulti-mate displacements. Walls are likely subject to flexural mecha-nisms, which induce normal tensile stresses in reinforcements.Thus, the ultimate displacement capacity of strengthened masonrywalls is, among other factors, certainly conditioned by the axialstiffness of reinforcements, which can explain at least partiallythe increase in stiffness for these reinforced walls.

The decrease in ultimate displacement is peculiar to the CGRWwall because anchorages have certainly been involved in the wallfailure (see below). As mentioned previously, the ductility factor(l = du/dcr) plays a crucial role in assessing the seismic behaviourof masonry walls in particular because it reduces the numberof elastic seismic design actions. Table 3 highlights thatTRC-reinforced walls exhibit nearly the same ductility factor asan unreinforced wall. In contrast, for the FRP-reinforced walls case,if lateral strength clearly increases, the ultimate ductility coeffi-cient is low compared with an unreinforced wall (Fig. 8d andFig. 10). This may be related to the capacity that TRC compositeshave, unlike FRP composites, to crack, to «follow» masonry walldisplacements and potentially to «absorb» wall damage, whichthus remains controlled.

Needless to say, the above assumptions require a more elabo-rate experimental campaign to be validated, but some of the pro-posed explanations will also be reviewed regarding availabledata at the local scale.

2.8. Dissipation energy

In addition to the ductility factor, the dissipation energy (Ediss)will help us position, at least in order of magnitude, the potentialsof FRP and TRC as masonry reinforcements. It is clear that for theadopted configurations, both with TRC (with the exception of theTRCRW1 wall with a very low reinforcement ratio) and FRP, resultsare conclusive because increases in the range of 200–500% can beexpected. To better understand, it is important to correlate theabove performance indicators with observed failure modes, dam-age mechanisms and kinematics and with data at the local scale.This is the purpose of the rest of this paper.

2.8.1. Failure modesThis section is devoted to providing information regarding fail-

ure modes based on reinforcement types and patterns, so that theefficiency of strengthening materials in the global lateral behaviourand their impact on damage mechanisms can be precise.

2.8.1.1. Unreinforced wall. An unreinforced wall exhibits a flexuralfailure mode (Fig. 11) characterised by horizontal cracks on the leftpart of the wall (on the side where lateral load is applied) due totensile stresses in bed mortar joints and by toe crushing (on theright part) at the compressed corner.

To better understand the failure process, the order in whichdamage has been visually detected is as follows. Horizontal cracksstart to open on the left side probably because the tensile bondstrength is exceeded and gradually develops as the crack openinggrows. These flexural cracks involve the first and second rows ofmortar bed joints. It is not excluded that high shear stresses atthe left bottom corner have contributed to the initial failure ofthe first line of bed joints first. The increasing lateral displacementstherefore induce the spreading of cracks through the second row ofbed mortar joints. At this point, the wall tends to rotate about theright bottom corner, thereby inducing the horizontal spread ofcracks (secondarily, vertical joints between the first and secondrows are damaged). As the resistant cross section is reduced, theright bottom corner is compressed. It eventually fails under a com-plex stress state as evidenced by concentrated crushings togetherwith cracks on an approximately 45-degree angle.

It is worth noting that observation of this failure mode hasgiven valuable information for designing strengthening configura-tions in the present study.

More specifically, the objective is to counteract or at least deferdamage of the horizontal joints that are subjected to predomi-nantly tensile stresses because of vertical strips at the ends of walls

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Fig. 8. Comparative diagrams of the different indicators (a-strength capacity, b-stiffness, c-ultimate displacement, d-ductility coefficient, e-dissipation energy).

Fig. 9. Relationship between strength capacity and reinforcement axial stiffness(qv :ER).

Table 3Summary of experimental results.

Walls tR � bR

(mm2)qv ¼ AR

l�tqv :ER

(MPa)Vcr = (70%Vmax)(kN)

dcr

(mm)Vmax

(kN)dVmax

(mm)Vu = (80%Vmax)(kN)

du

(mm)Kel

(kN/mm)

l = du/dcr

Ediss

(kN mm)Failure modes

URW – – – 7.88 0.88 11.25 4.59 9.00 12.98 8.97 14.78 123.90 FlexuralCGRW CFRP 0.48 � 60 3.82 82,020 35.79 3.66 51.13 8.39 40.90 8.70 9.78 2.38 285.30 Flexural + shear

GFRP 1.7 � 400CRW CFRP 0.48 � 60 0.3 31,316 22.40 2.50 32.00 9.94 25.60 13.20 8.96 5.28 324.26 Flexural + shearGRW GFRP 1.7 � 400 3.52 50,702 35.53 3.06 50.75 8.89 40.60 17.70 11.61 5.78 708.24 Flexural + shearTRCRW1 TRC 3 � 200 3.31 4854 8.61 0.75 12.30 5.85 9.84 12.80 11.48 17.07 133.49 Flexural + shearTRCRW2 TRC 9 � 200 13.98 22,135 18.92 1.07 27.03 7.32 21.62 17.60 17.68 16.45 403.80 Flexural + shear

tR- thickness of composite band; bR- width of composite band; AR- total cross section area of strengthening (=P

bR:tR); l and t- width and thickness of masonry wall; qv -vertical reinforcement ratio.

T.-L. Bui et al. / Composite Structures 132 (2015) 923–932 929

(areas where tensile stresses mainly develop during an earthquakeloading cycle).

Moreover, the use of wide strips (or even central strips) aims atcovering a large surface to bridge cracks, while at the same time,

unreinforced zones are maintained as concentrated damage zonesin which energy dissipation can advantageously take place.

Of course, wall behaviour – in terms of lateral strength orenergy dissipation capacity – highly depends on the strength ofanchorages.

2.8.1.2. FRP-reinforced walls. Failure modes of FRP-reinforced wallsare presented in Fig. 12. First, it must be noted that they are differ-ent and depend, among other things, on adopted reinforcementconfigurations. Walls predominantly exhibit shear failures withdifferent crack patterns that also correlate with the damage nature– either brittle failure or progressive degradation – which hasmajor implications on the dissipation capabilities of walls.

The sudden failure of the CGRW wall is singular and resultsfrom the premature failure of anchorages, which are subject to ten-sile forces in the left part (side where lateral load is applied). Thiswall has been reinforced over a large surface area on which bothglass and carbon textiles have been laid up, resulting in a signifi-cant increase in the wall stiffness (compared with an unreinforcedwall but also to other strengthened specimens). Under loading, thewall begins a rigid body motion that induces high shear stresses at

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Fig. 10. Reassessment of ultimate ductility factors by using the experimental elasticlimit rather than the conventional one.

Cracks of mortar joints Crushing of the

compressive brick

Fig. 11. Flexural failure for unreinforced wall (URW).

930 T.-L. Bui et al. / Composite Structures 132 (2015) 923–932

the bottom of the wall which could explain, even partially, theobserved failure mode.

As a consequence, the use of anchorage devices certainlyimproves the lateral strength of reinforced walls. However, toavoid sudden ultimate failures of walls and to ensure sufficientenergy dissipation capabilities, an anchorage design cannot bedecoupled from the stiffness or reinforced walls (taking intoaccount both reinforced surface area and strengthening materials).

The other two FRP-reinforced walls are not affected by anchorfailure and show similar damage mechanisms.

These walls exhibit coupled shear–flexure failure modes. Inboth specimens, shear cracks initiate in the middle of the wall(unreinforced zone) and propagate along the compressed diagonal

(a) (b)

Failure of anchors

Fig. 12. Failure modes and damage mechanisms for FRP-reinforced walls: (a)- sudden fail

for the CRW wall (which is a typical shear failure pattern), evenalong the edge of the unreinforced zone (in particular for GRWwall), thus reflecting that reinforcements can bridge cracks andalso influence their propagation.

Nevertheless, the collapse of these walls (or very marked dete-rioration) occurs by the crushing of the lower right corner (highlysubject to compression) and ends and by the splitting of theextreme unit block. With the failure of this block, the wall tendsto turn about the right toe, as found for an unreinforced wall(Fig. 12(b) and (c)).

2.8.1.3. TRC-reinforced walls. As indicated above for FRP-reinforcedwalls, the addition of TRC reinforcements changes the failuremode. TRC-reinforced walls have failed by combined shear–flexure(Fig. 13). Both reinforcement patterns lead to cracks at the left bot-tom part (side where load is applied) and in strengthening strips(unlike FRP) over horizontal joints.

In the case of the TRCRW1 wall (low reinforcement ratio andlow reinforced surface area), cracks develop along a horizontaljoint and go through a block unit in the centre with an inclinedcrack that suggests a reaction to a shear solicitation.

At the ultimate limit state, a macro crack has been observedthat corresponds to the failure of mortar used in TRC withoutany textile degradation. In scientific literature, this failure modeis commonly referred to as the «peeling-off» failure [18](Fig. 13a).

In the case of the TRCRW2 wall (reinforced by three TRC layerson a larger area), damage and failure mechanisms are different.Cracks grow horizontally and spread over the height of thestrengthening strips. Finally, multi-crack initiation contributes,beyond what is observed on the wall, to the global damage ofthe specimen. It must be noted that until failure, which ultimatelyoccurs by crushing the right bottom corner, crack opening dis-placement (without having been measured) remains limited(Fig. 13b).

In light of the obtained results, it is important to emphasisethat:

- Anchors can be considered useful and efficient systems, buttheir use must be correlated at this time with the global stiff-ness of the wall (reinforcements’ nature and surface area) toavoid their sudden and premature failure.

- Dissipation mechanisms are improved and concentrated for allreinforced walls, regardless of the adopted reinforcements’ typeand pattern. However, the dissipation process is controlledbecause it occurs on the reinforcement itself for TRC materials,whereas it is uncontrolled with FRP-debonding.

(c)

Crushing of the

compressive brick

Crushing ofthe

compressive brick

Cracks’

propagation

Local debonding of FRP

Cracks’

propagation

ure of CGRW wall; (b) and (c)- coupled shear–flexure failure of CRW and GRW walls.

Page 9: shear wall 8

(a) (b)

Macro-crack of TRC Multi-cracks of TRC

Fig. 13. Failure modes and damage mechanisms for TRC-reinforced walls (a)- TRCRW1 wall and (b)- TRCRW2 wall.

Fig. 14. Evolutions of strains in FRP reinforcements along wall length for CGRW wall (a): GRW wall (b) and CRW wall (c).

T.-L. Bui et al. / Composite Structures 132 (2015) 923–932 931

2.8.2. Local behaviourOnly strains measured in FRP-reinforced walls are displayed

because gauges applied on TRC have failed early, without givingvaluable information. Moreover, because of complex local effects(in zones where anchorage cords and FRP composites are heldtogether), only strain gauges (3; 6; 9) that are located aboveanchorage cords are considered. The evolution of strains in verticalFRP reinforcements along the wall length is given in Fig. 14 for dif-ferent loading levels ranging from 0 to Vmax, the maximum lateralstrength. In general, maximum tensile strains are measured withgauges located at the tensile side of the wall, and strains graduallydecrease, quasi linearly, to become negative at the compressedside. These evolutions give credibility to the assumption that planesections remain plane, despite wall heterogeneity and until lateralloading values are close to failure load.

Furthermore, at maximum load, maximum strains in FRP rein-forcements reach values of 4000 lm/m in CFRP strips and9000 lm/m in GFRP strips, respectively, for CRW and GRW walls.These values are very close to the ultimate strain values for glassfibre-reinforced composites and account for only 50% of the ulti-mate strain value for carbon-based reinforcements. Thus, the use-fulness of glass fibres is evident. In contrast, in the case of thehybrid solution with both glass and carbon fibres, strains remainsmall (2500 lm/m – CGRW wall) because of anchor failure. Thislimited strain highlights that without efficient anchorage, rein-forcements cannot significantly contribute to load transfer. Thisis the reason why the global performance of this wall is limited.

3. Conclusions

The present experimental study has focused on masonry wallsreinforced by TRC and FRP composites that are subject to mono-tonic in-plane combined shear–compression tests. The main find-ings are as follows:

- Reinforcements, regardless of their nature and the adopted lay-out diagram (provided that reinforcement ratio is sufficient),enable us to extend the structural integrity field of masonrywalls.

- TRC reinforcements lead to lower performance levels than FRPreinforcements in terms of lateral strength capacity, but theysignificantly increase their ductility capacity.

- TRC and GFRP seem to be more appropriate than CFRP in termsof ultimate displacement capacity.

- A low TRC reinforcement ratio only marginally modifies globalmasonry wall performances.

- Dissipation mechanisms, which differ between FRP andTRC-reinforced walls, have been clarified; in particular, theydiffuse dissipation processes (micro-cracks) in TRCcomposites.

- Anchorage systems are appropriate (and technologically possi-ble) to improve the in-plane performances of reinforcedmasonry walls.

- Reinforcement design is limited by the compressive strength ofconcrete hollow blocks.

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- The rigid body motion of a strengthened wall depends on thereinforcement pattern (reinforcement ratio, axial rigidity andreinforced surface area) and is likely to substantially limit thecontributions of anchorage systems.

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