shear behaviour of masonry walls...

11th International Symposium on Ferrocement and Textile Reinforced Concrete 3rd ICTRC

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SHEAR BEHAVIOUR OF MASONRY WALLS STRENGTHENED BY EXTERNAL BONDED FRP AND TRC

L. Bui1, A. Si Larbi2, N. Reboul1, E. Ferrier 1, 1LGCIE, Université Claude Bernard LYON 1, 82 bd Niels Bohr, 69622 Villeurbanne, FRANCE, 2, 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

Abstract: This experimental study focuses on the behaviour of hollow concrete brick ma-sonry 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 submitted 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 eval-uated 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.

INTRODUCTION

Masonry has a long history as a building technique. 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 direc-tions. Although the 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 seis-mic loading.

Solutions for repairing or strengthening masonry structures are many and are varied. Nev-ertheless, externally bonded fibre-reinforced polymer (FRP) composites are often preferen-tially chosen by prime contractors [2], 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 of the debonding of the composite sheets. To address this issue, an adequate mechanical an-chorage system needs to be set up to enhance the bond (between a masonry structure and its foundation) performance. The benefits of this solution – in terms of the FRP efficiency


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and lateral load resistance of a masonry wall – have now been widely acknowledged in the case of out-of-plane actions [3].

In addition, in the context of sustainable development and health and safety conditions for workers, consideration should be given to an alternative material to FRP, which is often manufactured with highly toxic epoxy resins. The idea is to substitute these resins with cementitious materials while preserving or even improving the dissipative capacity of rein-forced structures. From this perspective, textile-reinforced concrete (TRC) composites, which combine a suitable fine-grained mortar with the latest generation of textile fabrics, would benefit from promotion.The efficiency of TRC for strengthening masonry structures has recently been investigated ([4], [5], [6]-[9]). Compared with FRP, TRC composites show a nonlinear tensile behaviour with multiple matrices cracking, giving them a greater deformation capacity, a priori more suitable for seismic reinforcement [6].

Although instructive, these studies lack diversity for studied materials, reinforcement con-figurations, applied normal loads and slenderness ratio of walls. This work therefore, on the one hand, is aimed at further developing the existing experimental database, with spe-cial emphasis on identifying the performances of anchorage devices, particularly in the framework of a comparative study between FRP and TRC composites. On the other hand, this paper tries to help identify and clarify damage dissipative mechanisms and their im-pact on the failure modes of the masonry walls. To attain the aforementioned objectives, an experimental campaign has been performed, based on static monotonic shear tests, which are a simplified way to simulate stress states resulting from earthquakes. Certainly, inertial effects and the inherent cyclical nature of seismic actions are not addressed in the present study. However, this work can be regarded as a first step towards the definition of efficient reinforcement solutions. The approach is to test some strengthening configurations to have relevant and valuable information and to offer prospects that would be appropriate to as-sess with more realistic loadings in terms of earthquake hazards.

EXPERIMENTAL PROGRAM

Masonry walls A series of six walls has been built. The hollow concrete block units, whose dimensions are 500 mm long, 200 mm high and 75 mm thick, belong to Group 2 according to Euro-code 6, with a strength class B40 (characteristic compression strength of 4 MPa). However, these blocks have been halved lengthwise before being assembled to make walls dimen-sions compatible with the limited means of the laboratory in terms of space and actuator capacity (Block work size at reduced scale : 250 x 200 x 75 mm3). These blocks are as-sembled with a mortar composed of Portland cement (CEM I 52.5) and sand in the propor-tion 1:3 with a water/cement ratio equal to 0.5. These walls are built on a reinforced con-crete foundation and are superposed by a concrete loading beam (Figure 3)

Reinforcement a. Strengthening materials

Two types of composites have been used: the first composite is a fibre-reinforced polymer (FRP) while the latter composite is a textile-reinforced cementitious composite (TRC).


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FRP composite

The fibre-reinforced composite materials consist of a two-component epoxy matrix and bi-directional fabrics made of either carbon (CFRP) or glass (GFRP). Their mechanical char-acteristics have been measured on six specimens according to ISO 527-1. The obtained results are listed in Table 1.

Table 1: Mechanical characteristics of composites

Composite strengthening

system

Nominal thick-ness (mm)

Young modulus

(GPa)

Tensile strength

(Mpa)

Ultimate strain

(μm/m)

CFRP 0,48 105 1700 16000

GFRP 1,7 7,2 100 13800

TRC composite

The components of selected composite, referenced to works performed by Contamine et al. [10], are listed in Table 2, that results from a compromise between all of the aforemen-tioned parameters, including workability and thixotropy (to apply strengthening materials more easily). The composite material contains the 4.36 % AR-glass fibre volume fraction, that is, 2.18% in each principal direction. The textile used for reinforcement is a bidirec-tional warp-knitted grid fabric.

Table 2: Mix 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 μm* Thixotropy yes*

Number of filament/yarn 1600* Shrinkage ∼ 0*

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

Tensile strength (yarn) 1102 MPa* Compressive strength 40 MPa*

Young modulus 74000 MPa* Young modulus 1700 MPa*

* Provided properties; ** laboratory characterization


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Direct quasi-static tensile tests, whose protocol has been validated [11], have been per-formed to mechanically characterize the TRC reinforcements. Stress-strain curves are giv-en in Figure 1. Two different behaviour laws appear in Figure 1. Indeed, without the im-pregnating resin (latex), TRC exhibits a bilinear behaviour whereas by using latex, the evo-lution law is nearly linear because of a more homogeneous yarn contribution.

b. Anchorages To reduce or ideally remove the overturning effects due to lateral loads on walls and to best maximize the potential of each reinforcement solution by a priori improving its effi-ciency, a connector, in the form of an anchorage, has been introduced between walls (on their lower part, on both sides) and their foundation footings. Given the existing solutions and their well-established performances (easy application and high strength), only connec-tion solutions based on carbon fibres and an epoxy resin have been considered. The an-chorage system (MAPEI) is an anchor made from monodirectional carbon fibres with at least 36 yarns, each including 12,000 fibres. The anchorage strength given by the manufac-turer is 30 kN at the ultimate limit state.

Figure 2: Description of a MAPEI anchor

c. Strengthening configurations On the objective in reinforcing a structure is to improve its strength capacity, to enhance its ductility, or both, different strengthening configurations with TRC and FRP composites have been proposed. The reinforcing material is always applied symmetrically on both wall surfaces.

Figure 1: Tensile behaviour of TRC materials during uniaxial tensile tests


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The experimental program consists of testing six masonry walls to failure, including an unreinforced wall (as the reference specimen) and five TRC or FRP-reinforced specimens (see Figure 3). With FRP composites, three strengthening patterns have been proposed. The first wall, referenced CGRW, has been reinforced with both carbon fibre-reinforced polymer (CFRP) and glass fibre-reinforced polymer (GFRP) to significantly improve strength capacity. The two remaining walls, for their part, have been reinforced by either CFRP sheets (CRW wall) or GFRP sheets (GRW wall) to combine strength capacity and ductility. Concerning TRC solutions, it is clear that without overlooking the loadbearing capacity, it is desirable to increase the ductility of reinforced walls, which is the main mo-tivation of our choices. In the first pattern (TRCW1 wall), only one TRC layer, is applied on each side of both faces. However, in the second selected pattern, the main objective is to improve the dissipation capacities, so a vertical strip is applied in the middle of the wall. Each strip is made of three TRC-layers.

CGRW CRW GRW

URW TRCRW1 (One layer without latex)

TRCRW2 (3 layers with latex)

« Concrete loading beam »

Figure 3: Unreinforced wall and FRP/TRC-reinforced walls

Testing procedure As mentioned above, monotonic in-plane combined compression-shear tests have been selected. The vertical axial load (N) is centred and applied by means of a hydraulic actua-tor with 200 kN capacity. This actuator is combined with a force sensor that is located on a stiff steel plate and is placed on the concrete loading beam. These devices are held by a steel profile connected to the strong floor through steel rods, which allows us to control the applied vertical load (Figure 4).


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Wall

Threaded rod

Beams

Hydraulic jack

Post-tension rod

Spreader beam

Vertical load actuatorLoad cell

Reaction wall

Load cell

Figure 4: Test set-up

The vertical load is should represent the weight of the upper floors. In literature, this value varies greatly and depends on the masonry compression strength. For a given wall geome-try, the axial compression load magnitude influences the overall performances, in particu-lar, the efficiency of strengthening solutions. Papanicolaou et al. [7] have noted that rein-forcement efficiency is reduced as vertical compression load rises. Because this study is motivated by assessing the shear contribution of the reinforcements, a relatively low value of approximately 6% of the masonry compression strength has been adopted. It corre-sponds to a mean vertical stress of 0.2 MPa (15 kN). In practical terms, the vertical load is very slowly applied up to the target value equal to 15 kN. It is kept constant during the test because of a force control of the vertical actuator. At this time, horizontal load is imposed under quasi-static monotonic conditions. This loading step is performed under displace-ment control at a rate of 0.015 mm/s to better capture post-pic behaviour. Lateral loading is stopped when masonry walls have obviously failed, that is, when lateral force drops signif-icantly. According to Tomazevic [12], failure is ascertained when horizontal load falls by 20%. This criterion will be adopted herein.

Experimental results

a. Global behaviour All walls show a nonlinear behaviour after an initial linear elastic branch. A significant deviation is experienced in the length of the linear zone although stiffnesses are close. A dramatic increase in the ultimate strength and in the second phase stiffness is clearly ob-served, even if it is conditioned by reinforcing materials and strengthening configurations. Similarly, dissipation capacities increase overall in varying degrees that must be assessed and quantified. It must be underlined that only the wall CGRW behaves as brittle (as a re-sult of a sudden and premature failure) (Figure 5a). To evaluate the performances of these walls in terms of appropriately chosen and unbiased indicators, the experimental load-displacement curves will be idealized according to the trilinear model proposed by Toma-zevic [12] (Figure 5b).


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(a) (b)

Figure 5: (a) - Curves of load versus horizontal displacement at the top of the wall; and (b) - Idealized tri-linear diagram

The three phases of the conventional trilinear diagram are bounded by three characteristic points, which facilitate discussion and comparison on wall behaviours. The first zone, called elastic, ends at lateral load Vcr and displacement dcr. They mark the formation of the first significant cracks in the wall, which entail a change in stiffness [1]. As masonry walls exhibit highly nonlinear behaviours, a conventional Vcr value is generally adopted. Accord-ing to Tomazevic’s study [12], Vcr is equal to 70% of the maximum resistance Vmax. The second zone extends to the maximum lateral load Vmax and displacement dVmax.

At last, the ultimate zone is characterized by the ultimate load Vdu, corresponding to 80% of the maximum load and the maximum displacement du on the softening branch. Kel, μu and Ediss are the initial stiffness (the slope of the first phase, considered as elastic), ductility coefficient and dissipation energy, respectively. This last parameter provides information on resistance and deformation capability. Indeed, it is determined as the area below the idealized load-displacement diagram and is given by the following equation:

(1)

The ductility coefficient, obtained by dividing the ultimate horizontal displacement (du) by the elastic displacement (dcr), reflects the deformation capacity of shear walls in the post-elastic zone. This parameter is a decisive criterion for paraseismic construction.


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Table 3: Summary of experimental results

Walls × (mm

2 ) = ×

. (MPa

)

Vcr

= (7

0%V

max

) (k

N)

d cr (m

m)

Vm

ax (k

N)

d Vm

ax (m

m)

Vu =

(80%

Vm

ax)

(kN

)

d u (m

m)

Kel(k

N/m

m)

μ =

d u/d

cr

Edi

ss(k

N.m

m)

Failure modes

URW - - - 7.88 0.88 11.25 4.59 9.00 12.98 8.97 14.78 123.90 Flexural

CGRW CFRP GFRP

0.48x60 1.7x400 3.82 82020 35.79 3.66 51.13 8.39 40.90 8.70 9.78 2.38 285.30 Flexural+

shear

CRW CFRP 0.48x60 0.3 31316 22.40 2.50 32.00 9.94 25.60 13.20 8.96 5.28 324.26 Flexural+ shear

GRW GFRP 1.7x400 3.52 50702 35.53 3.06 50.75 8.89 40.60 17.70 11.61 5.78 708.24 Flexural+ shear

TRCRW1 TRC 3x200 3.31 4854 8.61 0.75 12.30 5.85 9.84 12.80 11.48 17.07 133.49 Flexural+ shear

TRCRW2 TRC 9x200 13.98 22135 18.92 1.07 27.03 7.32 21.62 17.60 17.68 16.45 403.80 Flexural+ shear

-thickness of composite band; - width of composite band; - total cross section area of strengthening

(= ; l and t- width and thickness of masonry wall; - vertical reinforcement ratio;

Strength capacity From these results, it appears that reinforced walls achieve substantially higher ultimate loads than the reference unreinforced wall, regardless of reinforcement type. Shear strength increases from 110% (TRCW1) to 450% (CGRW). TRC reinforcements are slightly less efficient (175% on average) than FRP reinforcements in terms of strength capacity. It must be noted that the TRCW1 wall appears as an exception because its reinforcement has a slight effect on the ultimate load. It can be explained by the excessively low stiffness of this reinforcement, only comprising one TRC layer.

Dissipation energy In addition to the ductility factor, the dissipation energy (Ediss) will help us position, at least in order of magnitude, the potentials of FRP and TRC as masonry reinforcements. It is clear that for the adopted configurations, both with TRC (with the exception of the TRCRW1 wall with a very low reinforcement ratio) and FRP, results are conclusive be-cause increases in the range of 200 to 500% can be expected. To better understand, it is important to correlate the above 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.

(a) (b) (c) Figure 6: Comparative diagrams of the different indicators (a-strength capacity, b-ductility coefficient, c-dissipation energy)


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b. Failure modes This section is devoted to providing information regarding failure modes based on rein-forcement types and patterns, so that the efficiency of strengthening materials in the global lateral behaviour and their impact on damage mechanisms can be precise.

Unreinforced wall: An unreinforced wall exhibits a flexural failure mode (Figure 7) characterized by horizon-tal cracks on the left part of the wall (on the side where lateral load is applied) due to ten-sile stresses in bed mortar joints and by toe crushing (on the right part) at the compressed corner. It is worth noting that observation of this failure mode has given valuable infor-mation for designing strengthening configurations in the present study.

FRP-reinforced walls: Failure modes of FRP-reinforced walls are presented in Figure 8. First, it must be noted that they are different and depend, among other things, on adopted reinforcement configu-rations.

The sudden failure of the CGRW wall is singular and results from the premature failure of anchorages, which are subject to tensile forces in the left part (side where lateral load is applied). This wall has been reinforced over a large surface area on which both glass and carbon textiles have been laid up, resulting in a significant increase in the wall stiffness (compared with an unreinforced wall but also to other strengthened specimens). Under loading, the wall begins a rigid body motion that induces high shear stresses at the bottom of the wall which could explain, even partially, the observed failure mode. As a conse-quence, the use of anchorage devices certainly improves the lateral strength of reinforced walls. However, to avoid sudden ultimate failures of walls and to ensure sufficient energy dissipation capabilities, an anchorage design cannot be decoupled from the stiffness or reinforced walls (taking into account both reinforced surface area and strengthening mate-rials).

The other two FRP-reinforced walls are not affected by anchor failure and show similar damage mechanisms. These walls exhibit coupled shear-flexure failure modes. In both specimens, shear cracks initiate in the middle of the wall (unreinforced zone) and propa-gate along the compressed diagonal for the CRW wall (which is a typical shear failure pat-tern), even along the edge of the unreinforced zone (in particular for GRW wall), thus re-flecting that reinforcements can bridge cracks and also influence their propagation. Never-theless, the collapse of these walls (or very marked deterioration) occurs by the crushing of the lower right corner (highly subject to compression) and ends and by the splitting of the extreme unit block. With the failure of this block, the wall tends to turn about the right toe, as found for an unreinforced wall (Figure 8 (b) and (c)).


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(b1) (b2) Figure 7: (a)- Flexural failure for unreinforced wall (URW) and Failure modes and damage mechanisms for TRC-reinforced walls (b1)- TRCRW1 wall and (b2)- TRCRW2 wall

(a) (b) (c)

Figure 8: Failure modes and damage mechanisms for FRP-reinforced walls: (a)- sudden failure of CGRW wall; (b) and (c)- coupled shear-flexure failure of CRW and GRW walls

TRC-reinforced walls: As indicated above for FRP-reinforced walls, the addition of TRC reinforcements changes the failure mode. TRC-reinforced walls have failed by combined shear-flexure (Figure 7b). Both reinforcement patterns lead to cracks at the left bottom part (side where load is ap-plied) and in strengthening strips (unlike FRP) over horizontal joints. In the case of the TRCRW1 wall (low reinforcement ratio and low reinforced surface area), cracks develop along a horizontal joint and go through a block unit in the centre with an inclined crack that suggests a reaction to a shear solicitation. At the ultimate limit state, a macro crack has been observed that corresponds to the failure of mortar used in TRC without any textile degradation. In scientific literature, this failure mode is commonly referred to as the « peel-ing-off » failure [13] (Figure 7 b1). In the case of the TRCRW2 wall (reinforced by three TRC layers on a larger area), damage and failure mechanisms are different. Cracks grow horizontally and spread over the height of the strengthening strips. Finally, multi-crack initiation contributes, beyond what is observed on the wall, to the global damage of the specimen. It must be noted that until failure, which ultimately occurs by crushing the right bottom corner, crack opening displacement (without having been measured) remains lim-ited (Figure 7 b2).


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In light of the obtained results, it is important to emphasize that: - anchors can be consid-ered useful and efficient systems, but their use must be correlated at this time with the global stiffness of the wall (reinforcements’ nature and surface area) to avoid their sudden and premature failure; - dissipation mechanisms are improved and concentrated for all re-inforced walls, regardless of the adopted reinforcements’ type and pattern. However, the dissipation process is controlled because it occurs on the reinforcement itself for TRC ma-terials, whereas it is uncontrolled with FRP-debonding.

CONCLUSIONS

The present experimental study has focused on masonry walls reinforced by TRC and FRP composites that are subject to monotonic in-plane combined shear-compression tests. The main findings are as follows: - Reinforcements, regardless of their nature and the adopted layout diagram (provided that reinforcement ratio is sufficient), enable us to extend the structural integrity field of masonry walls; - TRC reinforcements lead to lower perfor-mance levels than FRP reinforcements in terms of lateral strength capacity, but they signif-icantly increase their ductility capacity;- TRC and GFRP seem to be more appropriate than CFRP in terms of ultimate displacement capacity; - A low TRC reinforcement ratio only marginally modifies global masonry wall performances; - Dissipation mechanisms, which differ between FRP and TRC-reinforced walls, have been clarified; in particular, they dif-fuse dissipation processes (micro-cracks) in TRC composites; - Anchorage systems are appropriate (and technologically possible) to improve the in-plane performances of rein-forced masonry walls ;- Reinforcement design is limited by the compressive strength of concrete hollow blocks; - The rigid body motion of a strengthened wall depends on the reinforcement pattern (reinforcement ratio, axial rigidity and reinforced surface area) and is likely to substantially limit the contributions of anchorage systems.

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