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Shear behaviour of masonry walls strengthened by external bonded FRP and TRC
Composite Structures 132 (2015) 923–932
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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 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 b
a r t i c l e
i n f o
Article history: Available online 2 July 2015 Keywords: Masonry wall Strengthening FRP TRC Shear behaviour
a b s t r a c t 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 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 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 reinforced concrete and steel prevail in the modern structures, masonry units are also used. In France, a significant part of buildings 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 structures – a complex behaviour. Therefore, it is obvious that we should pay attention to these structures in a seismic context, particularly when a seismic hazard assessment has been revised, leading to a tightening of the safety rules in France. Indeed, past earthquakes have revealed that unreinforced masonry structures can suffer extensive damage. Their vulnerability 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]. ⇑ Corresponding author. Tel.: +33 4 77 43 75 38; fax: +33 4 78 43 33 83. E-mail address: [email protected] (A. Si Larbi). http://dx.doi.org/10.1016/j.compstruct.2015.06.057 0263-8223/Ó 2015 Elsevier Ltd. All rights reserved.
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 fibrereinforced polymer (FRP) composites are often preferentially chosen 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
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of the debonding of the composite sheets. To address this issue, an adequate mechanical anchorage 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 and lateral load resistance of a masonry wall – have now been widely acknowledged in the case of out-of-plane actions [5]. 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 reinforced 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 [6–11]. 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 [8]. Although instructive, these studies lack diversity for studied materials, reinforcement configurations, applied normal loads and slenderness ratio of walls. Sometimes, a small amount of information is known regarding damage and failure mechanisms or regarding the interaction between masonry material and reinforcements. On the one hand, this work is aimed at further developing the existing experimental database, with special emphasis on identifying the performances of anchorage devices, particularly in the framework of a comparative study between FRP and TRC composites. This comparison will cover criteria at the global scale and, to a lesser extent, at the local scale. On the other hand, this paper tries to help identify and clarify damage dissipative mechanisms and their impact 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 assess with more realistic loadings in terms of earthquake hazards.
2. Experimental program 2.1. Masonry walls A series of six walls has been built with the same dimension given in the Fig. 1. It should be mentioned that all of the specimens were built by a professional mason and must be considered to be in compliance with the practices. The hollow concrete block units, whose dimensions are 500 mm long, 200 mm high and 75 mm thick, belong to Group 2 according to Eurocode 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 dimensions compatible with the limited means 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 has been determined and ranges from 4 to 10 MPa (6.5 MPa on average with a standard deviation of 2.33). These blocks are assembled with a mortar composed of Portland cement (CEM I 52.5) and sand in the proportion 1:3 with a water/cement ratio equal to 0.5. Mortar test prisms of 40 40 160 mm3 were tested for compressive and flexural strengths. At 31 days, these strengths are 48 MPa and 10 MPa, respectively. 2.2. Reinforcement 2.2.1. 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). 2.2.1.1. 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 characteristics have been measured on six specimens according to ISO 527-1. The obtained results are listed in Table 1. 2.2.1.2. TRC composite. Knowledge on TRC composites is notably less significant than knowledge relating to FRP. However, it is Table 1 Mechanical characteristics of composites. Composite strengthening system
Nominal thickness (mm)
Young modulus (GPa)
Tensile strength (Mpa)
Ultimate strain (lm/ m)
CFRP GFRP
0.48 1.7
105 7.2
1700 100
16000 13.800
« Concrete loading beam »
« Reinforced concrete foundation »
Fig. 1. Description of unreinforced masonry wall (reference).
T.-L. Bui et al. / Composite Structures 132 (2015) 923–932 Table 2 Mix design for TRC composite. Micro-mortar**
Textile reinforcement
* **
Nature of fibres TEX Fibre diameter Number of filament/yarn Knitted grid size Tensile strength (yarn)
Glass-AR 1200* 19 lm* 1600*
Grain size Silica-fume Thixotropy Shrinkage
<2 mm* Yes* Yes* 0*
5 5 mm* 1102 MPa*
5 MPa** 40 MPa*
Young modulus
74,000 MPa*
Tensile strength Compressive strength Young modulus
1700 MPa*
Provided properties. Laboratory characterisation.
currently well established that various levers (roving diameters, yarn number, impregnation, etc.) can be mobilised to optimise TRC – with occasionally opposite consequences – and many of them 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 components are listed in Table 2, that results from a compromise between all of the aforementioned 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 bidirectional warp-knitted grid fabric. Direct quasi-static tensile tests, whose protocol has been validated [13], have been performed to mechanically characterise the TRC reinforcements. Stress–strain curves are given in Fig. 2. Two different behaviour laws appear in Fig. 2. Indeed, without the impregnating resin (latex), TRC exhibits a bilinear behaviour whereas by using latex, the evolution law is nearly linear because of a more homogeneous yarn contribution. 2.2.2. Anchorages To reduce or ideally remove the overturning effects due to lateral loads on walls and to best maximise the potential of each reinforcement solution by a priori improving its efficiency, a connector, in the form of an anchorage (see Fig. 3), has been introduced between walls (on their lower part, on both sides) and their foundation footings.
Fig. 2. Tensile behaviour of TRC materials during uniaxial tensile tests.
Fig. 3. Description of a MAPEI anchor.
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Given the existing solutions and their well-established performances (easy application and high strength), only connection solutions based on carbon fibres and an epoxy resin have been considered. The anchorage 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 manufacturer is 30 kN at the ultimate limit state. 2.2.3. Strengthening configurations The definition of strengthening patterns must be part of a strategy aimed at finding a «balance» between lateral strength and energy dissipation capacity [14]. Thus, the objective in reinforcing a structure is to improve its strength capacity, to enhance its ductility, or both. According to this strategy, different strengthening configurations with TRC and FRP composites have been proposed. The reinforcing material is always applied symmetrically on both wall surfaces. 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 Fig. 4). 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. Each side of both faces is reinforced by a continuous sheet, is comprised of two glass layers over a width of 400 mm along the entire height of the masonry and is combined with two discontinuous carbon sheets, which are 1410 mm long and 60 mm wide; the horizontal distance between these two carbon sheets is 100 mm. 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, embedding glass fabrics (rather than carbon fabrics) in an epoxy matrix tends to produce composites with better energy dissipation capacities. Thus, it is appropriate to take advantage of this property in the context of reinforcing masonry walls. This choice to use only glass fibres is more appropriate because it adds value to TRC materials (compared with carbon–epoxy composites), which achieves a smaller ecological footprint and has fewer problems with the hygiene and safety conditions for workers. As a consequence, it is clear that without overlooking the loadbearing capacity, it is desirable to increase the ductility of reinforced walls, which is the main motivation of our choices. In the first pattern (TRCW1 wall), only one TRC layer, 1410 mm high and 200 mm wide, 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, which are 1410 mm high and 200 mm wide. 2.2.4. Application of strengthening systems and anchor placement FRP/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 its reinforcements. The application of a strengthening system first involves covering the wall with a layer of epoxy resin (FRP) or mortar (TRC). Then, the fabric (FRP) or textile (TRC) is placed along the wall and is pressed against the resin or mortar. A second layer of resin or mortar is eventually applied to ensure fabric impregnation. For walls reinforced with several layers, the last two steps are repeated as necessary. These reinforcements are anchored to the foundation because of the CFRP structural connections. An anchorage anchor is composed of two parts: the anchorage strictly speaking and the «whip». The
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anchorage that is first impregnated with epoxy-resin is inserted into holes that have been drilled before in the reinforced concrete footing and is filled with resin. The remaining part of the anchor (the whip) is deployed similar to a fan on the wall, and resin is again applied so that the adhesive completely penetrates into the fibres. This splayed part ensures the connection of strengthening materials with the masonry wall. 2.3. 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 actuator with 200 kN capacity. This actuator is combined with a force sensor that is located on a stiff steel plate (100 mm wide 100 mm long 20 mm thick) 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 (Fig. 5). 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. Tomazevic [15] has proposed to adopt a constant value equal to 20% of the compression strength for confined brick masonry. Papanicolaou et al. [9] have worked with large ranges from 2.5% to 10% for rectangular walls and from 10% to 25% for slender walls. This value is between 11% and 16% in F.da Porto’s study [10] regarding steel-reinforced masonry. For a given wall geometry, the axial compression load magnitude influences the overall performances, in particular, the efficiency of strengthening solutions. Papanicolaou et al. [9] have noted that reinforcement 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 corresponds 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 displacement 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 significantly. According to Tomazevic [15], failure is ascertained when horizontal load falls by 20%. This criterion will be adopted herein. In addition to the force sensors that have been presented before in the description of the test set-up, for the sake of clarity, several displacement transducers and strain gauges complete the instrumentation. Their locations and purposes will now be discussed. The displacement transducer C1 (LVDT ± 100 mm) measures the lateral displacement at the top of the wall and will enable us to characterise the overall behaviour of the wall through load–displacement curves. The assumption that the foundation footing has no slips is controlled because of the displacement transducer C3 (LVDT ± 100 mm). To have information at the local scale and to assess the contributions of different reinforcements, strain gauges (120 X), numbered from J1 to J9, are bonded on only one face of the walls (Fig. 4).
walls do not heave. Concerning measures from LVDT C3, they emphasise that the concrete footing relative to the rigid floor does not slide for all six walls. As a consequence, the obtained relationships of lateral loads versus horizontal displacements at the top of the wall can be considered without caution towards boundary conditions. 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. Until the end of this first phase, wall integrity is preserved at the global scale. Next, a nonlinear phase starts, which differs depending on the nature of the walls and reinforcements. This nonlinear zone can be related to the damage (in tension or in shear) of one or more masonry components, the strengthening materials or even the reinforcement/block (or mortar joint) interface (interphase). A dramatic increase in the ultimate strength and in the second phase stiffness is clearly observed, 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 result of a sudden and premature failure) (Fig. 6). To evaluate the performances of these walls in terms of appropriately chosen and unbiased indicators, the experimental load–displacement curves will be idealised according to the trilinear model proposed by Tomazevic [15] (Fig. 7). Given the experimentally observed behaviours, this approach seems to be more consistent than the idealised bilinear relationship proposed by Magenes and Calvi [16]. The three phases of the conventional trilinear diagram are bounded by three characteristic points, which facilitate discussion and comparison on wall behaviours. Thus, energy dissipation, calculated from the idealised diagrams, will be useful to assess the suitability of strengthening solutions with dissipation needs in the context of earthquakes. 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 [3]. As masonry walls exhibit highly nonlinear behaviours, a conventional Vcr value is generally adopted. According to Tomazevic’s study [15], 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 characterised by the ultimate load Vdu, corresponding to 80% of the maximum load and the maximum displacement du on the softening branch. Kel, lu 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 idealised load–displacement diagram and is given by the following equation:
Ediss ¼
1 ½dcr :V cr þ ðdVmax dcr ÞðV max þ V cr Þ þ ðdu dVmax ÞðV max 2 þ V du Þ
ð1Þ
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.
2.4. Experimental results 2.5. Strength capacity 2.4.1. Global behaviour The number of steel rods, their high axial rigidity and their suitable tightening, without overlooking the high flexural rigidity of transverse steel girders, are many arguments for assuming that
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
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URW
TRCRW2
TRCRW1
(One layer without latex)
(3 layers with latex)
CRW
GRW
CGRW
Fig. 4. Unreinforced wall and FRP/TRC-reinforced walls.
Spreader beam Load cell Hydraulic jack
Load cell
Vertical load actuator
Post-tension rod Wall Beams
Reaction wall
Threaded rod
Fig. 5. Test set-up.
from 110% (TRCW1) to 450% (CGRW). TRC reinforcements are slightly less efficient (175% on average) than FRP reinforcements in terms of strength capacity. A more detailed analysis allows a quantitative comparison of the different configurations with TRC and FRP materials and highlights that strength capacity gains grow proportionally to reinforcement axial stiffness (qv :ER ) (The case of the CGRW wall has been eluded because it collapsed when the anchorages prematurely failed) – Fig. 9. This relationship is consistent with Mahmood et al. [17].
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. 2.6. Initial stiffness Initial stiffness is marginally affected by reinforcements. This can stem from the low thickness of strengthening systems, in
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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 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 because increases in the range of 200–500% can be expected. To better understand, it is important to correlate the above performance indicators with observed failure modes, damage mechanisms and kinematics and with data at the local scale. This is the purpose of the rest of this paper.
Fig. 6. Curves of load versus horizontal displacement at the top of the wall.
particular with FRP materials and to a lesser extent with TRC materials. By contrast, the size of the linear zone is strongly subordinated to the adopted reinforcements’ nature and configuration. Indeed, according to experimental results (Vcr indicator seems irrelevant or at least unsuitable for results obtained in the present case and presented as load–displacement curves), overall, wall macroscopic damage is delayed, except for the TRCW1 wall whose reinforcement ratio is low. In the first stage, the discernible increase in initial stiffness may be attributed to the ability of the reinforcement (mainly when its axial stiffness is sufficient) to bridge emerging microcracks and cracks because of a suitable load transfer. 2.7. Ultimate displacement and ductility Furthermore, it is important to note that the ultimate displacement is a minimum (except for the CGRW wall) maintained when compared with the reference wall. This emphasises the interest of chosen reinforcement solutions and a priori attests to the usefulness of anchorages between footings and walls. It is also worth noting that two reinforced walls (TRCW2 and GRW) see a tangible increase, of approximately 36% in their ultimate displacements. Walls are likely subject to flexural mechanisms, which induce normal tensile stresses in reinforcements. Thus, the ultimate displacement capacity of strengthened masonry walls is, among other factors, certainly conditioned by the axial stiffness of reinforcements, which can explain at least partially the increase in stiffness for these reinforced walls. The decrease in ultimate displacement is peculiar to the CGRW wall because anchorages have certainly been involved in the wall failure (see below). As mentioned previously, the ductility factor (l = du/dcr) plays a crucial role in assessing the seismic behaviour of masonry walls in particular because it reduces the number of elastic seismic design actions. Table 3 highlights that TRC-reinforced walls exhibit nearly the same ductility factor as an unreinforced wall. In contrast, for the FRP-reinforced walls case, if lateral strength clearly increases, the ultimate ductility coefficient is low compared with an unreinforced wall (Fig. 8d and Fig. 10). This may be related to the capacity that TRC composites have, unlike FRP composites, to crack, to «follow» masonry wall displacements and potentially to «absorb» wall damage, which thus remains controlled. Needless to say, the above assumptions require a more elaborate experimental campaign to be validated, but some of the proposed explanations will also be reviewed regarding available data at the local scale.
2.8.1. Failure modes This section is devoted to providing information regarding failure modes based on reinforcement types and patterns, so that the efficiency of strengthening materials in the global lateral behaviour and their impact on damage mechanisms can be precise. 2.8.1.1. Unreinforced wall. An unreinforced wall exhibits a flexural failure mode (Fig. 11) characterised by horizontal cracks on the left part of the wall (on the side where lateral load is applied) due to tensile stresses in bed mortar joints and by toe crushing (on the right part) at the compressed corner. To better understand the failure process, the order in which damage has been visually detected is as follows. Horizontal cracks start to open on the left side probably because the tensile bond strength is exceeded and gradually develops as the crack opening grows. These flexural cracks involve the first and second rows of mortar bed joints. It is not excluded that high shear stresses at the left bottom corner have contributed to the initial failure of the first line of bed joints first. The increasing lateral displacements therefore induce the spreading of cracks through the second row of bed mortar joints. At this point, the wall tends to rotate about the right bottom corner, thereby inducing the horizontal spread of cracks (secondarily, vertical joints between the first and second rows are damaged). As the resistant cross section is reduced, the right bottom corner is compressed. It eventually fails under a complex stress state as evidenced by concentrated crushings together with cracks on an approximately 45-degree angle. It is worth noting that observation of this failure mode has given valuable information for designing strengthening configurations in the present study. More specifically, the objective is to counteract or at least defer damage of the horizontal joints that are subjected to predominantly tensile stresses because of vertical strips at the ends of walls
Fig. 7. Idealised tri-linear diagram.
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URW CGRW CRW GRW TRCRW1 TRCRW2
– CFRP GFRP CFRP GFRP TRC TRC
t R bR (mm2)
AR qv ¼ lt
qv :ER (MPa)
Vcr = (70%Vmax) (kN)
dcr (mm)
Vmax (kN)
dVmax (mm)
Vu = (80%Vmax) (kN)
du (mm)
– 0.48 60 1.7 400 0.48 60 1.7 400 3 200 9 200
– 3.82
82,020
7.88 35.79
0.88 3.66
11.25 51.13
4.59 8.39
9.00 40.90
12.98 8.70
0.3 3.52 3.31 13.98
31,316 50,702 4854 22,135
22.40 35.53 8.61 18.92
2.50 3.06 0.75 1.07
32.00 50.75 12.30 27.03
9.94 8.89 5.85 7.32
25.60 40.60 9.84 21.62
13.20 17.70 12.80 17.60
l = du/ dcr
Ediss (kN mm)
Failure modes
8.97 9.78
14.78 2.38
123.90 285.30
Flexural Flexural + shear
8.96 11.61 11.48 17.68
5.28 5.78 17.07 16.45
324.26 708.24 133.49 403.80
Flexural + shear Flexural + shear Flexural + shear Flexural + shear
Kel (kN/ mm)
P tR - thickness of composite band; bR - width of composite band; AR - total cross section area of strengthening (= bR :tR ); l and t- width and thickness of masonry wall; qv vertical reinforcement ratio.
Fig. 8. Comparative diagrams of the different indicators (a-strength capacity, b-stiffness, c-ultimate displacement, d-ductility coefficient, e-dissipation energy).
unreinforced zones are maintained as concentrated damage zones in which energy dissipation can advantageously take place. Of course, wall behaviour – in terms of lateral strength or energy dissipation capacity – highly depends on the strength of anchorages.
Fig. 9. Relationship between strength capacity and reinforcement axial stiffness (qv :ER ).
(areas where tensile stresses mainly develop during an earthquake loading cycle). Moreover, the use of wide strips (or even central strips) aims at covering a large surface to bridge cracks, while at the same time,
2.8.1.2. FRP-reinforced walls. Failure modes of FRP-reinforced walls are presented in Fig. 12. First, it must be noted that they are different and depend, among other things, on adopted reinforcement configurations. Walls predominantly exhibit shear failures with different crack patterns that also correlate with the damage nature – either brittle failure or progressive degradation – which has major implications on the dissipation capabilities of walls. 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
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for the CRW wall (which is a typical shear failure pattern), even along the edge of the unreinforced zone (in particular for GRW wall), thus reflecting that reinforcements can bridge cracks and also influence their propagation. Nevertheless, 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 (Fig. 12(b) and (c)). Fig. 10. Reassessment of ultimate ductility factors by using the experimental elastic limit rather than the conventional one.
Crushing of the compressive brick
Cracks of mortar joints
Fig. 11. Flexural failure for unreinforced wall (URW).
the bottom of the wall which could explain, even partially, the observed failure mode. As a consequence, 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 materials). 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 propagate along the compressed diagonal
2.8.1.3. 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 (Fig. 13). Both reinforcement patterns lead to cracks at the left bottom 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 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 «peeling-off» failure [18] (Fig. 13a). 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 limited (Fig. 13b). In light of the obtained results, it is important to emphasise that: - Anchors can be considered 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 reinforced walls, regardless of the adopted reinforcements’ type and pattern. However, the dissipation process is controlled because it occurs on the reinforcement itself for TRC materials, whereas it is uncontrolled with FRP-debonding.
Crushing of the compressive brick
Crack of FRP Local debonding s’ propa gation Crack Crushing of the s’ propa compressive brick gation
Failure of anchors
(a)
(b)
(c)
Fig. 12. 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.
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Macro-crack of TRC
Multi-cracks of TRC
(a)
(b)
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).
2.8.2. Local behaviour Only strains measured in FRP-reinforced walls are displayed because gauges applied on TRC have failed early, without giving valuable information. Moreover, because of complex local effects (in zones where anchorage cords and FRP composites are held together), only strain gauges (3; 6; 9) that are located above anchorage cords are considered. The evolution of strains in vertical FRP reinforcements along the wall length is given in Fig. 14 for different loading levels ranging from 0 to Vmax, the maximum lateral strength. In general, maximum tensile strains are measured with gauges located at the tensile side of the wall, and strains gradually decrease, quasi linearly, to become negative at the compressed side. These evolutions give credibility to the assumption that plane sections remain plane, despite wall heterogeneity and until lateral loading values are close to failure load. Furthermore, at maximum load, maximum strains in FRP reinforcements reach values of 4000 lm/m in CFRP strips and 9000 lm/m in GFRP strips, respectively, for CRW and GRW walls. These values are very close to the ultimate strain values for glass fibre-reinforced composites and account for only 50% of the ultimate strain value for carbon-based reinforcements. Thus, the usefulness of glass fibres is evident. In contrast, in the case of the hybrid solution with both glass and carbon fibres, strains remain small (2500 lm/m – CGRW wall) because of anchor failure. This limited strain highlights that without efficient anchorage, reinforcements cannot significantly contribute to load transfer. This is the reason why the global performance of this wall is limited.
3. 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 performance levels than FRP reinforcements in terms of lateral strength capacity, but they significantly 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 diffuse dissipation processes (micro-cracks) in TRC composites. - Anchorage systems are appropriate (and technologically possible) to improve the in-plane performances of reinforced masonry walls. - Reinforcement design is limited by the compressive strength of concrete hollow blocks.
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- 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.
References [1] Fallahi A, Alaghebadian R, Miyajima M. Microtremor measurements and building damage during the Changureh-Avaj, Iran earthquake of June 2002. J Nat Disaster Sci 2003;25(3):37–46. [2] Klingner R. ‘‘Behavior of masonry in the Northridge (US) and Tecoman-Colima (Mexico) earthquakes: lessons learned, and changes in design provisions’’. 6_Congresso Nacional de Sismologia e Engenharia Sismica, Guimarães, Portugal; 2004. [3] Capozucca R. Experimental analysis of historic masonry walls reinforced by CFRP under in-plane cyclic loading. Compos Struct 2011;94:277–89. [4] Lanivschi CE. ‘‘State of the art for strengthening masonry with Fibre Reinforced Polymers’’. Buletinul Institutului Politehnic Din IASI Publicat de Universitatea Tehnica˘, Gheorghe Asachi’’ din Iasßi Tomul LVIII (LXII), Fasc. 2, 2012 Sectßia ConstrucTII. Arhitechtura. [5] Lunn Dillon S, Rizkalla SH, Maeda S, Ueda T. ‘‘FRP anchorage systems for infill masonry structures’’; APFIS 2012 Hokkaido University Japan, 2–4 february 2012. [6] Bernat E, Gil L, Roca P, Escrig C. Experimental and analytical study of TRM strengthened brickwork walls under eccentric compressive loading. Constr Build Mater 2013;44:35–47. [7] Carozzi FG, Milani G, Poggi C. Mechanical properties and numerical modeling of fabric reinforced cementitious matrix (FRCM) systems for strengthening of masonry structures. Compos Struct 2014;107:711–25.
[8] Papanicolaou CG, Triantafillou TC, Karlos K, Papathanasiou M. Textilereinforced mortar (TRM) versus FRP as strengthening material of URM walls: in-plane cyclic loading. Mater Struct 2007;40:1081–97. [9] Papanicolaou CG, Triantafillou T, Lekka M. Externally bonded grids as strengthening and seismic retrofitting materials of masonry panels. Constr Build Mater 2011;25:504–14. [10] da Porto Franca, Mosele Flavio. Claudio Modena ‘‘in-plane cyclic behaviour of a new reinforced masonry system: experimental results’’. Eng Struct 2011;33:2584–96. [11] Prota A, Marcari G, Fabbrocino G, Manfredi G, Aldea C. Experimental in-plane behavior of tuff masonry strengthened with cementitious matrix-grid composites. J Compos Constr ASCE 2006;10(3):223–33. [12] Contamine R, Si Larbi A, Hamelin P. Contribution to direct tensile testing of textile reinforced concrete (TRC) composites. Mater Sci Eng: A 2011;528:8589–98. [13] Contamine R, Si Larbi A, Hamelin P. Identifying the contributing mechanisms of textile reinforced concrete (TRC) in the case of shear repairing damaged and reinforced concrete beams. Eng Struct 2013;46:447–58. [14] Davidovici: ‘‘Risque sismique: renforcer des bâtiments existants»; ENIT October 2008. [15] Tomazevic M, Lutman M, Petkovic L. Seismic behavior of masonry walls: experimental simulation. J Struct Eng ASCE 1996;122(9):1040–7. [16] Magenes G, Calvi GM. In-plane seismic response of brick masonry walls. Earthquake Eng Struct Dyn 1997;26:1091–112. [17] Mahmood H, Ingham JM. Diagonal compression testing of FRP-retrofitted unreinforced clay brick masonry wallettes. J Compos Constr ASCE 2011;15(5):810–20. [18] Banholzer B. Bond behaviour of a multi-filament yarn embedded in a cementious matrix [Philosophy doctoral thesis]. Schriftenreihe Aachener Beiträge zur Bauforschung. Institut für Bauforschung der RWTH Aachen 2004; 12.
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Composite Structures journal homepage: www.elsevier.com/locate/compstruct
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 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 b
a r t i c l e
i n f o
Article history: Available online 2 July 2015 Keywords: Masonry wall Strengthening FRP TRC Shear behaviour
a b s t r a c t 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 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 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 reinforced concrete and steel prevail in the modern structures, masonry units are also used. In France, a significant part of buildings 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 structures – a complex behaviour. Therefore, it is obvious that we should pay attention to these structures in a seismic context, particularly when a seismic hazard assessment has been revised, leading to a tightening of the safety rules in France. Indeed, past earthquakes have revealed that unreinforced masonry structures can suffer extensive damage. Their vulnerability 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]. ⇑ Corresponding author. Tel.: +33 4 77 43 75 38; fax: +33 4 78 43 33 83. E-mail address: [email protected] (A. Si Larbi). http://dx.doi.org/10.1016/j.compstruct.2015.06.057 0263-8223/Ó 2015 Elsevier Ltd. All rights reserved.
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 fibrereinforced polymer (FRP) composites are often preferentially chosen 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
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of the debonding of the composite sheets. To address this issue, an adequate mechanical anchorage 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 and lateral load resistance of a masonry wall – have now been widely acknowledged in the case of out-of-plane actions [5]. 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 reinforced 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 [6–11]. 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 [8]. Although instructive, these studies lack diversity for studied materials, reinforcement configurations, applied normal loads and slenderness ratio of walls. Sometimes, a small amount of information is known regarding damage and failure mechanisms or regarding the interaction between masonry material and reinforcements. On the one hand, this work is aimed at further developing the existing experimental database, with special emphasis on identifying the performances of anchorage devices, particularly in the framework of a comparative study between FRP and TRC composites. This comparison will cover criteria at the global scale and, to a lesser extent, at the local scale. On the other hand, this paper tries to help identify and clarify damage dissipative mechanisms and their impact 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 assess with more realistic loadings in terms of earthquake hazards.
2. Experimental program 2.1. Masonry walls A series of six walls has been built with the same dimension given in the Fig. 1. It should be mentioned that all of the specimens were built by a professional mason and must be considered to be in compliance with the practices. The hollow concrete block units, whose dimensions are 500 mm long, 200 mm high and 75 mm thick, belong to Group 2 according to Eurocode 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 dimensions compatible with the limited means 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 has been determined and ranges from 4 to 10 MPa (6.5 MPa on average with a standard deviation of 2.33). These blocks are assembled with a mortar composed of Portland cement (CEM I 52.5) and sand in the proportion 1:3 with a water/cement ratio equal to 0.5. Mortar test prisms of 40 40 160 mm3 were tested for compressive and flexural strengths. At 31 days, these strengths are 48 MPa and 10 MPa, respectively. 2.2. Reinforcement 2.2.1. 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). 2.2.1.1. 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 characteristics have been measured on six specimens according to ISO 527-1. The obtained results are listed in Table 1. 2.2.1.2. TRC composite. Knowledge on TRC composites is notably less significant than knowledge relating to FRP. However, it is Table 1 Mechanical characteristics of composites. Composite strengthening system
Nominal thickness (mm)
Young modulus (GPa)
Tensile strength (Mpa)
Ultimate strain (lm/ m)
CFRP GFRP
0.48 1.7
105 7.2
1700 100
16000 13.800
« Concrete loading beam »
« Reinforced concrete foundation »
Fig. 1. Description of unreinforced masonry wall (reference).
T.-L. Bui et al. / Composite Structures 132 (2015) 923–932 Table 2 Mix design for TRC composite. Micro-mortar**
Textile reinforcement
* **
Nature of fibres TEX Fibre diameter Number of filament/yarn Knitted grid size Tensile strength (yarn)
Glass-AR 1200* 19 lm* 1600*
Grain size Silica-fume Thixotropy Shrinkage
<2 mm* Yes* Yes* 0*
5 5 mm* 1102 MPa*
5 MPa** 40 MPa*
Young modulus
74,000 MPa*
Tensile strength Compressive strength Young modulus
1700 MPa*
Provided properties. Laboratory characterisation.
currently well established that various levers (roving diameters, yarn number, impregnation, etc.) can be mobilised to optimise TRC – with occasionally opposite consequences – and many of them 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 components are listed in Table 2, that results from a compromise between all of the aforementioned 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 bidirectional warp-knitted grid fabric. Direct quasi-static tensile tests, whose protocol has been validated [13], have been performed to mechanically characterise the TRC reinforcements. Stress–strain curves are given in Fig. 2. Two different behaviour laws appear in Fig. 2. Indeed, without the impregnating resin (latex), TRC exhibits a bilinear behaviour whereas by using latex, the evolution law is nearly linear because of a more homogeneous yarn contribution. 2.2.2. Anchorages To reduce or ideally remove the overturning effects due to lateral loads on walls and to best maximise the potential of each reinforcement solution by a priori improving its efficiency, a connector, in the form of an anchorage (see Fig. 3), has been introduced between walls (on their lower part, on both sides) and their foundation footings.
Fig. 2. Tensile behaviour of TRC materials during uniaxial tensile tests.
Fig. 3. Description of a MAPEI anchor.
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Given the existing solutions and their well-established performances (easy application and high strength), only connection solutions based on carbon fibres and an epoxy resin have been considered. The anchorage 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 manufacturer is 30 kN at the ultimate limit state. 2.2.3. Strengthening configurations The definition of strengthening patterns must be part of a strategy aimed at finding a «balance» between lateral strength and energy dissipation capacity [14]. Thus, the objective in reinforcing a structure is to improve its strength capacity, to enhance its ductility, or both. According to this strategy, different strengthening configurations with TRC and FRP composites have been proposed. The reinforcing material is always applied symmetrically on both wall surfaces. 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 Fig. 4). 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. Each side of both faces is reinforced by a continuous sheet, is comprised of two glass layers over a width of 400 mm along the entire height of the masonry and is combined with two discontinuous carbon sheets, which are 1410 mm long and 60 mm wide; the horizontal distance between these two carbon sheets is 100 mm. 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, embedding glass fabrics (rather than carbon fabrics) in an epoxy matrix tends to produce composites with better energy dissipation capacities. Thus, it is appropriate to take advantage of this property in the context of reinforcing masonry walls. This choice to use only glass fibres is more appropriate because it adds value to TRC materials (compared with carbon–epoxy composites), which achieves a smaller ecological footprint and has fewer problems with the hygiene and safety conditions for workers. As a consequence, it is clear that without overlooking the loadbearing capacity, it is desirable to increase the ductility of reinforced walls, which is the main motivation of our choices. In the first pattern (TRCW1 wall), only one TRC layer, 1410 mm high and 200 mm wide, 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, which are 1410 mm high and 200 mm wide. 2.2.4. Application of strengthening systems and anchor placement FRP/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 its reinforcements. The application of a strengthening system first involves covering the wall with a layer of epoxy resin (FRP) or mortar (TRC). Then, the fabric (FRP) or textile (TRC) is placed along the wall and is pressed against the resin or mortar. A second layer of resin or mortar is eventually applied to ensure fabric impregnation. For walls reinforced with several layers, the last two steps are repeated as necessary. These reinforcements are anchored to the foundation because of the CFRP structural connections. An anchorage anchor is composed of two parts: the anchorage strictly speaking and the «whip». The
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anchorage that is first impregnated with epoxy-resin is inserted into holes that have been drilled before in the reinforced concrete footing and is filled with resin. The remaining part of the anchor (the whip) is deployed similar to a fan on the wall, and resin is again applied so that the adhesive completely penetrates into the fibres. This splayed part ensures the connection of strengthening materials with the masonry wall. 2.3. 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 actuator with 200 kN capacity. This actuator is combined with a force sensor that is located on a stiff steel plate (100 mm wide 100 mm long 20 mm thick) 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 (Fig. 5). 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. Tomazevic [15] has proposed to adopt a constant value equal to 20% of the compression strength for confined brick masonry. Papanicolaou et al. [9] have worked with large ranges from 2.5% to 10% for rectangular walls and from 10% to 25% for slender walls. This value is between 11% and 16% in F.da Porto’s study [10] regarding steel-reinforced masonry. For a given wall geometry, the axial compression load magnitude influences the overall performances, in particular, the efficiency of strengthening solutions. Papanicolaou et al. [9] have noted that reinforcement 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 corresponds 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 displacement 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 significantly. According to Tomazevic [15], failure is ascertained when horizontal load falls by 20%. This criterion will be adopted herein. In addition to the force sensors that have been presented before in the description of the test set-up, for the sake of clarity, several displacement transducers and strain gauges complete the instrumentation. Their locations and purposes will now be discussed. The displacement transducer C1 (LVDT ± 100 mm) measures the lateral displacement at the top of the wall and will enable us to characterise the overall behaviour of the wall through load–displacement curves. The assumption that the foundation footing has no slips is controlled because of the displacement transducer C3 (LVDT ± 100 mm). To have information at the local scale and to assess the contributions of different reinforcements, strain gauges (120 X), numbered from J1 to J9, are bonded on only one face of the walls (Fig. 4).
walls do not heave. Concerning measures from LVDT C3, they emphasise that the concrete footing relative to the rigid floor does not slide for all six walls. As a consequence, the obtained relationships of lateral loads versus horizontal displacements at the top of the wall can be considered without caution towards boundary conditions. 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. Until the end of this first phase, wall integrity is preserved at the global scale. Next, a nonlinear phase starts, which differs depending on the nature of the walls and reinforcements. This nonlinear zone can be related to the damage (in tension or in shear) of one or more masonry components, the strengthening materials or even the reinforcement/block (or mortar joint) interface (interphase). A dramatic increase in the ultimate strength and in the second phase stiffness is clearly observed, 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 result of a sudden and premature failure) (Fig. 6). To evaluate the performances of these walls in terms of appropriately chosen and unbiased indicators, the experimental load–displacement curves will be idealised according to the trilinear model proposed by Tomazevic [15] (Fig. 7). Given the experimentally observed behaviours, this approach seems to be more consistent than the idealised bilinear relationship proposed by Magenes and Calvi [16]. The three phases of the conventional trilinear diagram are bounded by three characteristic points, which facilitate discussion and comparison on wall behaviours. Thus, energy dissipation, calculated from the idealised diagrams, will be useful to assess the suitability of strengthening solutions with dissipation needs in the context of earthquakes. 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 [3]. As masonry walls exhibit highly nonlinear behaviours, a conventional Vcr value is generally adopted. According to Tomazevic’s study [15], 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 characterised by the ultimate load Vdu, corresponding to 80% of the maximum load and the maximum displacement du on the softening branch. Kel, lu 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 idealised load–displacement diagram and is given by the following equation:
Ediss ¼
1 ½dcr :V cr þ ðdVmax dcr ÞðV max þ V cr Þ þ ðdu dVmax ÞðV max 2 þ V du Þ
ð1Þ
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.
2.4. Experimental results 2.5. Strength capacity 2.4.1. Global behaviour The number of steel rods, their high axial rigidity and their suitable tightening, without overlooking the high flexural rigidity of transverse steel girders, are many arguments for assuming that
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
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URW
TRCRW2
TRCRW1
(One layer without latex)
(3 layers with latex)
CRW
GRW
CGRW
Fig. 4. Unreinforced wall and FRP/TRC-reinforced walls.
Spreader beam Load cell Hydraulic jack
Load cell
Vertical load actuator
Post-tension rod Wall Beams
Reaction wall
Threaded rod
Fig. 5. Test set-up.
from 110% (TRCW1) to 450% (CGRW). TRC reinforcements are slightly less efficient (175% on average) than FRP reinforcements in terms of strength capacity. A more detailed analysis allows a quantitative comparison of the different configurations with TRC and FRP materials and highlights that strength capacity gains grow proportionally to reinforcement axial stiffness (qv :ER ) (The case of the CGRW wall has been eluded because it collapsed when the anchorages prematurely failed) – Fig. 9. This relationship is consistent with Mahmood et al. [17].
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. 2.6. Initial stiffness Initial stiffness is marginally affected by reinforcements. This can stem from the low thickness of strengthening systems, in
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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 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 because increases in the range of 200–500% can be expected. To better understand, it is important to correlate the above performance indicators with observed failure modes, damage mechanisms and kinematics and with data at the local scale. This is the purpose of the rest of this paper.
Fig. 6. Curves of load versus horizontal displacement at the top of the wall.
particular with FRP materials and to a lesser extent with TRC materials. By contrast, the size of the linear zone is strongly subordinated to the adopted reinforcements’ nature and configuration. Indeed, according to experimental results (Vcr indicator seems irrelevant or at least unsuitable for results obtained in the present case and presented as load–displacement curves), overall, wall macroscopic damage is delayed, except for the TRCW1 wall whose reinforcement ratio is low. In the first stage, the discernible increase in initial stiffness may be attributed to the ability of the reinforcement (mainly when its axial stiffness is sufficient) to bridge emerging microcracks and cracks because of a suitable load transfer. 2.7. Ultimate displacement and ductility Furthermore, it is important to note that the ultimate displacement is a minimum (except for the CGRW wall) maintained when compared with the reference wall. This emphasises the interest of chosen reinforcement solutions and a priori attests to the usefulness of anchorages between footings and walls. It is also worth noting that two reinforced walls (TRCW2 and GRW) see a tangible increase, of approximately 36% in their ultimate displacements. Walls are likely subject to flexural mechanisms, which induce normal tensile stresses in reinforcements. Thus, the ultimate displacement capacity of strengthened masonry walls is, among other factors, certainly conditioned by the axial stiffness of reinforcements, which can explain at least partially the increase in stiffness for these reinforced walls. The decrease in ultimate displacement is peculiar to the CGRW wall because anchorages have certainly been involved in the wall failure (see below). As mentioned previously, the ductility factor (l = du/dcr) plays a crucial role in assessing the seismic behaviour of masonry walls in particular because it reduces the number of elastic seismic design actions. Table 3 highlights that TRC-reinforced walls exhibit nearly the same ductility factor as an unreinforced wall. In contrast, for the FRP-reinforced walls case, if lateral strength clearly increases, the ultimate ductility coefficient is low compared with an unreinforced wall (Fig. 8d and Fig. 10). This may be related to the capacity that TRC composites have, unlike FRP composites, to crack, to «follow» masonry wall displacements and potentially to «absorb» wall damage, which thus remains controlled. Needless to say, the above assumptions require a more elaborate experimental campaign to be validated, but some of the proposed explanations will also be reviewed regarding available data at the local scale.
2.8.1. Failure modes This section is devoted to providing information regarding failure modes based on reinforcement types and patterns, so that the efficiency of strengthening materials in the global lateral behaviour and their impact on damage mechanisms can be precise. 2.8.1.1. Unreinforced wall. An unreinforced wall exhibits a flexural failure mode (Fig. 11) characterised by horizontal cracks on the left part of the wall (on the side where lateral load is applied) due to tensile stresses in bed mortar joints and by toe crushing (on the right part) at the compressed corner. To better understand the failure process, the order in which damage has been visually detected is as follows. Horizontal cracks start to open on the left side probably because the tensile bond strength is exceeded and gradually develops as the crack opening grows. These flexural cracks involve the first and second rows of mortar bed joints. It is not excluded that high shear stresses at the left bottom corner have contributed to the initial failure of the first line of bed joints first. The increasing lateral displacements therefore induce the spreading of cracks through the second row of bed mortar joints. At this point, the wall tends to rotate about the right bottom corner, thereby inducing the horizontal spread of cracks (secondarily, vertical joints between the first and second rows are damaged). As the resistant cross section is reduced, the right bottom corner is compressed. It eventually fails under a complex stress state as evidenced by concentrated crushings together with cracks on an approximately 45-degree angle. It is worth noting that observation of this failure mode has given valuable information for designing strengthening configurations in the present study. More specifically, the objective is to counteract or at least defer damage of the horizontal joints that are subjected to predominantly tensile stresses because of vertical strips at the ends of walls
Fig. 7. Idealised tri-linear diagram.
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URW CGRW CRW GRW TRCRW1 TRCRW2
– CFRP GFRP CFRP GFRP TRC TRC
t R bR (mm2)
AR qv ¼ lt
qv :ER (MPa)
Vcr = (70%Vmax) (kN)
dcr (mm)
Vmax (kN)
dVmax (mm)
Vu = (80%Vmax) (kN)
du (mm)
– 0.48 60 1.7 400 0.48 60 1.7 400 3 200 9 200
– 3.82
82,020
7.88 35.79
0.88 3.66
11.25 51.13
4.59 8.39
9.00 40.90
12.98 8.70
0.3 3.52 3.31 13.98
31,316 50,702 4854 22,135
22.40 35.53 8.61 18.92
2.50 3.06 0.75 1.07
32.00 50.75 12.30 27.03
9.94 8.89 5.85 7.32
25.60 40.60 9.84 21.62
13.20 17.70 12.80 17.60
l = du/ dcr
Ediss (kN mm)
Failure modes
8.97 9.78
14.78 2.38
123.90 285.30
Flexural Flexural + shear
8.96 11.61 11.48 17.68
5.28 5.78 17.07 16.45
324.26 708.24 133.49 403.80
Flexural + shear Flexural + shear Flexural + shear Flexural + shear
Kel (kN/ mm)
P tR - thickness of composite band; bR - width of composite band; AR - total cross section area of strengthening (= bR :tR ); l and t- width and thickness of masonry wall; qv vertical reinforcement ratio.
Fig. 8. Comparative diagrams of the different indicators (a-strength capacity, b-stiffness, c-ultimate displacement, d-ductility coefficient, e-dissipation energy).
unreinforced zones are maintained as concentrated damage zones in which energy dissipation can advantageously take place. Of course, wall behaviour – in terms of lateral strength or energy dissipation capacity – highly depends on the strength of anchorages.
Fig. 9. Relationship between strength capacity and reinforcement axial stiffness (qv :ER ).
(areas where tensile stresses mainly develop during an earthquake loading cycle). Moreover, the use of wide strips (or even central strips) aims at covering a large surface to bridge cracks, while at the same time,
2.8.1.2. FRP-reinforced walls. Failure modes of FRP-reinforced walls are presented in Fig. 12. First, it must be noted that they are different and depend, among other things, on adopted reinforcement configurations. Walls predominantly exhibit shear failures with different crack patterns that also correlate with the damage nature – either brittle failure or progressive degradation – which has major implications on the dissipation capabilities of walls. 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
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for the CRW wall (which is a typical shear failure pattern), even along the edge of the unreinforced zone (in particular for GRW wall), thus reflecting that reinforcements can bridge cracks and also influence their propagation. Nevertheless, 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 (Fig. 12(b) and (c)). Fig. 10. Reassessment of ultimate ductility factors by using the experimental elastic limit rather than the conventional one.
Crushing of the compressive brick
Cracks of mortar joints
Fig. 11. Flexural failure for unreinforced wall (URW).
the bottom of the wall which could explain, even partially, the observed failure mode. As a consequence, 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 materials). 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 propagate along the compressed diagonal
2.8.1.3. 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 (Fig. 13). Both reinforcement patterns lead to cracks at the left bottom 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 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 «peeling-off» failure [18] (Fig. 13a). 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 limited (Fig. 13b). In light of the obtained results, it is important to emphasise that: - Anchors can be considered 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 reinforced walls, regardless of the adopted reinforcements’ type and pattern. However, the dissipation process is controlled because it occurs on the reinforcement itself for TRC materials, whereas it is uncontrolled with FRP-debonding.
Crushing of the compressive brick
Crack of FRP Local debonding s’ propa gation Crack Crushing of the s’ propa compressive brick gation
Failure of anchors
(a)
(b)
(c)
Fig. 12. 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.
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Macro-crack of TRC
Multi-cracks of TRC
(a)
(b)
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).
2.8.2. Local behaviour Only strains measured in FRP-reinforced walls are displayed because gauges applied on TRC have failed early, without giving valuable information. Moreover, because of complex local effects (in zones where anchorage cords and FRP composites are held together), only strain gauges (3; 6; 9) that are located above anchorage cords are considered. The evolution of strains in vertical FRP reinforcements along the wall length is given in Fig. 14 for different loading levels ranging from 0 to Vmax, the maximum lateral strength. In general, maximum tensile strains are measured with gauges located at the tensile side of the wall, and strains gradually decrease, quasi linearly, to become negative at the compressed side. These evolutions give credibility to the assumption that plane sections remain plane, despite wall heterogeneity and until lateral loading values are close to failure load. Furthermore, at maximum load, maximum strains in FRP reinforcements reach values of 4000 lm/m in CFRP strips and 9000 lm/m in GFRP strips, respectively, for CRW and GRW walls. These values are very close to the ultimate strain values for glass fibre-reinforced composites and account for only 50% of the ultimate strain value for carbon-based reinforcements. Thus, the usefulness of glass fibres is evident. In contrast, in the case of the hybrid solution with both glass and carbon fibres, strains remain small (2500 lm/m – CGRW wall) because of anchor failure. This limited strain highlights that without efficient anchorage, reinforcements cannot significantly contribute to load transfer. This is the reason why the global performance of this wall is limited.
3. 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 performance levels than FRP reinforcements in terms of lateral strength capacity, but they significantly 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 diffuse dissipation processes (micro-cracks) in TRC composites. - Anchorage systems are appropriate (and technologically possible) to improve the in-plane performances of reinforced masonry walls. - Reinforcement design is limited by the compressive strength of concrete hollow blocks.
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- 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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