Out-of-Plane Strengthening of Masonry-Infilled RC Frames with Textile-Reinforced Mortar Jackets
Publication: Journal of Composites for Construction
Volume 23, Issue 1
Abstract
This paper presents an experimental investigation on the use of textile-reinforced mortar (TRM) jackets as a means of improving the out-of-plane performance of masonry infill walls in reinforced concrete (RC) frames during natural or humanmade catastrophic events (e.g., blasts, earthquakes). The experimental program included testing of six half-scale, single-story masonry-infilled RC frames. All specimens were subjected to out-of-plane monotonic loading with the load being distributed at four points on the infill’s body. Two specimens were tested without receiving any retrofitting, thus serving as reference specimens; one with single-wythe and one with double-wythe infill wall. The remaining four specimens were first retrofitted with carbon-fiber TRM jackets and then tested to failure. The examined parameters included: (1) the connection configuration between the masonry infill wall and the surrounding RC frame members and (2) the thickness of the wall. It was found that the out-of-plane performance was dramatically improved in all cases of retrofitted walls, with the examined parameters playing a crucial role in certain aspects of the wall’s behavior. Overall, the risk of collapse was drastically mitigated when the walls were connected to the RC frame, thus enhancing the resilience of masonry-infilled RC buildings subjected to humanmade or natural catastrophic events.
Introduction and Background
The issue of upgrading unreinforced masonry (URM) walls serving as infills in existing reinforced concrete (RC) buildings is of great importance due to (1) their vulnerability to humanmade hazards such as acts of terrorism and (2) their poor performance under extreme natural events, such as strong earthquakes or even tornados. Acts of terrorism usually involve blasts next to buildings, imposing extreme loads on the building facades, which typically comprise URM infill walls and are extremely vulnerable to such a type of loading. Earthquakes can frequently result in out-of-plane collapse of masonry infills (e.g., Vicente et al. 2012), possibly triggering the progressive collapse of the building. Several out-of-plane collapses of external masonry infill walls have been attributed to weak connections between the interior and exterior walls (i.e., Verderame et al. 2009); sometimes the connections were absent and the brick walls possessed high slenderness. Finally, when tornados strike in a residential area, the building facades in the perimeter are exposed to high differential pressure, which ultimately leads to partial or global collapse of the wall facades.
Particularly in the case of moderate or severe seismic actions, although the presence of masonry infills can be beneficial for the overall performance of a building, they tend to crack early and detach from the surrounding RC frame members either due to in-plane or out-of-plane action (or a combination of the two). Collapse of the damaged infill wall can then be triggered by forces acting in the out-of-plane direction of the wall. The impact of such behavior of masonry infill walls is socioeconomic because it results in huge economic losses and casualties. Thus, an effective strengthening solution of masonry infilled RC frames can improve the resilience of such structures.
There are many studies in the literature investigating various techniques and materials for the in-plane seismic retrofitting of masonry-infilled RC frames. Conventional techniques include RC jacketing (i.e., Pinto et al. 2002), whereas more modern studies include the application of thin layers of lightweight epoxy-based composite materials, such as fiber-reinforced polymers (FRPs) (e.g., Ozcebe et al. 2003; Almusallam and Al-Salloum 2007; Ozden et al. 2011). Compared to the RC jacketing technique, the application of thin layers including noncorrosive lightweight materials yields important benefits, such as increased durability, reduced mass, easier and quicker application, and small thickness increase, which is important for architectural reasons. Moreover, the RC jacketing technique involves more labor, higher quantities of materials, and, most importantly, will interrupt the function of the building under renovation.
Recently, Koutas et al. (2014, 2015) proposed the use of textile-reinforced mortar (TRM) jackets for the in-plane strengthening of masonry-infilled RC frames, with successful results. TRM comprises open-mesh high-strength textiles (i.e., carbon, glass, basalt) combined with inorganic mortars (lime or cement based), resulting in composite materials with many advantages over FRP systems [i.e., resistance to high temperature (Tetta and Bournas 2016; Raoof and Bournas 2017a, b), compatibility with masonry or concrete substrates, applicability at low temperatures or on wet surfaces, lower costs]. Based on the current state-of-the-art studies, TRM has been proven effective for strengthening both concrete and masonry structures (i.e., Triantafillou and Papanicolaou 2006; Bournas et al. 2007; Papanicolaou et al. 2008; Augenti et al. 2011; D’Ambrisi and Focacci 2011; Elsanadedy et al. 2013; De Felice et al. 2014; Loreto et al. 2015; El-Maaddawy and El Refai 2015; Tetta et al. 2016; Askouni and Papanicolaou 2017; Awani et al. 2017; Azam et al. 2017; Koutas and Bournas 2017; Raoof et al. 2017; Akhoundi et al. 2018; Koutas et al. 2018; Wakjira and Ebead 2018) and even for the concurrent seismic and energy retrofitting of building envelopes when TRM jackets are combined with thermal insulation materials (Triantafillou et al. 2017; Bournas 2018). Examples of real applications of a TRM system were presented by Bournas (2016). The acronym FRCM is also used in the literature for the same material (Carloni et al. 2015).
Although several studies exist in the literature investigating the out-of-plane performance of simply supported unreinforced masonry walls strengthened with TRM layers (Papanicolaou et al. 2008; Harajli et al. 2010; Papanicolaou et al. 2011; Babaeidarabad et al. 2014; Valluzi et al. 2014; Carozzi et al. 2015; Martins et al. 2015; Triantafillou et al. 2017; D’Ambra et al. 2018; Kariou et al. 2018), there is only one study (da Porto et al. 2013) where the out-of-plane performance of a masonry-infilled RC frame strengthened with glass-fiber net-reinforced plaster was examined. The limited results in the study of da Porto et al. (2013) indicated a textile-reinforced mortar overlay can prevent collapse of the infill wall and increase the out-of-plane strength. As also highlighted by Lunn and Rizkalla (2011), who investigated the out-of-plane strengthening of masonry infill walls with FRP, the simply supported boundary conditions may be appropriate for certain cases of masonry wall types but do not simulate the boundary conditions of masonry infill walls in RC frames. The interface between the wall and RC frame members, as well as the restraint provided by the RC frame, cannot be taken into account in simply supported masonry elements. Hence, significant sources of overstrength (such as the arching action) are ignored.
This paper aims to investigate the out-of-plane strengthening of masonry infill walls in RC frames with TRM to improve the resistance and resilience of such structures in catastrophic events by simulating as realistically as possible various boundary conditions that may exist between the infill walls and surrounding RC frame members. The effectiveness of the novel strengthening technique is assessed in terms of out-of-plane strength, stiffness and deformability enhancement, energy dissipation capacity, and residual strength. More details are given in the following sections.
Experimental Program
Test Specimens and Parameters
The experimental program aimed to study the effectiveness of externally bonded TRM jackets in enhancing the out-of-plane performance of masonry infills in RC frames. For this purpose, six RC frames with identical geometry were constructed and infilled with single-wythe or double-wythe masonry walls. Two of them were tested without strengthening and therefore served as reference specimens. The remaining four frames were first retrofitted with carbon-fiber TRM jackets and then tested to failure. Construction of the infilled frames by the same technicians allowed assessment of the effectiveness of the retrofitting method.
The geometry of the tested frames is shown in Fig. 1(a). For the half-model-to-prototype scale selected, the resulted story-height was 1.50 m (3.0 m in the prototype), and the distance between column centerlines was 1.90 m (3.8 m in the prototype). The length and height of the infill wall panels were 1.70 and 1.25 m, respectively [Fig. 1(b)], which yields a typical (for old RC buildings) length-to-height aspect ratio of 1.36.
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The columns were of rectangular cross-section (with the long side parallel to the plane of the frame), whereas the beam had a T-section with a web’s width equal to the width of the columns (140 mm) and a height equal to 250 mm. Figs. 2(a and b) show the columns and T-beam sections, respectively, with details of the steel reinforcement. For better visualization, a three-dimensional (3D) sketch of the steel reinforcement of the whole RC frame is also provided in Fig. 2(c). The column longitudinal reinforcement consisted of four 10-mm deformed bars placed at the corners. The T-beam longitudinal reinforcement consisted of two 10-mm deformed bars at the bottom and four 10-mm deformed bars at the top (two of which were placed in the flanges). The transverse reinforcement for all concrete members consisted of 6-mm-diameter smooth steel stirrups (with 90-deg hooks at the ends) placed at 150 mm distances to emulate old detailing practices. Also, as typical of substandard structures, the thickness of the concrete cover to stirrups was 10 mm, apart from the web of the beam (20 mm). The longitudinal reinforcement of all members was continuous with no lap splices. The concrete base of the RC frame, which was used to fix the frame to the strong floor of the lab, was heavily reinforced to avoid the development of cracks during testing.
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The masonry infill walls were constructed using solid (fired clay) bricks, which were laid with the small side (65 mm) and bonded to the bed mortar joints. As a result, the single wall had a thickness of 65 mm (half of the column width), whereas the double-wythe wall had a total thickness of 140 mm and was equal to the column width. In the latter case, a 10-mm mortar filling was applied between the two wythes of the wall. The bed and head mortar joints of the walls in all cases had a thickness of approximately 10 mm. To ensure a good level of confinement between the walls and RC frame, the construction of the masonry infills was completed in two stages; the last row of bricks was laid 2 days after building the rest part of the wall, allowing for free drying shrinkage development.
The role of two parameters on the effectiveness of TRM strengthening technique was investigated, namely (1) the connection configuration between the wall and the surrounding RC frame members, which depends on the boundary conditions, and (2) the thickness of the wall. A description of the specimens follows, supported by Fig. 3 and Table 1:
•
: Control specimen with unretrofitted single-wythe infill wall.
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: Double-sided TRM-strengthened single-wythe wall specimen; no connection (NOC) between the wall and the RC frame members is applied.
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: Double-sided TRM-strengthened single-wythe wall specimen, with back-side (BCK) connection of the wall to the RC frame members.
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: Double-sided TRM-strengthened single-wythe wall specimen, with front-side (FRN, side of loading) connection of the wall to the RC frame members.
•
: Control specimen with unretrofitted double-wythe infill wall.
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: Specimen with double-wythe infill wall, fully-wrapped (WRP) with TRM jacket (infill connected to the RC frame members on both sides).
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| Specimen | Strengthening configuration | Concrete compressive strengtha (MPa) | Mortar for joints | Strengthening mortar | ||
|---|---|---|---|---|---|---|
| Compressive strengtha (MPa) | Flexural strengtha (MPa) | Compressive strengtha (MPa) | Flexural strengtha (MPa) | |||
| Single-wythe walls | ||||||
| S_CON | — | 22.1 (1.5) | 11.5 (0.9) | 3.1 (0.33) | N/A | N/A |
| S_NOC | Two TRM layers on both sides of the wall; no connection to the frame | 23.4 (1.3) | 11.8 (1.2) | 2.8 (0.45) | 37.3 (0.76) | 9.2 (0.49) |
| S_BCK | Two TRM layers on both sides of the wall; connection to the back face of frame | 23.4 (1.3) | 12.9 (1.0) | 2.9 (0.28) | 36.7 (0.59) | 9.5 (0.87) |
| S_FRN | Two TRM layers on both sides of the wall; connection to the front face of frame | 22.1 (1.5) | 12.3 (0.7) | 2.6 (0.36) | 38.6 (0.91) | 8.8 (0.68) |
| Double-wythe walls | ||||||
| D_CON | — | 24.0 (1.1) | 12.1 (1.1) | 2.9 (0.21) | N/A | N/A |
| D_WRP | Full-wrapping of the infilled frame with two TRM layers | 24.0 (1.1) | 12.2 (0.9) | 2.4 (0.32) | 38.2 (1.15) | 8.6 (0.64) |
a
Standard deviation in parentheses.
In all strengthened specimens, two layers of carbon-fiber TRM were applied (for details of the strengthening material, see the section “Strengthening Material”). As shown in Fig. 3, wherever a connection was provided between the infill wall and frame columns, the TRM layers were anchored on the small sides of the columns to avoid early debonding (apart from D_WRP specimens, where full wrapping around the specimen was applied). In real-practice applications, access to the sides of the columns may not be possible due to the presence of more than one frame span. In this case, the TRM layers can be extended to the adjacent infill wall to provide the necessary anchorage length or, if needed, the external reinforcement can cover all frame spans infilled with masonry and be anchored at the sides of the external columns of the frame.
Materials
RC Frame and Masonry Infill Walls
Casting of the RC frames was done in three groups on different dates by the same mix design targeting a concrete compressive strength of 20 MPa. The average compressive strength on the day of testing the infilled frames, measured on cubes (average values from three specimens), is given for each specimen in Table 1. The 6-mm-diameter smooth longitudinal bars had a yield stress of 470 MPa, a tensile strength of 508 MPa, and an ultimate strain of 7.2%. The respective values for the 10-mm-diameter deformed bars were 545 MPa, 637 MPa, and 11.2% (average values from three specimens).
The compressive strength of the nonengineered bricks used in this study to build the masonry infill walls was equal to 21.2 MPa. This value was obtained as the average from three compressive tests on bricks capped with rapid hardening mortar and loaded perpendicular to their length with a bearing area of .
A typical mix design was used for the mortar to bind the bricks, with 1:4 cement-to-sand proportions. To obtain its compressive and flexural strength, three mortar prism samples (dimensions of ) were taken during the construction of the infill wall of each test specimen. Compressive and three-point bending tests were conducted according to EN 1015-11 (CEN 1999b) on the day of the large-scale testing; the results are summarized in Table 1 (average values from three specimens).
Apart from testing individual masonry units and the mortar used for the joints, tests were also conducted to obtain the compressive strength and elastic modulus of the masonry perpendicular to the bed joints. For this purpose, three masonry walls with dimensions of and a thickness of 65 mm were built and tested under compression after 28 days (Fig. 4), following the EN 1052-1 (CEN 1999a) specifications. As shown in Fig. 4, the deformation of the wall was measured within a gauge length of 250 mm in the center of both faces using a potentiometer. According to the results, the mean compressive strength of the masonry was 9.7 MPa, and the secant elastic modulus obtained from the stress-strain curves in the region of 0%–30% of the maximum stress was equal to 2.5 GPa.
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Strengthening Material
A carbon-fiber textile was used as reinforcement of the TRM composite material, which was externally applied in layers to retrofit the masonry infill walls. The textile used [Fig. 5(a)] had a weight of with uncoated carbon fiber rovings in two orthogonal directions and a 50%–50% distribution of fibers in each direction. The nominal thickness of the textile in each direction was 0.095 mm (based on the smeared distribution of fibers). According to the manufacturer data sheets, the tensile strength and modulus of elasticity of the carbon fibers were 3,800 MPa and 225 GPa, respectively.
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The mortar used as the matrix of the TRM composite and binding material between the textile and substrates (masonry or concrete) was a polymer-modified cement-based mortar with an 8:1 cement-to-polymers ratio by weight. The water-to-cementitious material ratio by weight was equal to 0.23, resulting in plastic consistency and good workability. Table 1 includes the strength properties of the mortar (average values of three specimens) obtained experimentally on the day of testing using prisms of dimensions, according to EN 1015-11 (CEN 1999b).
To obtain the mechanical properties of the composite material, tensile tests on TRM coupons were conducted. Three dumbbell specimens [Fig. 5(b)] were fabricated, comprising one layer of carbon-fiber textile (same as the textile used for strengthening) embedded in the middle of a 10-mm-thick layer of mortar (same as the mortar used for strengthening). The test setup used in this study included clamping of the specimen ends using curved steel flanges with rubber pads in between in order to avoid stress concentration. In-plane and out-of-plane rotation of the specimens was allowed through the use of spherical joints, whereas deformations were measured via Linear Variable Differential Transformer (LVDTs) attached on both sides of the specimens within the gauge length (200 mm). According to the results, the mean tensile strength was 1,382 MPa (calculated on the basis of the textile-fiber cross-sectional area), the mean ultimate strain was 0.79% (calculated as the average strain over a gauge length of 200 mm), and the modulus of elasticity was 163.3 GPa (calculated as the secant modulus of elasticity of the postcracking response, which reflects the activation of the textile fibers in tension). Fig. 5(c) shows the stress-strain curves obtained from the coupon tests. Failure of all specimens was due to partial rupture of the fibers at an intermediate crack within the gauge length, accompanied by slippage of the fibers within the matrix. Importantly, different test methods [e.g., the clevis-grip test method required by ACI 549 (ACI 2013)] yield different results according to the recent study of D’Antino and Papanicolaou (2018), owing to the difference in the clamping method of the TRM coupon ends. The setup used in this study is believed by the authors to provide more realistic results for strengthening applications of TRM.
Strengthening Procedure
The strengthening procedure was a typical wet lay-up application and included the following steps:
1.
Removal of a thin layer of concrete and formation of a grid of grooves (2 mm deep) at the surface of the columns and beams to receive strengthening.
2.
Rounding of the column corners with a radius of 15 mm, to avoid stress concentration of the TRM jackets.
3.
Dampening of both concrete and masonry surfaces to receive strengthening [Fig. 6(a)].
4.
Application of the first mortar layer with a thickness of 3–4 mm using a smooth metal trowel [Fig. 6(b)].
5.
Application of the first textile layer into the mortar by hand pressure [Fig. 6(c)].
6.
Application of a mortar layer to completely cover the textile layer previously applied.
7.
Application of the second TRM layer by repeating the same procedure as for the first TRM layer [steps shown in Figs. 6(d and e)] while the previous layer was still in the fresh state.
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Each specimen was strengthened on both sides, but the strengthening configuration on each side depended on the boundary conditions between the wall and the adjoining RC frame members (Fig. 3). Specifically, when there was no step between the wall and the frame members, the TRM jacket was extended to the faces of all the frame members (columns and top beam) and to the sides of the columns. In contrast, when there was a step, the TRM jacket was applied only on the face of the wall. The width of the textile roll (equal to 1,250 mm) was adequate to cover up to the clear height of the infill wall, not including the beam’s height. For that reason, when the strengthening configuration included extension of the textile to the beam’s face, two separate textile parts were overlapped over a length of 200 mm. The overlap region was not the same for the two TRM layers; overlapping of the first TRM layer took place close to the bottom of the wall, whereas the second TRM layer was overlapped close to the top of the wall [Fig. 6(d)].
Test Setup and Testing Procedure
All specimens were subjected to monotonic out-of-plane loading and tested to failure. Fig. 7 illustrates the setup implemented for the tests. The specimens were fixed to the lab’s strong floor via prestressing 12 steel rods passing through the RC base [Fig. 7(a)]. The load was applied by a 500 kN-capacity servo-hydraulic actuator, which was mounted on a stiff steel reaction frame. A system of steel beams was used to spread the load into four points as shown in Fig. 7(a); allowing to achieve more uniform load distribution. The four load-application areas were centrally located at horizontal and vertical distances equal to one-third of the clear length () and height () of the infill walls, respectively [Fig. 7(b)]. To avoid concentration of high local stresses in these areas, square rubber pads () were placed between the wall and spreading beams. For convenience, the face of the specimen that was in contact with the loading configuration is hereafter called the front face, whereas the other is called the back face.
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As shown in Fig. 7(c), the horizontal out-of-plane displacement of the two RC frame’s beam-column joints was restrained by adding two stiff steel beams on the back face of the specimens, which were mounted on an external stiff steel reaction frame. The contact area between the steel beams and the beam-column joints was equal to . Recorded measurements included load and displacement values from the actuator and out-of-plane displacement values from nine potentiometers attached on the back face of the specimen. Specifically, as illustrated in Figs. 7(c and d), four potentiometers were fixed on the perimeter of the infill (POT1–POT4) to measure the relative out-of-plane movement of the wall with respect to the frame, whereas five potentiometers (POT5–POT9) were mounted on an external reference system to monitor the absolute out-of-plane displacements of the masonry wall in both bending directions. The exact locations of these sensors are provided in Fig. 7(d).
The testing procedure included application of the load monotonically at a constant displacement rate of . A data acquisition system was used to monitor and record synchronized data from all sensors at a sampling rate of 4 Hz. The termination point of the tests was decided individually for each specimen, taking into consideration risks from possible wall collapse.
Test Results
Table 2 summarizes the test results in terms of peak load, displacement at maximum load, strengthening effectiveness, initial stiffness, and energy absorption. The strengthening effectiveness is defined as the ratio of the maximum load attained by a retrofitted specimen to the capacity of the corresponding control specimen. The initial stiffness values have been calculated as the secant stiffness of the load versus displacement curves (Fig. 8) from 0% to 40% of the peak load. Finally, the energy absorption values represent the area under the load versus displacement curves from 0 to 57 mm, which is the minimum ultimate displacement value recorded among all specimens. The out-of-plane behavior of the specimens is described subsequently, grouped based on the thickness of the infill wall. The out-of-plane displacement used to describe the behavior of all specimens is the displacement measured at the center of the wall (POT9).
| Specimen | Peak load, (kN) | Central displacement at (mm) | Strengthening effectiveness factor, | Initial stiffnessa () | Energy absorptionb (kJ) |
|---|---|---|---|---|---|
| S_CON | 29 | 22.5 | — | 7.4 | 1.53 |
| S_NOC | 110 | 28.4 | 3.79 | 13.1 (77%) | 3.89 (154%) |
| S_BCK | 118 | 31.1 | 4.07 | 19.5 (164%) | 5.03 (229%) |
| S_FRN | 158 | 19.8 | 5.45 | 20.1 (172%) | 5.52 (261%) |
| D_CON | 56 | 10.5 | — | 21.8 | 2.87 |
| D_WRP | 137 | 32.7 | 2.45 | 51.5 (136%) | 6.85 (138%) |
a
Secant stiffness of the load versus central displacement curve, from 0% to 40% of the peak load. The values in the brackets represent the percentage increase with respect to the corresponding control specimen.
b
Cumulative energy absorption at a central out-of-plane displacement of 57 mm. The values in the brackets represent the percentage increase with respect to the corresponding control specimen.
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Single-Wythe Infill Walls
The reference frame with a single-wythe masonry infill wall (S_CON) attained a maximum load of 29 kN at a corresponding out-of-plane displacement at the center of the wall equal to 22.5 mm, whereas its initial stiffness was equal to . The first cracks appeared at a load level of 10 kN, at a bed joint in the central region of the wall’s back face. With increasing load, the cracks on the back face propagated toward the four corners, mainly following the mortar joints [step-cracking; Fig. 9(a)], whereas on the front face (loading side), the wall was separated from the RC frame on the perimeter. With the main mechanism of carrying forces being the arching action, the max load of 29 kN was reached and remained almost constant until the arching action started degrading at a displacement of approximately 30 mm (Fig. 8). With increasing out-of-plane displacements, the load dropped gradually until the test was terminated. The cumulative energy absorption at 57 mm displacement was 1.53 kJ. Fig. 9(b) shows specimen S_CON after the termination of the test.
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The specimen with the retrofitted single-wythe wall not connected to the surrounding RC frame members (S_NOC) reached a peak load of 110 kN at a displacement of 28.4 mm. The initial stiffness of the S_NOC specimen was equal to . The first cracks in the TRM jacket appeared at a load of 50 kN on the back face of the wall (at a displacement of 4 mm), whereas at 60 kN, the infill panel was detached from the middle region of the top beam and experienced shear sliding [visible in Fig. 9(c), after the test completion]. With the wall being supported only at the three sides of the frame (the two columns and the base), the load continued to increase, and further cracks developed in the TRM jacket due to bending. Fig. 9(d) shows the crack pattern at the back face of the specimen just before the maximum load of 110 kN was reached. Full detachment of the wall from the beam and columns led to sudden load drop. Insignificant residual strength was provided by the wall, which moved as a solid panel after failure occurred. The cumulative energy absorption at 57 mm displacement was equal to 3.89 kJ.
Specimen S_BCK reached a maximum load of 118 kN at a displacement of 31.1 mm. The initial stiffness of the S_BCK specimen was . Up to a load of 70 kN, there was no visible damage on the TRM jacket. As the load increased, the wall detached from the top beam’s middle region (shear sliding), resulting in a significant change in the slope of the load versus displacement curve (Fig. 8). Until the maximum load was reached (118 kN), the wall’s back face experienced multiple cracking, whereas extensive shear sliding occurred between the wall and beam. As a result, the TRM overlays started to gradually debond from the beam’s surface; debonding was at the interface between the TRM and the concrete substrate. When full debonding occurred [Fig. 9(e)], the load dropped to approximately 80 kN (at a displacement of 38 mm). Redistribution of stresses led to stiffness and load recovery as the wall started detaching from the two columns (shear sliding), thus activating the TRM overlays in the region. After a load recovery of 20 kN, the TRM jacket debonded from the face of the left column [looking from the back; Fig. 9(f)] and the load dropped again at 80 kN (at a displacement of 55 mm). Ultimately, at very large displacements (65 mm), the TRM jacket also debonded from the face of the right column. Slippage of the textile fibers through the mortar at the side faces of the two columns [Fig. 9(f)] resulted in gradual load drop. This failure mode was also observed during the TRM characterization tests, as described in the relevant section. The cumulative energy absorption at 57 mm displacement was equal to 5.03 kJ. Fig. 9(g) shows the back face of specimen S_BCK after test termination.
Specimen S_FRN reached a maximum load of 158 kN at a displacement of 19.8 mm. The initial stiffness of the S_FRN specimen was . The behavior of this specimen was smooth until the peak load was reached. At a load of 70 kN, the first cracks appeared on the wall’s back face (at a displacement of 3.5 mm), as well as on the back face of the beam and columns. Multiple cracks on the wall’s back face continued to develop and propagate toward the four corners [Fig. 9(h)], indicating an excellent connection between the infill wall and frame. This behavior resulted in smooth stiffness reduction. On the loading side, minor cracks developed on the face of the TRM jacket at the boundary between the infill and RC frame members. At a load of 145 kN (at a displacement of 13 mm), shear sliding between bricks occurred close to the loading area [Fig. 9(i)]. The TRM jacket was then activated in tension and helped to reach a peak load of 158 kN (at about 20 mm displacement). Then, local rupture of fibers led to a significant load drop (to 110 kN), accompanied by redistribution of the stresses and activation of the TRM jacket in an undamaged region within the close perimeter of the loading area. Further resistance was then provided by the jacket until a second load drop was recorded (to 65 kN) due to fiber rupture (at a displacement of 33 mm). The wall exhibited significant residual strength of approximately 80 kN owing to the continuous stress redistribution at the perimeter of the loading area. Stresses were redistributed by the textile with the matrix-fiber interface area engaged, being shifted to adjacent undamaged regions. This favorable characteristic of TRM has also been observed in other past studies (e.g., Triantafillou and Papanicolaou 2006; Bournas et al. 2007; Papanicolaou et al. 2008; Koutas et al. 2015) and constitutes a main advantage over the use of competitive materials such as FRPs.
The cumulative energy absorption at 57 mm displacement was 5.52 kJ. The damage of the wall’s back face after the termination of the test is shown in Fig. 9(j).
Double-Wythe Infill Walls
The reference frame with a double-wythe masonry infill wall (D_CON) attained a maximum load of 56 kN at a displacement of 10.5 mm. The initial stiffness was as high as . The first cracks appeared on the infill’s body (at the back face) at a load of 40 kN (at a displacement of 1.5 mm). The propagation of the cracks toward the four corners and the development of minor cracks at the beam and columns resulted in gradual reduction of the stiffness. After reaching a maximum value of 56 kN, the load followed a large plateau (Fig. 8), and the cracks already developed became wider [Fig. 10(a)]. Only when the displacement reached 30 mm did the load start to gradually decrease. With further increase of displacements, crushing of a few bricks was observed on the loading side [Fig. 10(b)]. The test was terminated at very large displacements due to the risk of the wall’s collapse. The cumulative energy absorption at 57 mm displacement was equal to 2.87 kJ. A picture of the specimen after the test termination is presented in Fig. 10(c).
CC.1943-5614.0000911/asset/7dae0234-82c0-4fd6-b9ca-d21b4f8bd277/assets/images/large/figure10.jpg)
Specimen D_WRP attained a maximum load of 137 kN at a displacement of 32.7 mm, whereas its initial stiffness was . The specimen had no visible damage up to a load of 80 kN. Detachment of the wall from the top beam (shear sliding) was recorded at 90 kN and resulted in stiffness reduction. Up to a load of 128 kN, multiple cracks developed in the TRM jacket on the back face of the wall [Fig. 10(d)], whereas on the front side, a single crack appeared at the boundary between the jacket and frame members. At a displacement of 12.5 mm, detachment of the wall from the bottom and two columns led to a small load drop at 115 kN, after which the wall recovered part of its stiffness due to stress redistribution and reached the maximum load. At that point, the TRM jacket experienced full debonding from the beam and partial debonding from the two faces of the columns at the back side, which resulted in a small load drop to 122 kN (at a displacement of 36 mm). Up to very large displacements, the specimen exhibited high residual strength approximately equal to 120 kN. The test was terminated at a displacement of 77 mm because of a limitation on the stroke of the actuator. The cumulative energy absorption at 57 mm central displacement was equal to 6.85 kJ. Figs. 10(e–g) present specimen D_WRP after test termination.
Discussion of Results
All TRM-retrofitted specimens dramatically enhanced the out-of-plane overall performance of the masonry-infilled RC frames. In terms of the parameters investigated in this experimental program, an examination of the results in terms of out-of-plane load-carrying capacity, initial stiffness, energy absorption, and residual strength revealed the information presented in the following sections. Fig. 11 supports the comparisons between the retrofitted specimens by providing effectiveness factors for the relevant mechanical properties (the factors were calculated by dividing the mechanical property of the retrofitted specimen with the corresponding value of the control specimen).
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Out-of-Plane Strength
The retrofitted single-wythe wall without connection to the frame (S_NOC) sustained a maximum load of 110 kN, achieving a 3.79 times higher capacity compared to the unretrofitted wall). This is explained by the significant contribution of the textile fibers in carrying the tensile forces as a result of the wall’s two-way bending.
When the wall was connected to the back face of the frame members (S_BCK), the maximum load attained was only 7.2% higher compared to S_NOC, resulting in a 4.07 times higher capacity with respect to the control specimen. In contrast, when the retrofitted wall was connected to the front face of the frame members (S_FRN), the infilled frame exhibited the best load capacity among all specimens, achieving: (1) a 43.6% higher load compared to the retrofitted wall without connection to the frame (S_NOC) and (2) a strengthening effectiveness factor of 5.45. The inability of the back side connection to provide a high strength increase is linked to the premature failure mechanism observed. In specific, TRM debonding from the faces of the RC frame members occurred as a result of the wall’s shear sliding on the boundaries. Thus, tensile forces were developed at the TRM-concrete interface, which in combination with the shear forces promoted early debonding of the jackets. On the other hand, the front-side connection proved to be highly effective in increasing the out-of-plane strength owing to the better anchorage conditions developed on the boundaries. It is therefore concluded that the type of connection (front or back) had a significant effect on the out-of-plane load-carrying capacity of the masonry infill walls, with the connection on the front (loading) side yielding better results.
Doubling the thickness of the infill walls by constructing two single-wythe walls in parallel resulted in a nearly double out-of-plane load-carrying capacity of the unretrofitted infill wall (from 29 to 56 kN). Having a wall thickness equal to the width of the columns and the beam allowed for the full wrapping of the infilled frame with a two-layer TRM jacket. As a result, the retrofitted specimen (D_WRP) reached a maximum load of 137 kN (a 2.45 times higher capacity compared to D_CON), which was 13% lower than the maximum load attained by the stronger of the single-wythe wall specimens, S_FRN. The relatively lower strengthening effectiveness factor of 2.45 for D_WRP specimen, compared to all the retrofitted single-wythe wall specimens (varying from 3.70 to 5.45), is because the ultimate load response was governed by the capacity of the connections and not by the flexural capacity of the wall itself. This result reveals the importance of the test setup used in this study, which is more representative of the real behavior of masonry-infilled walls; flexural tests on masonry subassemblages would miss this masonry infill-RC frame interaction and therefore would not yield realistic conclusions.
Initial Stiffness
The initial stiffness of the specimens was calculated as the secant stiffness of the load versus central displacement curve, from 0% to 40% of the peak load. As shown in Table 2 and Fig. 11, due to the presence of the TRM jacket on both sides of the infill wall, its initial stiffness was increased in all retrofitted specimens from 77% to 172% with respect to the unretrofitted one (effectiveness factor from 1.77 to 2.72; Fig. 11).
The connection of the TRM jacket to the back or front face of the frame members increased the initial stiffness of the wall by 49% and 53% (S_BCK and S_FRN), respectively, compared to the retrofitted wall without a connection to the frame (S_NOC). Hence, it is concluded that, although connecting the wall to the frame members substantially increased its initial stiffness, the connection configuration (connection to the front or back face of the frame) did not play a critical role at the early stage of loading. This is expected because no detachment (shear sliding) of the wall was observed up to the point where the initial stiffness was measured. The initiation of shear sliding between the wall and frame members was the key point where different connections started having a different impact on the results.
The initial stiffness of the control specimen with the double-wythe wall (D_CON) was more than doubled (almost tripled) when compared to the control specimen with the single-wythe wall (S_CON). It is believed that the mortar used to fill the gap between the two wall layers (10-mm gap) contributed significantly to the nonproportional increase in the initial stiffness. Assuming that the initial bending stiffness of the wall is proportional to the section’s moment of inertia, a double wall consisting of two wythes without any connection between them should have a bending stiffness that is two times the stiffness of a single wall. However, in this study, the presence of mortar between the two wall wythes developed shear forces at the interfaces, thus providing a degree of connection. Nevertheless, this connection was far from monolithic; a monolithic wall with double thickness would be expected to increase the bending stiffness as large as 8 times.
The effectiveness factor of the full-wrapping technique in increasing the initial stiffness of the double-wythe wall specimen was equal to 2.36, which is within the range of values observed in the single-wythe wall specimens.
Energy Absorption
The comparative behavior of the specimens in terms of cumulative energy absorption (represented by the area under the force versus central out-of-plane displacement curves) is illustrated in Fig. 12. All retrofitted specimens exhibited substantially improved behavior in terms of energy absorption compared to the corresponding control specimen, with an effectiveness factor varying from 2.39 to 3.61 times at an out-of-plane displacement of 57 mm (Fig. 11).
CC.1943-5614.0000911/asset/6ddf66ea-ad21-42ab-b7f7-3304c8e799bf/assets/images/large/figure12.jpg)
The retrofitted wall without connection to the frame (S_NOC) behaved identically to the retrofitted wall connected to the back face of the frame members (S_BCK), up to a displacement of approximately 30 mm (where the peak load was recorded for both specimens). However, the energy absorbed in the postpeak response of specimen S_BCK was higher than that of S_NOC owing to the contribution of the connection, resulting in a 29% increase recorded at an out-of-plane displacement of 57 mm. The favorable connection of the wall to the front (loading) face of the frame in specimen S_FRN, though, resulted in an overall better performance compared to the S_BCK and S_NOC specimens. Ultimately, at an out-of-plane displacement of 57 mm, the cumulative energy absorption of the S_FRN specimen was approximately 10% and 42% higher with respect to the S_BCK and S_NOC specimens, respectively. The better behavior of the S_FRN specimen can be attributed to the energy dissipated through the stress-redistribution mechanism enabled after the peak load reached, as described in the “Test Results” section. The importance of the connection between the wall and frame is therefore verified in terms of energy absorption as well.
As shown in Fig. 12, specimen D_WRP exhibited the best behavior among all specimens in terms of absolute values of energy absorption. Specifically, D_WRP reached a value of 6.85 kJ at an out-of-plane displacement of 57 mm, which is 24% higher compared to S_FRN (the best among the single-wythe wall specimens). Nevertheless, normalized to the D_CON specimen, the resulting effectiveness factor is 2.39, which is the lowest among all strengthened specimens, but still very high.
Residual Strength and Resilience
Another interesting aspect in most of the specimens’ behavior is the residual strength they exhibited at the postpeak spectrum of out-of-plane displacements. By comparing the load-displacement curves in Fig. 8 among specimens S_CON, S_NOC, S_BCK, and S_FRN, it was observed that both S_CON and S_NOC specimens exhibited poor behavior in terms of residual strength, with the load dropping at low levels for large out-of-plane displacements. In contrast, in specimens S_BCK and S_FRN, the load, after reaching its maximum value, dropped to a load level of approximately 80 kN, exhibiting signs of residual strength for a wide range of large out-of-plane displacements. Key to the residual strength observed in those two specimens were the connection details of the retrofitted wall to the RC frame members. In particular, anchoring the TRM overlays on the two small sides of the columns helped avoid a full debonding failure mechanism.
The best behavior in terms of residual strength was exhibited by D_WRP specimen, which, after the maximum capacity was reached (137 kN), retained most of its capacity (approximately equal to 120 kN) without any signs of deterioration up to very large displacements. This was a result of the excellent connection provided by the full-wrapped jackets. In this case, the risk of the wall’s collapse is drastically mitigated, effectively increasing its resilience. When full wrapping is not feasible due to geometrical constraints, two-sided (specimens S_BCK or S_FRN) or U-shaped TRM jackets that also have high effectiveness could be used.
Conclusions
This paper presented an experimental study on the use of textile-reinforced mortar jackets as a means of improving the out-of-plane performance of masonry infill walls in reinforced concrete frames. The effect of different connection configurations between the infill wall and the RC frame members, as well as the effect of the wall’s thickness, were carefully examined and discussed in terms of ultimate load-carrying capacity, initial stiffness, energy absorption, and residual strength. The main conclusions drawn from this study are summarized subsequently:
•
The suggested strengthening technique was highly effective in increasing the out-of-plane load carrying capacity of the infill walls. The strengthening effectiveness factor varied from 3.79 to 5.45 for single-wythe wall specimens and was equal to 2.45 in the case of the double-wythe wall specimen.
•
The peak load was governed by the capacity of the connections and not by the flexural capacity of the wall itself. The retrofitted single-wythe wall with a connection on the front side (loading side) of the frame exhibited the highest capacity among all specimens.
•
For the specimens with a single-wythe wall, it was concluded that, although connecting the wall to frame members substantially increased its initial stiffness (approximately 20%), the different connection configuration did not play a critical role at the early stage of loading.
•
The energy absorption of the retrofitted walls was enhanced from 138% to 261% compared to the control ones, with the connection configuration playing an important role. Connecting the single-wythe wall to the front or back side of the frame resulted in a 29% and 42% increase, respectively, compared to the wall without a connection to the frame.
•
The residual strength of the single-wythe walls was only improved when the walls were connected to the RC frame, thus mitigating the risk of collapse and enhancing the resilience of such retrofitted structures.
To summarize, this study proved the effectiveness of the TRM strengthening technique in improving the out-of-plane performance of masonry infill walls in RC frames, highlighting the importance of the connection configuration, which depends on the boundary conditions between the infill and frame members. Importantly, the results of this study neglect the possible adverse effects of cyclic loading or the possible beneficial effects of an impact or blast loading. In this context, the observed performance is believed by the authors to represent the maximum possible for the case of earthquake loading or the minimum for the case of blast loads.
Future research could be directed toward examining the combined effect of in-plane and out-of-plane actions or the effect of cyclic or blast loading, as well as developing appropriate numerical models to simulate the out-of-plane behavior of TRM-strengthened masonry-infilled RC frames.
Acknowledgments
The authors wish to acknowledge the help of the former PhD students Saad Raoof and Zoi Tetta in the preparation of the specimens.
References
ACI (American Concrete Institute). 2013. Guide to design and construction of externally bonded fabric-reinforced cementitious matrix (FRCM) systems for repair and strengthening concrete and masonry structures. ACI 549. Farmington Hills, MI: ACI.
Akhoundi, F., G. Vasconcelos, P. Lourenço, L. M. Silva, F. Cunha, and R. Fangueiro. 2018. “In-plane behavior of cavity masonry infills and strengthening with textile reinforced mortar.” Eng. Struct. 156: 145–160. https://doi.org/10.1016/j.engstruct.2017.11.002.
Almusallam, T. H., and Y. A. Al-Salloum. 2007. “Behavior of FRP strengthened infill walls under in-plane seismic loading.” J. Compos. Constr. 11 (3): 308–318. https://doi.org/10.1061/(ASCE)1090-0268(2007)11:3(308).
Askouni, P. D., and C. G. Papanicolaou. 2017. “Experimental investigation of bond between glass textile reinforced mortar overlays and masonry: The effect of bond length.” Mater. Struct. 50 (2): 164. https://doi.org/10.1617/s11527-017-1033-7.
Augenti, N., F. Parisi, A. Prota, and G. Manfredi. 2011. “In-plane lateral response of a full-scale masonry subassemblage with and without an inorganic matrix-grid strengthening system.” J. Compos. Constr. 15 (4): 578–590. https://doi.org/10.1061/(ASCE)CC.1943-5614.0000193.
Awani, O., T. El-Maaddawy, and N. Ismail. 2017. “Fabric-reinforced cementitious matrix: A promising strengthening technique for concrete structures.” Constr. Build. Mater. 132: 94–111. https://doi.org/10.1016/j.conbuildmat.2016.11.125.
Azam, R., K. Soudki, J. S. West, and M. Noël. 2017. “Behavior of shear-critical RC beams strengthened with CFRCM.” J. Compos. Constr. 22 (1): 04017046. https://doi.org/10.1061/(ASCE)CC.1943-5614.0000829.
Babaeidarabad, S., F. De Caso, and A. Nanni. 2014. “URM walls strengthened with fabric-reinforced cementitious matrix composite subjected to diagonal compression.” J. Compos. Constr. 18 (2): 04013045. https://doi.org/10.1061/(ASCE)CC.1943-5614.0000441.
Bournas, D. A. 2016. “Strengthening of existing structures: Selected case studies.” Chap. 17 in Textile fibre composites in civil engineering, edited by T. C. Triantafillou, 389–411. Duxford, UK: Woodhead.
Bournas, D. A. 2018. “Concurrent seismic and energy retrofitting of RC and masonry building envelopes using inorganic textile-based composites with insulation materials: A new concept.” Compos. Part B 148: 166–179. https://doi.org/10.1016/j.compositesb.2018.04.002.
Bournas, D. A., P. V. Lontou, C. G. Papanicolaou, and T. C. Triantafillou. 2007. “Textile-reinforced mortar versus fiber-reinforced polymer confinement in reinforced concrete columns.” ACI Struct. J. 104 (6): 740–748. https://doi.org/10.14359/18956.
Carloni, C., et al. 2015. “Fiber reinforced composites with cementitious (inorganic) matrix.” Chap. 9 in Design procedures for the use of composites in strengthening of reinforced concrete structures: State of the art report of the RILEM TC 234-DUC, edited by C. Pellegrino, and J. Sena-Cruz, 349–391. Dordrecht, Netherlands: Springer.
Carozzi, F. G., P. Colombi, and C. Poggi. 2015. “Fabric reinforced cementitious matrix (FRCM) systems for strengthening of masonry elements subjected to out-of-plane loads.” In Proc., 7th Biennial Conf. on Advanced Composites in Construction, Cambridge, UK.
CEN (European Committee for Standardization). 1999a. Methods of test for masonry. Part 1: Determination of compressive strength. EN 1052-1. Brussels, Belgium: CEN.
CEN (European Committee for Standardization). 1999b. Methods of test for mortar for masonry. Part 11: Determination of flexural and compressive strength of hardened mortar. EN 1015-11. Brussels, Belgium: CEN.
D’Ambra, C., G. P. Lignola, A. Prota, E. Sacco, and F. Fabbrocino. 2018. “Experimental performance of FRCM retrofit on out-of-plane behaviour of clay brick walls.” Compos. Part B 148: 198–206. https://doi.org/10.1016/j.compositesb.2018.04.062.
D’Ambrisi, A., and F. Focacci. 2011. “Flexural strengthening of RC beams with cement-based composites.” J. Compos. Constr. 15 (5): 707–720. https://doi.org/10.1061/(ASCE)CC.1943-5614.0000218.
D’Antino, T., and C. Papanicolaou. 2018. “Comparison between different tensile test set-ups for the mechanical characterization of inorganic-matrix composites.” Constr. Build. Mater. 171: 140–151. https://doi.org/10.1016/j.conbuildmat.2018.03.041.
da Porto, F., G. Guidi, M. Dalla Benetta, and N. Verlato. 2013. “Combined in-plane/out-of-plane experimental behaviour of reinforced and strengthened infill masonry walls.” In Proc., 12th Canadian Masonry Symp. Vancouver, Canada.
De Felice, G., et al. 2014. “Mortar-based systems for externally bonded strengthening of masonry.” Mater. Struct. 47 (12): 2021–2037. https://doi.org/10.1617/s11527-014-0360-1.
El-Maaddawy, T., and A. El Refai. 2015. “Innovative repair of severely corroded T-beams using fabric-reinforced cementitious matrix.” J. Compos. Constr. 20 (3): 04015073. https://doi.org/10.1061/(ASCE)CC.1943-5614.0000641.
Elsanadedy, H. M., T. H. Almusallam, S. H. Alsayed, and Y. A. Al-Salloum. 2013. “Flexural strengthening of RC beams using textile reinforced mortar—Experimental and numerical study.” Compos. Struct. 97: 40–55. https://doi.org/10.1016/j.compstruct.2012.09.053.
Harajli, M., H. El Khatib, and J. T. San-Jose. 2010. “Static and cyclic out-of-plane response of masonry walls strengthened using textile-mortar system.” J. Mater. Civ. Eng. 22 (11): 1171–1180. https://doi.org/10.1061/(ASCE)MT.1943-5533.0000128.
Kariou, F. A., S. P. Triantafyllou, D. A. Bournas, and L. N. Koutas. 2018. “Out-of-plane response of masonry walls strengthened using textile-mortar system.” Constr. Build. Mater. 165: 769–781. https://doi.org/10.1016/j.conbuildmat.2018.01.026.
Koutas, L., S. N. Bousias, and T. C. Triantafillou. 2015. “Seismic strengthening of masonry-infilled RC frames with TRM: Experimental study.” J. Compos. Constr. 19 (2): 04014048. https://doi.org/10.1061/(ASCE)CC.1943-5614.0000507.
Koutas, L., A. Pitytzogia, T. C. Triantafillou, and S. N. Bousias. 2014. “Strengthening of infilled reinforced concrete frames with TRM: Study on the development and testing of textile-based anchors.” J. Compos. Constr. 18 (3): A4013015. https://doi.org/10.1061/(ASCE)CC.1943-5614.0000390.
Koutas, L. N., and D. A. Bournas. 2017. “Flexural strengthening of two-way RC slabs with textile-reinforced mortar: Experimental investigation and design equations.” J. Compos. Constr. 21 (1): 04016065. https://doi.org/10.1061/(ASCE)CC.1943-5614.0000713.
Koutas, L. N., Z. Tetta, D. A. Bournas, and T. C. Triantafillou. 2018. “Strengthening of concrete structures with textile reinforced mortars (TRM): State-of-the-art review.” J. Compos. Constr. 23 (1), in press.
Loreto, G., S. Babaeidarabad, L. Leardini, and A. Nanni. 2015. “RC beams shear-strengthened with fabric-reinforced-cementitious-matrix (FRCM) composite.” Int. J. Adv. Struct. Eng. (IJASE) 7 (4): 341–352. https://doi.org/10.1007/s40091-015-0102-9.
Lunn, D. S., and S. H. Rizkalla. 2011. “Strengthening of infill masonry walls with FRP materials.” J. Compos. Constr. 15 (2): 206–214. https://doi.org/10.1061/(ASCE)CC.1943-5614.0000088.
Martins, A., G. Vasconcelos, R. Fangueiro, and F. Cunha. 2015. “Experimental assessment of an innovative strengthening material for brick masonry infills.” Compos. Part B 80: 328–342. https://doi.org/10.1016/j.compositesb.2015.06.012.
Ozcebe, G., U. Ersoy, T. Tankut, E. Erduran, O. Keskin, and C. Mertol. 2003. Strengthening of brick-infilled RC frames with CFRP. Ankara, Turkey: METU.
Ozden, S., U. Akguzel, and T. Ozturan. 2011. “Seismic strengthening of infilled reinforced concrete frames with composite materials.” ACI Struct. J. 108 (4): 414–422.
Papanicolaou, C. G., T. C. Triantafillou, and M. Lekka. 2011. “Externally bonded grids as strengthening and seismic retrofitting materials of masonry panels.” Constr. Build. Mater. 25 (2): 504–514. https://doi.org/10.1016/j.conbuildmat.2010.07.018.
Papanicolaou, C. G., T. C. Triantafillou, M. Papathanasiou, and K. Karlos. 2008. “Textile-reinforced mortar (TRM) versus FRP as strengthening material of URM walls: Out-of-plane cyclic loading.” Mater. Struct. 41 (1): 143–157. https://doi.org/10.1617/s11527-007-9226-0.
Pinto, A., H. Varum, and J. Molina. 2002. “Experimental assessment and retrofit of full-scale models of existing RC frames.” In Proc., 12th European Conf. on Earthquake Engineering (ECEE). Amsterdam, Netherlands: Elsevier.
Raoof, S. M., and D. A. Bournas. 2017a. “Bond between TRM versus FRP composites and concrete at high temperatures.” Compos. Part B 127: 150–165. https://doi.org/10.1016/j.compositesb.2017.05.064.
Raoof, S. M., and D. A. Bournas. 2017b. “TRM versus FRP in flexural strengthening of RC beams: Behaviour at high temperatures.” Constr. Build. Mater. 154: 424–437. https://doi.org/10.1016/j.conbuildmat.2017.07.195.
Raoof, S. M., L. N. Koutas, and D. A. Bournas. 2017. “Textile-reinforced mortar (TRM) versus fibre-reinforced polymers (FRP) in flexural strengthening of RC beams.” Constr. Build. Mater. 151: 279–291. https://doi.org/10.1016/j.conbuildmat.2017.05.023.
Tetta, Z. C., and D. A. Bournas. 2016. “TRM vs FRP jacketing in shear strengthening of concrete members subjected to high temperatures.” Compos. Part B 106: 190–205. https://doi.org/10.1016/j.compositesb.2016.09.026.
Tetta, Z. C., L. N. Koutas, and D. A. Bournas. 2016. “Shear strengthening of full-scale RC T-beams using textile-reinforced mortar and textile-based anchors.” Compos. Part B 95: 225–239. https://doi.org/10.1016/j.compositesb.2016.03.076.
Triantafillou, T. C., K. Karlos, K. Kefalou, and E. Argyropoulou. 2017. “An innovative structural and energy retrofitting system for URM walls using textile reinforced mortars combined with thermal insulation: Mechanical and fire behavior.” Constr. Build. Mater. 133: 1–13. https://doi.org/10.1016/j.conbuildmat.2016.12.032.
Triantafillou, T. C., and C. G. Papanicolaou. 2006. “Shear strengthening of reinforced concrete members with textile reinforced mortar (TRM) jackets.” Mater. Struct. 39 (1): 93–103. https://doi.org/10.1007/s11527-005-9034-3.
Valluzi, M. R., F. da Porto, E. Garbin, and M. Panizza. 2014. “Out-of-plane behaviour of infill masonry panels strengthened with composite materials.” Mater. Struct. 47 (12): 2131–2145. https://doi.org/10.1617/s11527-014-0384-6.
Verderame, G. M., I. Iervolino, and P. Ricci. 2009. “Report on the damages on buildings following the seismic event of the 6th of April 2009.” Accessed November 1, 2017. http://www.reluis.it.
Vicente, R., H. Rodrigues, H. Varum, A. Costa, and J. M. Silva. 2012. “Performance of masonry enclosure walls: Lessons learned from recent earthquakes.” Earthquake Eng. Eng. Vibr. 11 (1): 23–34. https://doi.org/10.1007/s11803-012-0095-3.
Wakjira, T. G., and U. Ebead. 2018. “FRCM/internal transverse shear reinforcement interaction in shear strengthened RC beams.” Compos. Struct. 201: 326–339. https://doi.org/10.1016/j.compstruct.2018.06.034.
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Journal of Composites for Construction
Volume 23 • Issue 1 • February 2019
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Received: Nov 2, 2017
Accepted: Jul 12, 2018
Published online: Dec 8, 2018
Published in print: Feb 1, 2019
Discussion open until: May 8, 2019
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