In-plane cyclic loading tests of concrete hollow block masonry walls retrofitted with high ductile fiber-reinforced concrete

Construction and Building Materials 238 (2020) 117758

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In-plane cyclic loading tests of concrete hollow block masonry walls retrofitted with high ductile fiber-reinforced concrete Mingke Deng a, Wei Zhang a,⇑, Ning Li b a b

School of Civil Engineering, Xi’an University of Architecture and Technology, Xi’an, Shanxi 710055, China Hualu Engineering & Technology Co., LTD. Xi’an, Shanxi 710065, China

h i g h l i g h t s
Reverse cyclic tests were performed on two types of concrete block walls.
Walls were retrofitted using single or double-sided HDC layers.
HDC layers can effectively reduce the damage of the masonry panels.
The seismic response of masonry walls is remarkably improved by HDC layers.

a r t i c l e

i n f o

Article history: Received 18 February 2019 Received in revised form 31 October 2019 Accepted 28 November 2019

Keywords: High ductile fiber-reinforced concrete (HDC) Concrete hollow block masonry walls Retrofitting In-plane cyclic loading Seismic behavior Ultimate shear strength

a b s t r a c t When subjected to horizontal loading induced by an earthquake, concrete block masonry walls generally exhibit vulnerability, particularly under high shear force. To improve the failure mode and lateral resistance of these walls, and to promote the application of concrete block units in multi-story masonry buildings, this study proposes techniques for retrofitting concrete hollow block masonry walls by installing coated high ductile fiber-reinforced concrete (HDC) layers. Six half-scale testing specimens, including three unreinforced concrete block masonry walls (URM) and three confined walls, were tested under in-plane cyclic loading to investigate their seismic performance before and after the upgrade. The test results indicate that these configurations significantly increased the lateral shear strength, stiffness, and energy dissipation capacity of the URM walls, and changed the brittle failure mode into a ductile mode. For the confined masonry walls, the lateral shear resistance and the stiffness were obviously enhanced, whereas the energy dissipation showed negligible improvement. The retrofitted walls showed higher residual strength and less damage, making them favorable for post-earthquake restoration. To estimate the ultimate shear strength of the tested specimens, relevant formulas developed in this study are presented. Ó 2019 Elsevier Ltd. All rights reserved.

1. Introduction Concrete blocks are widely used in building construction worldwide, owing to their lightweight, ease of installation, and energysaving qualities. However, despite accounting for a large part of the building stock, many concrete block masonry walls have exhibited low seismic performance because their outdated designs are not in compliance with recent masonry code provisions. As a result, such walls are vulnerable to severe damage and have posed a threat to human life, when responding to the force of earthquakes. Therefore, upgrading these walls with effective techniques is imperative to maintain safe working conditions and to improve ⇑ Corresponding author. E-mail address: [email protected] (W. Zhang). https://doi.org/10.1016/j.conbuildmat.2019.117758 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.

their seismic performance, which has contributed to the application of concrete blocks in multi-story buildings. Multiple techniques are used to improve the seismic performance of masonry walls, such as installation of externally bonded fiber-reinforced polymer (FRP) strips or laminates [1–4], shotcreting masonry panels in combination with steel wire mesh reinforcement [5,6], and other conventional retrofitting methods [7–9]. However, FRP used in retrofitting has exhibited poor bonding and anchoring between the FRP and the substrate; moreover, the construction processes are difficult under humid and low-temperature environments. Other techniques are expensive and timeconsuming and require abundant space. In addition, some retrofitting techniques can actually exacerbate the damage caused by earthquakes, because they add substantial mass to the structures. Thus, the present study proposes retrofitting structural members

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M. Deng et al. / Construction and Building Materials 238 (2020) 117758

with high ductile fiber-reinforced concrete (HDC) layers to address these limitations. High ductile fiber-reinforced concrete is a type of cement composite reinforced with synthetic fibers. It is also known as engineered cementitious composite (ECC) [10], strain-hardening cement-based composite (SHCC) [11], ultra-high toughness cementitious composite (UHTCC) [12], and high-performance fiber-reinforced cementitious composite (HPFRCC) [13]. HDC exhibits strain-hardening characteristics through the development of multiple fine cracks that are bridged by the stress in the fibers [14]. Owing to its superior mechanical properties, HDC has been used to strengthen and repair structural members [15–19] and is widely used in civil engineering [20]. This study investigates the seismic performance of concrete block masonry walls before and after being reinforced with HDC layers. The seismic behavior of the tested walls subjected to lateral in-plane cyclic loading, was analyzed to evaluate the failure modes, failure processes, hysteretic characteristics, deformation, and bearing capacities. In addition, corresponding formulas are proposed to predict the lateral resistance.

2. Experimental program 2.1. Test specimens The concrete masonry blocks could not be altered to customize the sample size owing to regional limitations. Thus, this experiment used three full-scale blocks with dimensions (length  height  thickness) of 390 mm  190 mm  190 mm, 190 mm  190 mm  190 mm, and 290 mm  190 mm  190 mm, respectively. Considering the laboratory restriction for the size of the test site, the walls were constructed as half-scale models with reduced-scale height and length but full-scale thickness. Murty and Jain [21] analyzed the influence of a reduced-scale masonry specimen constructed with full-scale bricks and indicated a reduction in the likelihood of crushing failure mode occurrence. The specimens were constructed on a precast reinforced concrete footing and were built to dimensions of 2400 mm  1500 mm  190 mm (length  height  thickness). The cast-in-place reinforced head beams and confined columns were fixed with cross-sections of 250 mm  190 mm and 200 mm  190 mm, respectively, after allowing the specimens to cure for approximately one week. The grade of the longitudinal reinforcements and the transverse stirrups for the confined columns was HPB300 with diameters of 8 mm and 6 mm, respectively, and the spacing of the stirrups was 200 mm. Details of the specimens are shown in Fig. 1, and the strengthened forms and the numbers of specimens are shown in Table 1.

2.2. Specimens retrofitting The walls were upgraded using the configurations of troweling 20 mm HDC layers on one side of the masonry interface or 15 mm HDC layers on both sides. The masonry surface required several treatments to improve the adhesion between the masonry substrate and HDC layers before coating the HDC layers. First, the mortar joints of the masonry walls were dug 8–10 mm by using a trowel. Then, the walls were cleaned with a wire brush and water to remove dust and lose particles, and an interfacial agent was applied on the masonry wall surface after that wall dried. Finally, after air-curing for about 12 h, the HDC layers were coated on the masonry wall surface. Once the strengthening treatments were completed, the retrofitted walls were cured for one week by being spraying with water twice daily.

Fig. 1. The details of test specimens: (a) unreinforced concrete block masonry (URM) walls; (b) confined concrete masonry wall.

Table 1 The retrofitted schemes of specimens. Specimen

The retrofitted schemes

W-0 W-1 W-2 WG-0 WG-1 WG-2

Unstrengthened URM wall Single-sided 20 mm HDC layer strengthened URM wall Double-sided 15 mm HDC layers strengthened URM wall Unstrengthened confined wall Single-sided 20 mm HDC layer strengthened confined wall Double-sided 15 mm HDC layers strengthened confined wall

2.3. Materials used in the experiment The main components of the HDC used in this research were cement, fly ash, water, river sand, water reducers, and polyvinyl alcohol (PVA) fibers (Fig. 2); the volume ratio of the PVA fibers was 2%. The tensile stress–strain curve of the HDC is shown in Fig. 3. All the samples used to determine the mechanical properties of the constitutive masonry materials were poured when the masonry walls were being constructed and were cured in the same environment that used for the constructed walls. After air-curing for approximately two months, the mechanical properties of the constitutive masonry materials were considered stable. The constructed specimens were then tested, as were the mechanical properties of the constitutive masonry materials; Table 2 summarizes the test results. The mean compressive strength (f1) of the concrete blocks was determined by using six samples (390 mm  190 mm  190 mm) based on GB/T 4111–2013 [22]; the tested value was 14.83 MPa. One type of mortar was selected as the construction material; its mean compressive strength (f2)

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2.4. Test setup, instrumentation, and loading procedures Fig. 4 shows the cyclic loading test setup and the instrumentation of the walls. A total vertical load of 365kN was applied to the rigid steel beam using two hydraulic jacks. The lateral force, which was applied to the head beam by a hydraulic actuator, was distributed uniformly to the masonry wall cross-section. The footing beam was fixed to the ground by four anchor bolts. The magnitude of the lateral force at the loading point of the wall top was measured using the load cell of the test setup. Four linear variable displacement transducers (LVDT) were employed to measure the horizontal displacement of the head beam and foundation, and that of the diagonal deformation of the specimen, respectively. The typical loading sequence, as shown in Fig. 5, consisted of a series of force and displacement-controlled cycles. First, horizontal cyclic loading was applied up to 20% of the predicted cracking load to check the normal operation of the test instrumentations. Then, the applied lateral load was gradually increased at 20kN intervals. After the yielding of the specimens, the loading scheme changed to the displacement-controlled stage. The displacement-controlled included loading at increments of 1 mm until the wall failed. Each amplitude included one complete cycle in the force-controlled stage and three cycles in the displacement-controlled stage. Finally, the test ended when the lateral force of the wall was reduced to 85% of its maximum lateral strength.

Fig. 2. Polyvinyl alcohol (PVA) fibers.

3. Experimental results 3.1. Failure modes and cracking distributions As shown in Fig. 6, the principal in-plane failure modes of URM walls consisted of shear failure, sliding failure, rocking failure, and toe-crushing failure [25,26]. The cracking patterns of all tested walls are shown in Figs. 7–12. The response in the two loading directions generally was not symmetrical; Table 3 presents the test results.

Fig. 3. Stress-strain curves of uniaxial tensile tests.

was determined by testing 36 mortar cubes of 70.7 mm3, which gave the value of 10.49 MPa. The average cubic compressive strength of the HDC (fcu, HDC) was 54.14 MPa, as determined by testing six HDC cubes of 100 mm3. Similarly, the average compressive strength of the concrete (fcu, m) used in the confined walls was tested to be 37.81 MPa. The mean yield stress of reinforcing steel (fy) was tested to be 395 MPa. On the basis of the experimental results, the masonry compressive strength (fm) and concrete tensile strength (ft) were be computed to be 9.03 MPa and 2.91 MPa, respectively, according to the Eq. (1) of GB 50003–2011 [23] and Eq. (2) of GB 50010–2010 [24]. 0:9

f m ¼ 0:46f 1 ð1 þ 0:07f 2 Þ ft ¼

ð1Þ

0:55 0:395f cu; m

ð2Þ

3.1.1. Specimen of URM wall (W-0) Specimen W-0 was a plain masonry panel dominated by shear failure. When the lateral force reached 100kN for the pulling direction, a flexural crack first formed at the bottom of the masonry panel. Stepped cracks appeared at the center of the wall panel until the lateral force reached 200kN for the pushing direction. Then, the loading procedure changed to the displacement-controlled stage. When the top displacement reached 3 mm for the pushing direction, diagonal cracks occurred in the masonry units at the wall toe. In addition, the wall achieved the maximum lateral force, and the existing cracks widened and were propagated in a slanted direction crossing through the masonry blocks and mortar joints. At the top displacement of 5 mm, the diagonal cracks were fully developed. The maximum diagonal crack width reached approximately 6 mm, and local spalling occurred at the compressive region. Finally, the wall assumed the mixed failure modes of shear and toe-crushing failure, as shown in Fig. 7. 3.1.2. Strengthened URM wall (W-1) Specimen W-1 was retrofitted by using a single-sided HDC layer. The HDC layer restricted the development of the diagonal

Table 2 Properties of constitutive masonry materials. Material tests

f1

f2

fcu,

Values (MPa)

14.83

10.49

54.14

HDC

ftu,

HDC

5.13

fcu,

m

37.81

Note: f1: compressive strength of the concrete block; f2: compressive strength of the mortar; fcu, HDC: cubic compressive strength of the HDC; ftu, HDC; fcu, m: cubic compressive strength of the concrete used in the confined walls; fy: yield strength of the reinforcement.

HDC:

fy 395 The tensile strength of

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M. Deng et al. / Construction and Building Materials 238 (2020) 117758

Fig. 4. Test set-up and instrumentation.

Fig. 5. Loading procedure.

cracks, which effectively reduced the damage of the masonry panel. For the pushing direction, a horizontal crack appeared at the bottom bed-joint of the masonry panel when the lateral force reached 120kN. After loading to 260kN, the displacementcontrolled state implemented the loading procedure. Stepped cracks first formed around the center of the masonry substrate when the top deflection reached 2 mm for the pushing direction. At displacement of 3 mm for the pushing direction, the wall experienced its ultimate lateral resistance, and multiple diagonal cracks appeared on the masonry panel. When the top deflection reached 4 mm for the pulling direction, fine diagonal cracks appeared on the HDC surface, and the horizontal cracks occurring at the bottom bed-joint of the masonry panel penetrated the entire wall length. The test ended when the loading beam reached 7 mm; the failure of specimen W-1 was ductile, in association with shear/flexural and sliding failure modes. As shown in Fig. 8, the cracking pattern showed several diagonal cracks in the masonry panel and the HDC layer as well as a penetrating horizontal flexural crack along the entire wall length. 3.1.3. Strengthened URM wall (W-2) Specimen W-2 was retrofitted using the double-sided HDC layers. A horizontal crack first formed at the bottom bed-joint of the masonry panel when the lateral force reached 180kN for the pulling direction. After loading of 280kN, displacement controlled the horizontal cyclic loading. When the top deflection reached 1 mm,

Fig. 6. In-plane failure modes: (a) shear failure; (b) sliding failure; (c) rocking failure.

the flexural cracks occurring at both loading directions extended to the mid-length of the wall. The specimen achieved the ultimate lateral resistance at the top deflection of 4 mm for the pushing direction, and the masonry panel began to slide. As the loading continued, the horizontal crack became fully developed and extended throughout the length of the wall; however, no diagonal cracks appeared on the HDC surface. Finally, the wall failed when the top deflection reached 9 mm and reached the sliding failure mode, as shown in Fig. 9.

M. Deng et al. / Construction and Building Materials 238 (2020) 117758

Fig. 7. The failure mode of specimen W-0.

5

Fig. 9. The failure mode of specimen W-2.

3.1.4. Confined wall (WG-0) Specimen WG-0 was a reference specimen of the confined masonry walls that reached the shear failure mode. When the lateral force reached 280kN for the pulling direction, a stepped crack first occurred from the upper corner to the masonry center. Then, the loading procedure changed to the displacement-controlled stage. As the loading continued, horizontal cracks appeared at the center and bottom of the confined columns, and multiple diagonal cracks appeared on the masonry panel and propagated along the masonry joints in a slanted direction. At the top deflection of 7 mm for the pushing direction and 9 mm for the opposite direction, the specimen achieved the ultimate lateral resistance. Finally, the test ended at the top deflection of 12 mm. As shown in Fig. 10, the wall showed a large amount of diagonal cracking, and a major diagonal crack penetrated the confined column. 3.1.5. Strengthened confined wall (WG-1) Specimen WG-1 was strengthened by using a single-sided HDC layer. At a load of 280kN for the pushing direction, a horizontal crack first appeared at the bottom of the confined columns. When the lateral force reached 320kN for the pulling direction, a diagonal crack formed on the masonry surface, and the loading procedure changed to the displacement-controlled state. Fine diagonal cracks appeared on the HDC layer when the top deflection reached 2 mm. At displacements of 3 mm for the pushing direction and 4 mm for the opposite direction, the wall reached the ultimate lateral resistance. Cyclic loading continued, multiple fresh diagonal cracks appearing on the masonry surface and the HDC layer, and partial spalling in the HDC layer occurred at the wall center. Finally, the wall loading ended at the top deflection of 11 mm, and the shear failure mode was reached, as shown in Fig. 11.

Fig. 10. The failure mode of specimen WG-0.

3.1.6. Strengthened confined wall (WG-2) Specimen WG-2 was strengthened by using the double-sided HDC layers. When the lateral force reached 360kN for the pushing direction, a horizontal crack first appeared at the bottom of the wall. After loading of 460kN, displacement controlled the loading procedure. At the top deflection of 2 mm for the pulling direction, the wall achieved the ultimate lateral resistance, and a few diagonal hairline cracks formed in the HDC layer. With an increase in lateral displacement, the horizontal crack widened and propagated along the wall length; no new diagonal cracks emerged in the HDC layers. When the specimen was near failure, the wall began

Fig. 8. The failure mode of specimen W-1.

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M. Deng et al. / Construction and Building Materials 238 (2020) 117758

Fig. 11. The failure mode of specimen WG-1.

Fig. 12. The failure mode of specimen WG-2.

to slide along the horizontal crack. At a displacement of 8 mm, the wall loading ended, and rocking and sliding failure modes were reached. The cracking pattern is shown in Fig. 12.

3.2. Hysteretic behavior The hysteretic curves and responses of the tested specimens are shown in Fig. 13. Before cracking, the load–displacement curve of specimen W-0 hardly changed, essentially remaining in a straight line with negligible residual displacement. After the first diagonal crack appeared in the masonry panel, the specimen entered the stage of elastic– plastic deformation, and the area of hysteretic loops widened with the loading process. At a peak load of 320.37kN for the pushing direction, inclined cracks appeared at the compressive region. Afterward, the cracks developed significantly, which led to local spalling at the wall toes. Therefore, the hysteretic curve of specimen W-0 exhibited a sharp decline, dropping to approximately 85% of the lateral resistance. At the later loading stage, the wall was damaged severely, and the residual displacement averaged for the two loading directions increased to about 3.6 mm. For specimen W-1, the hysteretic curve was approximately in a linear elastic state before cracking. The pinching behavior of the hysteretic loops started at a load of 260kN. Then, the residual displacement increased distinctly with the loading. The HDC layers were able to reduce the damage of the wall, and failure mode was converted to a ductile mode. Therefore, the retrofitted specimen resisted the lateral force for several cycles with little strength degradation after a peak load of 374.39kN for the pushing

direction. Compared with specimen W-0, the numbers of hysteretic loops of specimen W-1 obviously increased, and the residual displacement averaged for the two loading directions decreased obviously to approximately 2.7 mm. Specimen W-2 reached the sliding failure mode, and its hysteretic curve showed the following characteristics. The wall showed a linear elastic response during the initial loading state. The pinching behavior of the hysteretic loops occurred when the horizontal crack extended to about mid-length of the wall. Before the sliding occurred, the wall showed nonlinear behavior with small residual displacement. After a peak load of 371.31kN for the pushing direction, the specimen slid, and the residual displacement of the wall increased continuously. Similar to that shown by specimen W-1, specimen W-2 resisted the lateral force for several cycles with little strength degradation. The numbers of hysteretic loops were higher for specimen W-2 than those for specimen W1, which reflects greater energy dissipation. For specimen WG-0, the pinching behavior was triggered until the first diagonal crack appeared in the masonry panel. Then, as the loading continued, the area of the hysteretic loops increased accordingly. After a peak load of 393.72kN for the pushing direction, the residual displacement increased gradually. Essentially, the confined columns improved the integrity of specimen WG-0, which contributed to reduced residual displacement, and the hysteretic loops were thicker and more stable at later loading stages. This specimen had the same failure mode as specimen WG-1 and analogous hysteretic loops; however, the maximum lateral resistance of specimen WG-1 was higher than that of the specimen WG-0. Specimen WG-2 reached the rocking failure mode, and its hysteretic curves had the following characteristics. The hysteretic curves were maintained in a straight line at the initial loading state. Under reverse cyclic loading, the wall resumed the original position with the opening and closing of the horizontal crack; therefore, the residual displacement for both loading directions was small. Before the wall slid, its hysteretic loops were narrow and stable. When it approached the failure state, the specimen slid with only one cycle to resist the lateral load, and its residual displacement increased suddenly to about 5 mm averaged for the two loading directions. The area of the hysteretic loops increased significantly after the sliding of the wall. 3.3. Skeleton curves and characteristic points Fig. 14 shows envelopes of hysteretic loops of the specimens, and Table 3 presents the characteristic points obtained from the skeleton curves. As shown in Table 3, the retrofitting techniques increased the lateral resistance of the strengthened walls by 19% for specimen

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M. Deng et al. / Construction and Building Materials 238 (2020) 117758 Table 3 Test results at characteristic points. Specimen

Cracking state

Maximum state

Failure state

Failure mode

Pcr/kN

Dcr/mm

Pu/kN

Average

Du/mm

Average

Pd/kN

Average

Dd/mm

Average

W-0

100

0.30

310.86

Shear and crushing

6.23

Shear/flexural and sliding

W-2

180

0.38

8.07

Sliding

WG-0

280

0.94

11.52

Shear

WG-1

+280

+0.59

10.25

Shear

WG-2

+360

+0.32

+3.76 5.09 +6.49 5.96 +7.57 8.57 +11.45 11.58 +10.56 9.93 +7.62 7.72

4.42

+0.22

+272.31 256.15 +318.23 310.34 +315.61 288.66 +334.66 336.45 +384.69 385.20 +450.15 417.01

264.23

+120

+2.92 3.99 +3.04 5.02 +4.07 6.00 +7.05 7.97 +3.01 4.03 +1.50 2.00

3.46

W-1

+320.37 301.35 +374.39 365.10 +371.31 339.61 +393.72 395.82 +452.57 455.17 +529.59 490.59

7.67

Rocking and sliding

369.75 355.46 394.77 453.87 510.09

4.03 5.04 7.51 3.52 1.75

314.28 302.13 335.56 384.95 433.58

Note: (+) presents the pushing loading direction; (-) presents the pulling loading direction; Pcr: cracking load; Dcr: displacement corresponding to cracking load; Pu: ultimate load; Du: displacement corresponding to ultimate load; Pd: failure load; Dd: displacement corresponding to failure load.

Fig. 13. Load-displacement hysteretic curves.

W-1 and 14.3% for specimen W-2. The ultimate shear strength of specimen W-1 was higher than that of specimen W-2, probably owing to the discreteness of the wall materials. The HDC layers improved the stiffness of the walls, and the confined columns restricted the sliding of the masonry panel. Therefore, when

loading to the ultimate load and failure load, the deformation of specimens WG-1 and WG-2 decreased to some degree compared with that shown by specimen WG-0. However, the displacement of the retrofitted URM walls increased because the masonry panel slid. The diagonal shear capacity of the URM wall retrofitted with

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M. Deng et al. / Construction and Building Materials 238 (2020) 117758

Fig. 15. Energy dissipation. Fig. 14. Skeleton curves.

HDC layers increased significantly, whereas the increment of the flexural capacity at the bottom bed-joint was comparatively small. Therefore, the horizontal crack at the bottom bed-joint of the masonry panel was able to develop fully to penetrate the entire wall length. As described in the literature [27,28], the formation of horizontal cracks at the bed joints can cause the masonry panels to slide along the wall length. Specimen W-1 and W-2 showed ultimate displacements of 1.41 and 1.83 times larger than that of the reference specimen, respectively; in addition, their lateral resistance increased by 15% and 29%, whereas their ultimate displacement decreased by 11% and 33%, respectively. As shown in Fig. 14, the residual shear strength of the strengthened walls was close to the peak loads of the controlled walls, and if loaded these walls could continue to resist deformation. This can partly explain why the ultimate displacement of specimens WG-1 and WG-2 decreased. The HDC layers enhanced the initial stiffness of the strengthened confined walls; as a result, the skeleton curves of the retrofitted specimens had smaller deformations than the controlled walls before the peak load. After the peak load, the HDC layers provided a constraint for the masonry wall, which effectively reduced the damage, and restricted the strength degradation. Therefore, the skeleton curves of the retrofitted specimens had a long, steady descent trajectory. 3.4. Energy dissipation As discussed in previous research [28], the energy dissipation of the walls consisted of friction along joints, the formation of new cracks, fracturing in the masonry units, and rupture of the HDC

layers. The amount of dissipation for the different specimens at different levels was determined by calculating the area enclosed by the corresponding hysteretic loops. As shown in Fig. 15, the cumulative energy dissipation of the third loop at each displacement level was calculated based on the hysteretic curves. Table 4 shows the cumulative energy dissipation of the walls at peak loads as well as the ultimate displacement. For the URM walls, the failure modes of the retrofitted specimens were ductile; therefore, the total accumulated energy dissipation of the strengthened specimens was higher than that of the reference specimen. Compared with specimen W-0, the total energy dissipation of specimens W-1 and W-2 improved by 56% and 111%, respectively. As shown in Fig. 15, at the same displacement, the accumulated energy dissipation of the strengthened specimens was lower than that of the reference specimen. This occurred because the HDC layers improved the failure mode of the test specimens; thus, when loading to the same displacement, the reference specimen was damaged severely and dissipated a substantial amount of energy. However, the strengthened specimens showed only slight damage. Compared with specimen W-0, the energy dissipation capacity of specimen WG-0 showed a considerable increase owing to the effect of the confined columns. The increased stiffness caused relatively little deformation in specimens WG-1 and WG-2, which in turn restricted the exertion of reinforced bars to dissipate the energy. Thus, the total accumulated energy dissipation of specimens WG-1 and WG-2 increased 12% and decreased 11%, respectively, compared with that of specimen WG-0. This reflects negligible improvement. In addition, the residual shear strength

M. Deng et al. / Construction and Building Materials 238 (2020) 117758

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Table 4 Cumulative energy dissipation of specimens. Specimen

Cumulated energy/kNmm Peak load (kN)

Ultimate displacement (mm)

W-0 W-1 W-2 WG-0 WG-1 WG-2

2661 4866 4926 15,006 4736 2355

12,587 19,636 26,558 35,143 39,210 31,310

of the retrofitted walls was high, which indicates that the energy dissipation capacity of the strengthened walls did not reach its maximum. Under the reverse cyclic loading test, the confined wall before and after retrofitting using HDC layers had less residual deformation, and the retrofitted wall had greater stiffness than that in the reference specimen. Therefore, the accumulated energy dissipation of the strengthened specimens at the same displacement was higher than the reference wall. 3.5. Hysteretic damping Hysteretic damping can be evaluated by using the coefficient of equivalent viscous damping (n), which is expressed by the relationship between the energy (DEhyst) dissipated in one cycle of loading and the input potential energy (Ep) in the same cycle [29]. The coefficient of equivalent viscous damping (n) is calculated by using Eq. (3); the specific calculations of the energy dissipation (DEhyst) and the input potential energy (Ep) are discussed in the literature [29].

n¼

DEhyst 2pEp

ð3Þ

According to the results of previous research [30,31], three limit states were defined to further analyze the energy dissipation in the test walls: cracking limit, ultimate limit, and failure limit. The cracking limit represents the state at which cracks first appear; the ultimate limit represents the state at which the specimen achieves the maximum load; and the failure limit represents the position at which the specimen reaches the ultimate displacement. Table 5 summarizes the coefficient of equivalent viscous damping (n) of the test specimens. As shown in Fig. 13, specimens W-0 and WG-2 produced large residual deformation at the failure limit; therefore, the area of their hysteretic loops was larger. In addition, the energy dissipation (DEhyst) increased significantly, which contributed to the high coefficient of equivalent viscous damping (n). For the URM walls after retrofitting, the coefficient of equivalent viscous damping (n) was smaller than that in the reference specimen at the ultimate limit and failure limit states, whereas, the coefficient of equivalent viscous damping (n) was higher than that in the reference specimen for the confined walls. This occurred likely because the HDC layers transformed the failure mode of the test specimens and changed the area of the hysteretic loops. Therefore, it can be inferred that the coefficient of equivalent viscous damping (n) was likely influenced by the failure modes.

Fig. 16. Stiffness degradation curves.

3.6. Stiffness degradation Based on the skeleton curves and the data obtained in the experiments, the lateral secant stiffness can be expressed by the following relationship to represent the laws of stiffness degradation:

Ki ¼

jþF i j þ jF i j jþDi j þ jDi j

ð4Þ

where F i is the positive and negative lateral peak load at the first loop of each loading stage, and Di represents the corresponding displacement at the peak load under each loading stage. The stiffness degradation can reflect the level of damage in walls subjected to in-plane lateral cyclic loads. The stiffness degradation curves are shown in Fig. 16. 4. Bearing capacity

Table 5 Equivalent viscous damping. Specimen

W-0 W-1 W-2 WG-0 WG-1 WG-2

4.1. Shear capacity of the URM walls

The coefficient of equivalent viscous damping (n) Cracking limit

Ultimate limit

Failure limit

0.076 0.063 0.073 0.083 0.069 0.091

0.149 0.101 0.078 0.084 0.115 0.135

0.314 0.229 0.142 0.088 0.110 0.226

As described in the literature [27], the cracked section of URM walls caused by flexure hardly contributed to the ultimate shear strength. The length of the uncracked section is calculated by neglecting the tensile strength of the bed-joints and assuming a simplified distribution of compression stress, which is most commonly constant or linear. Therefore, Magenes and Calvi [27] proposed the formula in Eq. (5) to predict the ultimate shear

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M. Deng et al. / Construction and Building Materials 238 (2020) 117758

strength of URM walls. In the present study, this formula was arranged to predict the ultimate lateral resistance of the test specimens as follows:

Vm ¼

1:5f v0 þ lry Am 1 þ 3f v0 av =ry

l ¼ 0:83  0:7

ry fm

ðry =f m 6 0:8Þ

ð5Þ ð6Þ

4.2. Shear capacity of the confined walls Tomazevic and Klemenc [32] predicted the lateral shear strength of confined walls by considering the contribution of both the masonry panel and the confined columns. They proposed the following equation to calculate the shear capacity of the confined walls:

ð7Þ

where Vwall is the total lateral resistance of confined masonry wall, Vpanel is the lateral resistance of the masonry panel, and Vties is the lateral resistance of the confined columns. Vpanel can be calculated by using Eq. (5). Abdelkrim [33] and Bourzam [34] considered that a certain amount of shear is transmitted by the dowel action of the vertical bars when the diagonal crack appears at the columns. The confined columns in this research showed only horizontal cracking. Therefore, the shear capacity of the confined columns is attributed to the tensile action of the concrete and longitudinal reinforcements and can be computed by using Eq. (8) [35]:

V tie ¼ 0:1Ay f y þ 0:7wAc f t

ð8Þ

where Ay is the cross-sectional area of the longitudinal reinforcement bars; Ac is the cross-sectional area of the tie column; w is the shear strength reduction coefficient of the concrete columns, w = 0.85; ft is the tensile strength of the concrete, and fy is the yield stress of reinforcing steel. The lateral resistance of the strengthened walls was predicted by using the following formula.

V u ¼ V þ V HDC

ð9Þ

where VHDC is the shear capacity of the HDC layer, and V is the shear capacity of the masonry wall. At present, no specific formula is used to predict the contribution of the HDC layer to the lateral resistance of masonry walls retrofitted using HDC layers. The conventional relation of Eq. (10) provided in ACI-318 (ACI 2008) [36] was arranged to calculate the shear strength of the reinforced concrete layers that were used to retrofit the masonry walls. In the present study, this formula was employed to predict the shear strength of the HDC layers:

qffiffiffiffiffiffiffiffiffiffiffiffiffi V HDC ¼ 0:17k f c; HDC bh

ð10Þ

where fc, HDC is the compressive strength of the HDC material, fc, = 0.8fcu, HDC [37]; k is a modification factor (k = 1.0); and h and b are the thickness and the length of the HDC layer, respectively. The calculated ultimate shear strength values of the tested walls are summarized in Table 6. HDC

Specimen

Vi/kN

Pi/kN

Pi/Vi

W-0 W-1 W-2 WG-0 WG-1

277.79 331.49 358.34 378.45 432.15

310.86 369.75 355.46 394.77 453.87

1.11 1.12 0.99 1.04 1.05

Note: Vi: calculated result; Pi: tested results.

where Vm represents the ultimate lateral resistance of the unreinforced masonry wall; fv0 represents the shear strength of the masonry wall, fv0 = 0.069 f1/2 [23]; ry is the vertical compressive 2 stress; Am is the cross-sectional area of the masonry panel; and av is the geometry aspect ratio (av = h/l). The friction coefficient l can be calculated using formula (6) of the China code for the design of masonry structures [23].

V wall ¼ V panel þ V ties

Table 6 Ultimate shear strength of tested walls.

5. Conclusions In this research, six walls (including three URM walls and three confined walls) were tested under an in-plane cyclic loading system, before and after retrofitting with high ductile fiberreinforced concrete (HDC) layers, to assess the effectiveness of the retrofitting techniques. The mechanical behavior of the tested walls was evaluated based on the failure mode, hysteretic behavior, stiffness degradation, and lateral shear resistance. The conclusions drawn in this research are summarized in the following points. 1. For URM walls, the retrofitting techniques effectively restricted the generation of stepped cracks and converted the brittle failure mode to a ductile mode. Specimens W-1 and W-2 reached the sliding failure mode and were damaged slightly. Therefore, measures employed to resist the sliding of the walls would further exploit the effect of the HDC layer and improve the seismic performance of retrofitted specimens. For the confined masonry walls, the HDC layer of specimen WG-1 ruptured to dissipate the energy, which effectively reduced the damage in the masonry substrate. The specimens strengthened with doublesided HDC layers experienced sliding failure and rocking failure modes, which indicates that these walls are desirable for postearthquake repair. 2. Retrofitting techniques including trowelling 20 mm into the HDC layer on the one-sided masonry surface and 15 mm on both sides of the double-sided masonry surface can increase the lateral resistance by 14%–19% for URM walls and 15%–29% for confined walls. The URM walls showed improved energy dissipation after retrofitting, with an increment of 56%–111%. However, for the confined walls, the energy dissipation was not improved significantly, probably because the maximum performance of the retrofitted walls was not reached. 3. The residual shear strength of the retrofitted specimens was close to (or even higher than) the ultimate shear strength of the original walls, which illustrates that the retrofitted specimens can continue to bear loads when the lateral forces of the walls reduce to below 85% of their ultimate lateral resistance. 4. The retrofitted URM walls had more hysteretic loops, which indicates that the HDC layers can improve the seismic performance of URM walls. The skeleton curves for all of the retrofitted specimens had a long steady descent trajectory, which demonstrates that the HDC layers can improve the capacity of damage tolerance. 5. The secant stiffness of the retrofitted specimens was higher than that of the controlled specimens at the same displacement, which confirms that the retrofitting technique can reduce wall damage. 6. Based on previous studies [27,32,35,36], the present study provided design equations for evaluating the ultimate shear strength of the tested walls. The calculated results were conservative and matched well with the tested results, although the accuracy of these formulas requires further verification owing to the lack of substantial test data. Therefore, additional relative studies should be conducted to complete this hypothesis.

M. Deng et al. / Construction and Building Materials 238 (2020) 117758

7. The formulas proposed in this study aimed to calculate the lateral resistance of the masonry walls with the shear failure mode. For the masonry walls with flexural failure mode, the corresponding formulas were not proposed due to the short of tested specimens.

Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper. Acknowledgments The research was funded by the National Natural Science Foundation of China (No. 51878545). Special thanks are given to prof. Mingke Deng and students who make great efforts to this project. Appendix A. Supplementary data Supplementary data to this article can be found online at https://doi.org/10.1016/j.conbuildmat.2019.117758. References [1] P. Roca, G. Araiza, Shear response of brick masonry small assemblages strengthened with bonded FRP laminates for in-plane reinforcement, Constr. Build. Mater. 24 (8) (2010) 1372–1384. [2] M.R. Valluzzi, D. Tinazzi, C. Modena, Shear behavior of masonry panels strengthened by FRP laminates, Constr. Build. Mater. 16 (7) (2002) 409–416. [3] H. Santa-Maria, P. Alcaino, Repair of in-plane shear damaged masonry walls with external FRP, Constr. Build. Mater. 25 (3) (2011) 1172–1180. [4] N.G. Shrive, The use of fibre reinforced polymers to improve seismic resistance of masonry, Constr. Build. Mater. 20 (4) (2006) 269–277. [5] F.V. Karantoni, M.N. Fardis, Effectiveness of seismic strengthening techniques for masonry buildings, ASCE J. Struct. Eng 118 (7) (1992) 1884–1902. [6] M.A. ElGawady, P. Lestuzzi, M. Badoux. Seismic behavior of URM walls retrofitted using Shotcrete. Proceedings of New Zealand Society of Earthquake Engineering Annual Conference. No. CONF. 2006. [7] M. ElGawady, P. Lestuzzi, M. Badoux, A review of conventional seismic retrofitting techniques for URM, in: 13th international brick and block masonry conference. 2004, pp. 1-10. [8] L. Binda, C. Modena, G. Baronio, S. Abbaneo, Repair and investigation techniques for stone masonry walls, Constr. Build. Mater. 11 (3) (1997) 133– 142. [9] M. Ashraf, A.N. Khan, Naseer A, et al., Seismic behavior of unreinforced and confined brick masonry walls before and after ferrocement overlay retrofitting, Int. J. Archit. Herit. 6(6) (2012) 665-688. [10] V.C. Li, S. Wang, C. Wu, Tensile strain-hardening behavior of polyvinyl alcohol engineered cementitious composite (PVA-ECC), ACI Mater. J.-American Concrete Institute 98 (6) (2001) 483–492. [11] G.P.A.G. van Zijl, F.H. Wittmann, B.H. Oh, et al., Durability of strain-hardening cement-based composites (SHCC), Mater. Struct. 45 (10) (2012) 1447–1463. [12] H. Li, S.L. Xu, C.K.Y. Leung, Tensile and flexural properties of ultra high toughness cemontious composite, J. Wuhan. Univ. Technol.-Mater. Sci. Ed. 24 (4) (2009) 677–683.

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