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Seismic behavior of masonry-infilled precast concrete frames considering effects of opening
Construction and Building Materials 211 (2019) 756–770
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Construction and Building Materials journal homepage: www.elsevier.com/locate/conbuildmat
Seismic behavior of masonry-infilled precast concrete frames considering effects of opening Lei Wang a, Zhen-Yun Tang b, Yue Li c, Kai Qian d,⇑ a
College of Civil Engineering and Architecture, Guilin University of Technology, Guilin 541004, China The Key Laboratory of Urban Security and Disaster Engineering of Ministry of Education, Beijing University of Technology, Beijing 100124, China c Department of Civil Engineering, Case Western Reserve University, Cleveland, OH, USA d School of Civil and Environmental Engineering at Nanyang Technological University, Singapore b
h i g h l i g h t s Eight 1/2 scale, single-bay, single-storey RC and PC frames were tested. The effects of proposed PC construction technology were evaluated. The strengthening effects of SFRC at the cast-in-place connections were quantified. The reliability of analytical model for predicting the capacity of RC and PC frames was evaluated.
a r t i c l e
i n f o
Article history: Received 30 December 2018 Received in revised form 22 March 2019 Accepted 23 March 2019
Keywords: Precast concrete Reinforced concrete Infilled masonry wall Seismic Experimental study
a b s t r a c t It is commonly accepted that addition of infilled masonry walls (IMW) could enhance the in-plane strength and stiffness of RC frame buildings subjected to seismic loads. However, once the IMW walls failed, the shear force initially resisted by the IMW has to be transferred to the surrounding frames and may incur shear failure to the frames. Despite the extensive studies carried out in the past decades, studies on the structural performance of masonry infilled precast concrete (PC) frames are few. In this study, eight half-scaled single-story one bay frame were tested subjected to seismic loading, including three reinforced concrete (RC) frames and five PC frames. The investigated PC frames deviates the beam potential plastic hinge zone away from beam-to-beam connection region. The variables investigated in the tests included the types of opening in masonry infill walls and the types of cast-in-place concrete at the connections (normal concrete or steel fiber reinforced concrete). Results indicated that the investigated PC construction method has minimal effects on inelastic behavior and ductility of bare and infilled frames. However, the PC construction method may affect the crack development, especially at the interfaces in between precast and cast-in-place concrete surfaces in the region of the connection. However, the load resisting capacity and ductility was insensitive to the usage of steel fiber reinforced concrete at the connections, especially for bare frames. Ó 2019 Elsevier Ltd. All rights reserved.
1. Introduction Infilled masonry walls (IMW) are frequently utilized as cladding or partitions in reinforced concrete (RC) and steel moment resisting frames. Although it is commonly accepted that IMW may interact with the surrounding frames under earthquake load, they are often ignored in design practice. Though as secondary members, it has also been generally accepted that the IMW may affect the stiffness and strength of the infilled frames and change the seismic
⇑ Corresponding author. E-mail address: [email protected] (K. Qian). https://doi.org/10.1016/j.conbuildmat.2019.03.287 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.
response of the infilled frame, as observed in post-earthquake investigation [1,2]. However, in highly seismic regions, ignoring the IMW and frames interaction may not always be safe as the enhancement of stiffness as IMW may change the dynamic response of the frame significantly. The presence of IMW may result in unexpected force distribution and cause local damage in the frame [3]. Regarding the complexity of the interactions, extensive studies have been carried out experimentally or analytically. Since 1950s, a number of experimental tests [4–9] have been carried out to evaluate the seismic behavior of RC frames with IMW. These tests indicated that IMW could improve the
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L. Wang et al. / Construction and Building Materials 211 (2019) 756–770 Table 1 Property of test specimens. Test ID
Dimensions 2
PC PCS PCW PCWD PCWDS
Joint Trans. Rebar
Connection Type Beam-to-Beam
Column-to-Column
0.2% 0.2% 0.2% 0.2% 0.2%
Overlapping Overlapping Overlapping Overlapping Overlapping
Mechanical Mechanical Mechanical Mechanical Mechanical
2
Beam (mm )
Column (mm )
130 230 130 230 130 230 130 230 130 230
250 250 250 250 250 250 250 250 250 250
performance of RC frames significantly in accordance with the strength and energy dissipation capacity. Moreover, they were indicated that the number of bays will affect the failure mode, peak and residual capacity, and shear stress distribution of the frames significantly. Moreover, Kakaletsis and Karayannis [10,11], Mohammadi and Nikfar [12] studied the influences of openings on the seismic behavior of masonry infilled RC frames. They were found that the openings should be located as close to the edge of the panel as possible to achieve an improvement in the behavior of the infilled frames. The reduction factor of the ultimate strength of infilled frames with openings depends highly on the material of the confining frame, but the reduction factor of stiffness does not. Moretti et al. [13] and Jiang et al. [14] experimentally investigated the influence of connection between IMW and the bounding frame. The test results indicated that IMW should be connected to the frame through dowels. Including the infill walls, which were rigidly connected to the frame, could increase the strength, stiffness, and energy-dissipation capacity significantly. However, the displacement-based ductility decreased considerably. Furthermore, the possible failure modes of infilled masonry frames (corner crushing, diagonal compression failure, sliding shear failure, diagonal cracking failure, and frame failure) were summarized by tests [3,15,16]. In the past decades, researchers found simplified methods to simulate the lateral action of the IMW. The most commonly used method is diagonal strut method, which is diagonally connected the opposite compression corners of the frame [17,18]. Smith and Carter [19] indicated that the width of the strut is associated with the contact length between the frame and IMW. An analytical model is proposed to determine the equivalent width through a parameter k; which is expressed as the relative stiffness of the frame to that of infill panel and as shown in Eq. (1):
k¼
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4 ðEm t inf sin2hÞ 4Ef Icol hinf
ð1Þ
where Em ; t inf and hinf are infill’s Young’s modulus of elasticity, thickness and height, respectively; Ef Icol is the flexural rigidity of
ah ¼
Sleeve Sleeve Sleeve Sleeve Sleeve
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4Ef Icol hinf 4 2 Em tinf sin2h
p
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4Ef Ib Linf 4 aL ¼ p Em t inf sin2h
Infilled Walls
Concrete in the Connections
N/A N/A Full Wall Opened Wall Opened Wall
Normal SFRC Normal Normal SFRC
ð4Þ
ð5Þ
In the preceding two decades, researchers found that singlestrut model cannot simulate the complex behavior (bending moment or shear force) of the masonry infilled frames well, although it can fairly accurately to simulate the general behavior of infilled frames [23,24]. Therefore, multiple-strut models were proposed by researchers [16,25–32]. PC structures have advantages of high productivity, easy quality control, and cost efficiency. In the developed countries, such as U. S., Japan, Singapore, and New Zealand, PC systems with different connection types are widely implemented in the moment resisting frames. However, the PC system still has a limit in its applications in high seismic areas mainly due to several issues still unsolved. In particular, the reliable behavior of beam-column joints is a critical issue restricting the application of PC frames. The post-earthquake investigation has indicated that the collapse of PC structures is due to inadequate design of the beam-column joints [33]. So far, the design of PC system is normally required to emulate the behavior of cast-in-place RC system [34–37]. Current practice in PC frame construction indicated that the connections between precast beams are commonly located at the beam-column joint region (refer to Fig. 1). However, it should be noted that connecting the beams at the joint region is unfavorable because it disturbs the continuity of the reinforcements. Moreover, the congestion of reinforcements in the joints many bring difficulties in erection stage. Therefore, some researchers [38–40] proposed beam-to-beam connection at the midspan of the beam (refer to Fig. 2a). The connection types include 180-degree overlapping hooks, 90-degree double hooks or welding. However, as the T-shaped or cruciform precast components are very heavy and difficult to transport due to their large dimensions, applying this method in practice is
the columns; and h ¼ tan1 hLinf is the angle whose tangent is the inf
infill’s height-to-length aspect ratio. Based on experimental and analytical results, Mainstone and Weeks [20] proposed an empirical equation to determine the equivalent width of the strut as shown in Eq. (2), which was adopted by FEMA [21]. The parameter k is given in Eq. (1).
w ¼ 0:175ðkhcol Þ
0:4
ð2Þ
rinf
In CSA S340.1-04 [22], the equivalent strut width, w; could be determined using the vertical, ah ; and horizontal contact, aL ; length between the strut and the frame, h ¼ tan1 hLinf is the angle whose inf
tangent is the infill’s height-to-length aspect ratio
w¼
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi a2h þ a2L
ð3Þ
Fig. 1. Assembling method by connecting the beam reinforcements at the joint core.
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Fig. 2. Assembling configurations (a) midspan connection (Park [38]), (b) relocated the connection at a distance away from the column face, proposed by Khoo et al. [43] and used in this study.
difficult especially for the frames with long beam span [41]. To solve this problem, French et al. [42] proposed a new method, in which the connections were relocated at the beam span at a distance away from the column faces (refer to Fig. 2b). In this method, the beam-column joint with short, protruding beam stubs were cast with the column, subsequently, the beam spanning in between columns would form another precast member. Khoo et al. [43] adopted similar idea as French et al. [42] but the 180degree hooks or 90-degree double hooks were used for reinforcement lap splice. However, in these studies, the beam-column joints or beam-to-beam connections were the main concerns. The effects of column-to-column connection on seismic behavior of PC frames were ignored. In this paper, to re-exam the efficiency of proposed beam-tobeam connection proposed by Khoo et al. [43] and to evaluate the global behavior of PC frames subjected to simulated seismic load, a series of PC and RC single-storey frames were tested. Moreover, the effects of IMW with or without openings on PC frames were also investigated. 2. Experimental program 2.1. Test specimens As tabulated in Table 1, eight single-storey, single-bay, 1/2 scaled planar frame specimens were tested in this study. These geometrically identical eight specimens were categorized into two groups based on construction methods: RC-series and PC-
series. However, based on the configuration, these specimens could be further categorized into three groups: bare frame, full infilled frames, and opened infilled frame. As it shown in Fig. 3, PC frames are constructed using three-steps: 1. Constructing the beamcolumn components (refer to Fig. 3a), 2. Assembling them at the site via cast-in-place beam-to-beam and column-to-column connections (refer to Fig. 3b), and 3. Casting concrete at the cast-inplace connections (refer to Fig. 3c). Fig. 4a shows the dimensions and reinforcement details of PC specimens, in ½ scales. The reinforcement ratio and connection types examined in this study were similar to that used in Khoo et al.[43]. The height of the frame is 1515 mm, while the span of the frame is 2250 mm from the column centres. Therefore, the height/length ratio is about 1/1.5. The beam and column cross-section were 130 mm (width) 230 mm (depth) and 250 mm (width) 250 mm (depth). These frames can be categorized to frames with limited ductility in accordance with NZS 3101 [44], however, double check with ACI 318-14 [45] was also made. The transverse reinforcement at the beam and column ends was designed to be R6@70 mm while the middle zones were placed to be R6@140 mm due to limited ductile detailing required in NZS 3101 [44]. Two transverse reinforcements were installed in the beam-column joints, which is only about 30% of that required by the NSZ3101 [44] code. Similar to Khoo et al. [43], as shown in Fig. 4a, the beam-to-beam connection was composed of overlapping 180-degree hooks. To avoid the connection at the potential plastic hinge zone, the beam-to-beam connection was placed at a distance of 350 mm (1.75d) away from the column face, where d is the beam effective depth. For column-to-column
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of 240 mm 115 mm 90 mm was used for this study. The bed and head joints were 9 mm thick. Thus, the size of brick and thickness of bed and head joints actually were not scaled down. Steel fiber reinforced concrete (SFRC) was used for Specimens PCS and PCWDS at the beam-to-beam connections and column-to-column connections to investigate the enhancement efficiency of SFRC at the connections. 2.2. Material properties The selected porous brick had a mean value of compressive 0 strength (parallel to the hole), f b ; of 15.9 MPa and a void ratio of 16%. As shown in Fig. 5, compressive and shear strength of the masonry is measured to be 5.0 MPa and 0.55 MPa, respectively. The Young’s modulus of the masonry is measured to be 2.8 GPa. The compressive strength of the mortar was 6.0 MPa based on the tests on 70.3 mm cubes. As 1/2 scaled specimens were designed, the aggregate and clear cover was also scaled down. The measured properties of the concrete and reinforcements are tabulated in Tables 2 and 3, respectively. 2.3. Test setup and instrumentation Fig. 6 shows the experimental setup of a typical specimen. As shown in the figure, the lateral load was applied by a hydraulic 0 actuator. The initial axial force 0:2f c Ag was applied by a hydraulic jack through four high strength strands at the top of each column. The hydraulic jack was manually adjusted at the end of each load step to ensure the axial force in the column was almost constant during tests. As shown in Fig. 7, displacement-controlled loading procedure was applied in this study. In the initial four increments (0.1% to 0.33% drift ratio), the specimens were subjected to one fully reversed loading cycle. After that, three fully reversed loading cycles were applied at each increment. The specimen was fixed to the strong floor by foundation beams, which was fixed to the strong floor by prestressed strands with diameter of 50 mm. Brackets were specially designed to prevent out-of-plane failure of the specimens. The applied lateral load was measured by the load cell embedded in the actuator while the displacement at the loading point was measured by a linear variable displacement transducer (LVDT), which was installed at the center of the top beam. As shown in Fig. 4c, LVDTs are also utilized to measure the shear deformation of the infill panel and flexural deformation of the columns. The rigid rotation and horizontal movement of the bottom strong beam was also monitored during tests. Electric wire strain gauges (TML FLA-5-11-5LT) are installed in the beam and column longitudinal reinforcement, as shown in Fig. 4a. Fig. 3. Construction procedure of precast concrete frames: (a) precast components, (b) joint established through overlapping hooks, and (c) cast-in-place concrete at joints.
3. Test observation and results 3.1. Global response and failure modes
connections, mechanical sleeves were used for connecting the column longitudinal reinforcements, as plastic hinge can develop there. Two transverse reinforcements were also installed at the beam-to-beam connections and column-to-column connections. For RC frames, as shown in Fig. 4b, the beam and column crosssection is identical to that of PC frames. The column height and beam span of RC frames were identical to that of PC frames. Different to PC frames, longitudinal reinforcements in the beam and column were continual and the RC frames were cast-in-place. For Specimens RCW and PCW, full IMW were installed, as shown in Fig. 4c. However, as shown in Fig. 4d, for Specimens RCWD and PCWD, a door opening with size of 900 mm 500 mm was designed. A lintel with size of 750 mm 115 mm 90 mm was also placed above the door opening. Porous brick with size
3.1.1. Bare specimens Fig. 8a presents the development of crack patterns of specimen PC in critical drift ratio (DR), which is defined as the ratio of lateral displacement at the loading point to storey height. Fig. 9a shows the load-displacement curve of the specimen PC. As shown in Fig. 8a, for specimen PC, when the DR reached 0.2%, cracking first occurred at the interface between the precast and cast-in-place concrete surface of the column-to-column joints. At a DR of 0.33%, a flexural crack first formed at the beam end. When further increasing the DR to 0.5%, more flexural cracks were observed at the column-to-column joints. In the first cycle of DR of 0.67%, the hysteresis loop expressed clear signs of rebar yielding, whereby the loop flattens significantly. Increasing DR to 1.3%, horizontal
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Fig. 4. Dimensions and reinforcement details of tested specimens: (a) PC&PCS, (b) RC, (c) RCW&PCW, and (d) RCWD&PCWD&PCWDS.
cracks were observed at the beam and column ends. Moreover, diagonal cracks occurred at the column-to-column joint. However, there were fewer cracks at the beam-to-beam connection. At this stage, some pinching was observed in the hysteresis loop. At the DR of 2.0%, concrete crushing first occurred at the beam and column ends. Severe concrete spalling was observed at a DR of 2.8%. The positive peak load resisting capacity (PLRC) of 171 kN was achieved. After that large area of concrete spalled at the beam ends. The test was terminated at a DR of 5.0%, which exceeded the DR required by ASCE/SEI 7-16 [46] of 2.0%, longitudinal reinforcements buckled severely in the beam and the load resistance
dropped over 15%. In general, the hysteresis loop of specimen PC was ductile. The failure mode of PC is shown in Fig. 10a. As shown in the figure, the concrete spalling is very severe in the beam ends. Almost the whole block of concrete at the left beam end fell out and thus, the longitudinal reinforcements buckled severely. Concrete crushing was also observed at the column ends. However, compared to the beam end, it was quite mild. Little damage occurred at the beam-to-beam connections while severe flexural cracks were observed the column-to-column connections. As shown in Figs. 8b and 9b, the global behavior of PCS is very similar to that of PC in terms of crack pattern and load-displacement
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Fig. 5. Property measurements for masonry: (a) compressive strength, (b) shear strength.
Table 2 Concrete properties in different construction stages. 0
Stages
Cylindrical compressive strength f c
Tensile strength f t
Elastic modulus
Precast beam, column components Cast-in-place in connections for PC, PCW, and PCWD Cast-in-place in connections (SFRC) for PCS and PCWDS
26.8 MPa 34.5 MPa 36.8 MPa
2.4 MPa 2.9 MPa 7.8 MPa
26.7 GPa 31.1 GPa 33.0 GPa
Table 3 Properties of reinforcements. Types
Diameter
Yield Strength MPa
Ultimate Strength MPa
Elastic Modulus GPa
Elongation
R6 T12 T16
6 12 16
318 348 359
529 488 543
198 203 206
15.1% 16.3% 16.6%
Note: R and T represents plain rebar and deformed rebar, respectively.
Fig. 6. Specimen PCW ready for test.
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Fig. 7. Applied lateral displacement history.
hysteresis. Unlike that of Specimen PC, at the beginning of the test of PCS, more flexural cracks developed at the zone beyond the columnto-column connection as the SFRC was used at the cast-in-place connections. Although PCS’s load-displacement hysteresis loop was similar to that of PC before reaching the PLRC, the degradation of the load resistance beyond the PLRC in PCS was mild. As shown in Fig. 9b, the positive and negative PLRC values of PCS are 192 kN and 189 kN, respectively. Thus, the PLRC of PCS was slightly larger than that of PC. However, similar to PC, the hysteresis loop is ductile with maximum DR of 5.0%. The failure mode of PCS is illustrated in Fig. 10b. In general, the failure mode of PCS was similar to that of PC. For cast-in-place specimen RC, as shown in Figs. 8c and 9c, at the beginning stage, the first crack occurs at the bottom of the column, rather than at the interface between the precast and cast-inplace concrete surfaces. After that, the behavior of Specimen RC was similar to those of specimens PC and PCS. The failure mode of this specimen is shown in Fig. 10c. Thus, it was confirmed that the PC bare specimens with proposed construction types could achieve similar behavior of RC counterpart. 3.1.2. Infilled specimens with full IMW As shown in Fig. 8d, for specimen PCW, flexural crack occurs at the column, where is not the interface between precast and castin-place concrete surfaces. The stepped crack was also observed at IMW. At a DR of 0.5%, flexural cracks were also observed at the beam ends and more flexural cracks were formed at the column. The hysteresis loops were significantly pinched as the initial stepped crack began to spread the wall diagonally and crushing was first observed at the right upper corner of the wall. Increasing the DR to 1.3%, horizontal cracks began to develop at the beam ends. A number of diagonal cracks formed at the column top end. Slight concrete crushing occurred at the column bottom. The load resistance began to drop suddenly and thus, the hysteresis loop of this specimen was brittle. When DR reached 2.0%, concrete crushing occurred at the beam ends and severe shear deformation was observed at the column top. Crushing also occurred at the IMW. At the final portion of test, the shear failure of the column top caused the sudden drop of load resistance. The load-displacement hysteresis of this specimen is illustrated in Fig. 9d. It can be seen that the positive and negative PLRC of PCW are 424 kN and 404 kN, respectively. The ultimate deformation capacity was 40 mm in accordance with 2.8% of DR, which is larger than the required DR of 2.0% in ASCE/SEI 7-16 [46]. The failure mode of PCW is illustrated in Fig. 10d. As shown in the figure, the shear failure is concentrated on top end of the right column. Severe crushing and spalling occurred in IMW. Although crushing was also observed in the beam ends, compared to specimen PC, it was less severe than what was in PCW due to the interaction of the frame. For RC specimen RCW, as shown in Fig. 8e, flexural cracks occur at the column base with a DR of 0.2%. When the DR was increased
to 0.33%, flexural cracks occurred at the mid-height of the columns. At this DR stage, flexural cracks also formed at the beam ends and diagonal stepped cracks formed at IMW. In general, the specimen only experienced elastic response with little residual deformation after the force was released. When further increasing DR, more flexural cracks formed at the beam ends and mid-height of the columns. Two cracks were also observed at the beam-column joints. However, no new cracks noticed at the infills. When the DR reached 1.0%, the infills at the right up corner began to crush and an obvious gap was observed between the infills and surrounding frame. Obvious pinching was observed in the hysteresis loop. Diagonal cracks suddenly formed at the top of right column at a DR of 1.3%. The load resistance began to drop suddenly after this DR. Further increasing the DR, more bricks were crushed and the gap between the infills and frame widened. At a DR of 2.8%, shear failure occurred at the top of the right column. Similar to PCW, the hysteresis loop of this specimen is brittle. As illustrated in Fig. 9e, the positive and negative PLRC of RCW are 417 kN and 396 kN, respectively. The failure mode of RCW is illustrated in Fig. 10e. In general, the failure mode of RCW was very similar to that of PCW, although the wall was crushed less. 3.1.3. Infilled specimens with opened IMW As shown in Fig. 8f, for specimen PCWD, stepped cracks diagonally occurs at the left panel of wall at a DR of 0.33%. At a DR of 0.5%, a number of flexural cracks occurred at the left column. The diagonally cracks developed across the entire left panel at a DR of 0.67%. After this DR, the hysteresis loop became more flatten and pinching was observed in the hysteresis loop. Further increasing DR to 1.0% caused more crack to develop and crushing was observed at the right up corner of the panel. The hysteresis loops became more pinched and showed greater stiffness degradation. At a DR of 1.3%, horizontal cracks occurred at the beam ends and slight panel crushing was observed at there. The crushing of the panel became more severe and the brick partially collapsed. At a DR of 2.0%, more bricks collapsed and more flexural cracks developed at the columns. In general, the hysteresis loop is ductile. The concrete crushing at the beam ends was severe, which was similar to that of the corresponding bare specimen. As shown in Fig. 9f, the measured positive and negative PLRC of specimen PCWD are 246 kN and 264 kN, respectively. In addition, the failure mode of the specimen is illustrated in Fig. 10f. It can be seen that severe crushing occurred at the right column base and beam ends. However, no shear crack occurred at the beam-column joints. Wide diagonal stepped cracks occurred at the left panel and some of the bricks collapsed. For strengthened specimen PCWDS and RC specimen RCWD, a similar behavior was observed. As shown in Fig. 9g, the measured positive and negative PLRC of specimen PCWDS are 254 kN and 283 kN, respectively. Comparing to that of PCWD, the positive PLRC increased by 7.4%. For specimen RCWD, the measured positive and negative PLRC were 251 kN and 275 kN, respectively. Thus, even for infilled frame, the investigated PC construction method could achieve similar load resisting capacity as the corresponding RC specimen. The failure modes of PCWDS and RCWD are shown in Fig. 10g and h. Comparing with that of PCWD, less damage was observed in the column bottom of PCWDS. For RCWD, the crushing at the beam ends was less severe. 3.2. Stiffness degradation Fig. 11 illustrates the stiffness degradation behavior of tested specimens. It can be seen that the initial stiffness, which is defined as the ratio of load resistance to the lateral displacement at first crack stage, of PC, PCS, RC, PCW, RCW, PCWD, PCWDS, and RCWD were 28.1 kN/mm, 28.9 kN/mm, 31.4 kN/mm, 89.5 kN/mm,
L. Wang et al. / Construction and Building Materials 211 (2019) 756–770
Fig. 8. Crack pattern development of the specimens: (a) PC, (b) PCS, (c) RC, (d) PCW, (e) RCW, (f) PCWD, (g) PCWDS, and (h) RCWD.
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Fig. 9. Lateral load versus displacement hysteresis curves: (a) PC, (b) PCS, (c) RC, (d) PCW, (e) RCW, (f) PCWD, (g) PCWDS, and (h) RCWD.
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Fig. 9 (continued)
95.8 kN/mm, 63.0 kN/mm, 69.3 kN/mm, and 65.7 kN/mm, respectively. Thus, the PC specimens achieved slightly lower initial stiffness compared to the corresponding RC specimens including both bare and infilled specimen. Adding steel fiber at the cast-in-place concrete of the connection could increase the initial stiffness of the specimen by 2.8% and 10.0% for bare frames and opened infilled frames, respectively. However, the degradation of the stiffness was similar for PC and RC specimens. Moreover, the IMW could increase the initial stiffness of the PC and RC specimen by 218.5% and 205.1%, respectively. However, the door opening will decrease the initial stiffness of the PC and RC specimen by 29.6% and 27.7%, respectively; this will be further discussed in later by comparing this data with that of analytical models. 3.3. Energy dissipation capacity The energy dissipation capacity was determined by the area enclosed by the lateral load–displacement hysteresis curve. Fig. 12 illustrates the comparison of the curves of cumulative energy dissipation capacity, which is calculated by summing up the energy dissipated in consecutive loops. It was found that the energy dissipation capacity of Specimen PC, PCS, RC, PCW, RCW, PCWD, PCWDS, and RCWDS were 3.1, 3.2, 3.3, 3.1, 2.8, 3.0, 3.1, and 3.0 kNm, respectively. However, it should be noted that the lower energy dissipation capacity measured in the infilled frames was mainly due to the tests terminating when the load resisting capacity dropped over 15% from the PLRC. If the energy dissipation capacity at a DR of only 2.8% was taken into consideration, the energy dissipation capacity of infilled frames would be much larger than that of the bare frame. Similarly, the infilled frames with full walls achieved the larger value than that of the frame with opened walls. Moreover, as shown in the figure, the PC specimens achieved similar energy dissipation capacity as the corresponding RC specimens for both bare and infilled frames. Furthermore, SFRC in the cast-in-place connections has little effects on the energy dissipation capacity of the specimens. 4. Discussion of the influence of design variables 4.1. Effects of PC construction Fig. 13a and b show the comparison of the envelope of hysteresis loop and Table 4 tabulated the key results. As shown in the
figure and table, the average PLRC of specimens PC and RC are 166.5 kN and 170.5 kN, respectively. Thus, specimen PC could achieve 97.7% of the strength of the corresponding RC bare frame. Moreover, the displacement-based ductility, which is defined as the ratio of ultimate deformation capacity (Du) to the nominal yield displacement (Dy) in positive direction, of PC and RC are 5.4 and 5.8, respectively. It should be noted that the nominal yield displacement is defined as shown in Fig. 14 by assuming the area of E1 equals to E2. Thus, the ductility of specimen PC is about 93.1% of that of specimen RC. Comparing their failure modes, similar failure modes with plastic hinges at the column base and beam ends were observed. However, the interface between precast and cast-inplace concrete surfaces was prone to crack early. The average PLRC of specimens PCW, RCW, PCWD, and RCWD were 414.0 kN, 406.5 kN, 255.0 kN, and 263.0 kN, respectively. Thus, for infilled frames, PCW and PCWD could achieve 101.8% and 97.0% of the strength of RCW and RCWD, respectively. Comparing their ductility, it was found that PCW only achieve 74.4% of that of RCW. 4.2. Effects of infilled masonry walls Fig. 13c shows the comparison of the envelope of hysteretic curve of PC specimens with or without IMW and Table 4 tabulated the key results. As shown in the figure and table, the average PLRC of PC, PCW, and PCWD were 166.5 kN, 414.0 kN, and 255.0 kN, respectively. Thus, the full IMW increased its PLRC by 148.6%. The walls with door opening increased the PLRC of the bare frame by 53.2%. Moreover, the displacement-based ductility of PC, PCW, and PCWD was 5.4, 2.9, and 7.2, respectively. Thus, the full IMW only achieved 53.7% of the ductility of PC bare frame. However, the openings will increase the ductility of the infilled PC frame. Comparing their failure modes, the IMW may result in the shear failure of the column due to the interaction between the walls and frames. Moreover, the openings may affect the stability of the walls. 4.3. Effects of steel fibre in cast-in-place concrete of the connection As shown in Fig. 13d and Table 4, the average PLRC of PC, PCS, PCWD, and PCWDS are 166.5 kN, 190.5 kN, 255.0 kN, and 268.5 kN, respectively. Thus, the SFRC at cast-in-place connections could increase the average PLRC of PC bare frame by 14.4%. For infilled PC frames, the SFRC could increase the average PLRC by 3.1%. This was due to the SFRC not increasing the bending moment capacity
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Fig. 10. Failure mode of the specimens: (a) PC, (b) PCS, (c) RC, (d) PCW, (e) RCW, (f) PCWD, (g) PCWDS, and (h) RCWD.
of the column-to-column connection effectively. However, as shown in the figure, the degradation of the strength in PCS and PCWDS is much milder than that of corresponding specimens PC and PCWD. Moreover, SFRC could increase the ductility of the bare and infilled PC frame by 14.8% and 5.6%, respectively. Comparing of the failure modes, it was observed that little cracks were seen in the bare and infilled frame with SFRC in the cast-in-place column connections. However, the SFRC in the beam-to-beam connection had little effects as the damage was mainly concentrated at the beam ends, which was beyond the position of beam-tobeam connections.
5. Analytical analysis Simplified analytical models were developed to predict the PLRC of tested specimens with or without infills. 5.1. Specimens PC & PCS & RC As shown in Fig. 15, for bare frame, it is assumed plastic hinges are formed at the bottom of the column, which is actually observed in specimens PC, PCS, and RC. Thus, the PLRC of the bare frame could be determined by Eqs. (6) and (7):
L. Wang et al. / Construction and Building Materials 211 (2019) 756–770
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Fig. 11. Comparison of the stiffness degradation of tested specimens. Fig. 12. Comparison of the energy dissipated capacity of tested specimens.
Fig. 13. Comparison of the envelope of the hysteresis curves: (a) effects of PC construction for bare frame, (b) effects of the PC construction for infilled frame, (c) effects of the opening, and (d) effects of the steel fiber reinforced concrete in the joints.
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Table 4 Comparison of the critical results. Test ID
Positive Peak Load (kN)
Negative Peak Load (kN)
Energy Dissipation at test final (kNm)
Energy Dissipation at DR* of 2.8% (kNm)
Initial Stiffness (kN/mm)
Yield Displacement (mm)
Yield Strength (kN)
Ductility
PC PCS RC PCW RCW PCWD PCWDS RCWD
171 192 175 424 417 246 254 251
162 189 166 404 396 264 283 275
3.1 3.2 3.3 3.1 2.8 3.0 3.1 3.0
1.8 1.9 2.0 3.1 2.8 2.4 2.5 2.4
28.1 28.9 31.4 89.5 98.5 63.0 69.3 65.7
10.3 9.1 9.5 9.7 7.2 7.8 7.4 9.4
134 158 140 346 341 202 209 204
5.4 6.2 5.8 2.9 3.9 7.2 7.6 6.0
Note: DR represents drift ratio.
F c hc þ D Nc ¼ 2 M pc
ð6Þ
V u ¼ 2F c
ð7Þ
where Fc is the shear force in each column; Mpc is ultimate moment strength of the column considering axial force effects; Nc is the ini-
tial axial force of the column; and D is the lateral displacement in accordance with PLRC. As tabulated in Table 5, the calculated PLRC was 165 kN, which was about 99%, 87%, and 97%, of the measured average PLRC of specimens PC, PCS, and RC, respectively. 5.2. Specimens PCW & RCW As shown in Fig. 16, for infilled frame with full walls, the IMW works like a single diagonal compression strut could help to resist the lateral load, as recommended by FEMA [21]. Thus, the PLRC of PCW and RCW could be determined as below:
F c hc þ D N c ¼ 2 M pc
ð8Þ
V u ¼ 2F c þ V W
ð9Þ
0
V W ¼ at inf f m cosh
ð10Þ 0:4
where V W is the lateral resistance from IMW; a ¼ 0:175ðkhc Þ r inf is the width of the strut; k is a factor, as shown in Eq. (1); t inf is the thickness of the infill panel and equivalent strut; r inf is the diagonal length of the infill panel; h is the angle whose tangent is the infill 0 height-to-length aspect ratio; f m is the compressive strength of infill panel; Efe is modulus of elasticity of frame material; Eme is modulus of elasticity of infill material; Icol is the moment inertial of column; hinf is the height of infill panel. As tabulated in Table 5, the calculated PLRC of PCW and RCW was 344 kN. The measured average PLRC of PCW and RCW were 414.0 kN and 406.5 kN, respectively. Thus, the calculated values were 83% and 85% of the measured one for PCW and RCW, respectively.
Fig. 14. Definition of the nominal yield strength and yield displacement.
5.3. Specimens PCWD & PCWDS & RCWD For the opened infilled frame with door opening, the layout of the struts is shown in Fig. 17a and b. Thus, similar to full infills, by using superposition principle, the negative and positive PLRC could be determined by Eqs. (11)–(13), respectively:
Fig. 15. Analytical model for bare frames.
F c hc þ D N c ¼ 2 M pc
ð11Þ
V u ¼ 2F c þ V W1 þ V W2 þ V W3
ð12Þ
V u ¼ 2F c þ V W2 þ V W3
ð13Þ
Table 5 Comparison of the calculated peak load resistance with that measured from tests. Results
Measured Calculated Calculated Measured
PC (average) (kN)
166.5 165 0.99
PCS (average) (kN)
190.5 165 0.87
RC (average) (kN)
170.5 165 0.97
PCW (average) (kN)
414 344 0.83
RCW (average) (kN)
406.5 344 0.85
PCWD (kN)
PCWDS (kN)
RCWD (kN)
Posi.
Neg.
Posi.
Neg.
Posi.
Neg.
246 269 1.09
264 306 1.16
254 269 1.06
283 306 1.08
251 269 1.07
275 306 1.11
L. Wang et al. / Construction and Building Materials 211 (2019) 756–770
2.
3.
Fig. 16. Analytical model for infilled frame with full walls.
4.
5.
769
148.6% and 218.5%, respectively. Moreover, the PC frames with full infilled masonry walls could increase the energy dissipation capacity by 72.2% at drift ratio of 2.8%. Similar results were observed in counterpart RC specimens. The masonry walls with opening ratio of 27.8% could increase the lateral load resistance and initial stiffness of PC frames by 63.0% and 124.2%, respectively. At drift ratio of 4.0%, the opened walls could increase the dissipation capacity by 33.3%. For RC frames, similar extent of increasing was measured. Steel fiber reinforced concrete (SFRC) used in cast-in-place connection has little effects on behavior of PC frames in terms of initial stiffness, lateral load resistance, and energy dissipation capacity. The proposed PC construction technology, connecting the column longitudinal reinforcement from a different storey by mechanical sleeves and designing the beam-to-beam connection beyond the potential beam’s plastic hinge zone (1.5d), could achieve similar behavior to that of a cast-in-place RC bare and infilled frame. Analytical models could predict the lateral load resistance of the bare and infilled frames with reasonable accuracy.
Authors’ contributions Dr. Kai Qian gives idea for the study, Dr. Lei Wang carried out experimental tests. All authors analyzed the data and were involved in writing the manuscript. Conflicts of interest The authors declare that there is no conflict of interest regarding the publication of this paper. Acknowledgements
Fig. 17. Analytical model for infilled frame with door opened walls.
For V W1 ; V W2 ; and V W3 ; they can be determined similar to how V W is determined [21]. As shown in Fig. 17a, when positive lateral load is applied, the resistance of V W3 should consider the reduction factor 0.5 as the strut was not directly connected the corner of the frame. Similarly, as shown in Fig. 17b, when negative load was applied, the resistance of V W3 should consider the reduction factor 0.5. As the strut of V W1 was not directly connected to the corner of the frame and the strut width was relatively small, the contribution of V W1 was ignored. As it is presented in Table 5, the calculated positive and negative PLRC of PCWD, PCWDS, and RCWD were 269 kN and 306 kN, respectively; the measured positive and negative PLRC of RCWD were 251 kN and 275 kN, respectively. The calculated positive and negative PLRC were 107% and 111% of the measured ones, respectively. For PCWD, the measured positive and negative PLRC were 109% and 116% of the measured ones, respectively. Similarly, for PCWDS, the calculated positive and negative PLRC were 106% and 108% of the measured ones, respectively 6. Conclusions Based on the results of this study, the following conclusions can be drawn: 1. Introducing full infilled masonry walls to PC frames could increase the lateral load resistance and initial stiffness by
The authors gratefully acknowledge the financial support provided by the Natural Science Foundation of China (Nos. 51778153, 51568004, 51478118). The high-level innovation team in colleges and universities and excellence scholar program in Guangxi (201738). Any opinions, findings and conclusions expressed in this paper do not necessary reflect the view of Natural Science Foundation of China. References [1] L.D. Decanini, A.D. Sortis, A. Goretti, L. Liberatore, F. Mollaioli, P. Bazzurro, Performance of reinforced concrete buildings during the 2002 Molise, Italy, Earthquake, Earthquake Spectra 20 (S1) (2004) 221–255. [2] B. Zhao, F. Taucer, T. Rossetto, Field investigation on the performance of building structures during the 12 May 2008 Wenchuan earthquake in China, Eng. Struct. 31 (8) (2009) 1707–1723. [3] P.G. Asteris, D.M. Cotsovos, C.Z. Chrysostomou, A. Mohebkhah, G.K. Al-Chaar, Mathematical micromodeling of infilled frames: state of the art, Eng. Struct. 56 (2013) 1905–1921. [4] M. Dolšek, P. Fajfar, The effect of masonry infills on the seismic response of a four-storey reinforced concrete frame—a deterministic assessment, Eng. Struct. 30 (7) (2008) 1991–2001. [5] B. Pantò, L. Calio, P.B. Lourenço, Seismic safety evaluation of reinforced concrete masonry infilled frames using macro modelling approach, Bull. Earthq. Eng. 15 (9) (2017) 3871–3895. [6] F. Pires, E.C. Carvalho, The behaviour of infilled reinforced concrete frames under horizontal cyclic loading, in: Proceedings of the 10th World Conference on Earthquake Engineering, 1992, pp. 3419–3422. [7] A.B. Mehrabi, P.B. Shing, M.P. Schuller, J.L. Noland, Experimental evaluation of masonry-infilled RC frames, J. Struct. Eng. 122 (3) (1996) 228–237. [8] P. Gulan, M.A. Sozen, Procedure for determining seismic vulnerability of building structures, ACI Struct. J. 96 (3) (1999) 336–342. [9] G. Al-Chaar, M. Issa, S. Sweeney, Behavior of masonry-infilled nonductile reinforced concrete frames, J. Struct. Eng. 128 (8) (2002) 1055–1063. [10] D.J. Kakaletsis, C.G. Karayannis, Experimental investigation of infilled R/C frames with eccentric openings, Struct. Eng. Mech. 26 (3) (2007) 231–250.
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Construction and Building Materials journal homepage: www.elsevier.com/locate/conbuildmat
Seismic behavior of masonry-infilled precast concrete frames considering effects of opening Lei Wang a, Zhen-Yun Tang b, Yue Li c, Kai Qian d,⇑ a
College of Civil Engineering and Architecture, Guilin University of Technology, Guilin 541004, China The Key Laboratory of Urban Security and Disaster Engineering of Ministry of Education, Beijing University of Technology, Beijing 100124, China c Department of Civil Engineering, Case Western Reserve University, Cleveland, OH, USA d School of Civil and Environmental Engineering at Nanyang Technological University, Singapore b
h i g h l i g h t s Eight 1/2 scale, single-bay, single-storey RC and PC frames were tested. The effects of proposed PC construction technology were evaluated. The strengthening effects of SFRC at the cast-in-place connections were quantified. The reliability of analytical model for predicting the capacity of RC and PC frames was evaluated.
a r t i c l e
i n f o
Article history: Received 30 December 2018 Received in revised form 22 March 2019 Accepted 23 March 2019
Keywords: Precast concrete Reinforced concrete Infilled masonry wall Seismic Experimental study
a b s t r a c t It is commonly accepted that addition of infilled masonry walls (IMW) could enhance the in-plane strength and stiffness of RC frame buildings subjected to seismic loads. However, once the IMW walls failed, the shear force initially resisted by the IMW has to be transferred to the surrounding frames and may incur shear failure to the frames. Despite the extensive studies carried out in the past decades, studies on the structural performance of masonry infilled precast concrete (PC) frames are few. In this study, eight half-scaled single-story one bay frame were tested subjected to seismic loading, including three reinforced concrete (RC) frames and five PC frames. The investigated PC frames deviates the beam potential plastic hinge zone away from beam-to-beam connection region. The variables investigated in the tests included the types of opening in masonry infill walls and the types of cast-in-place concrete at the connections (normal concrete or steel fiber reinforced concrete). Results indicated that the investigated PC construction method has minimal effects on inelastic behavior and ductility of bare and infilled frames. However, the PC construction method may affect the crack development, especially at the interfaces in between precast and cast-in-place concrete surfaces in the region of the connection. However, the load resisting capacity and ductility was insensitive to the usage of steel fiber reinforced concrete at the connections, especially for bare frames. Ó 2019 Elsevier Ltd. All rights reserved.
1. Introduction Infilled masonry walls (IMW) are frequently utilized as cladding or partitions in reinforced concrete (RC) and steel moment resisting frames. Although it is commonly accepted that IMW may interact with the surrounding frames under earthquake load, they are often ignored in design practice. Though as secondary members, it has also been generally accepted that the IMW may affect the stiffness and strength of the infilled frames and change the seismic
⇑ Corresponding author. E-mail address: [email protected] (K. Qian). https://doi.org/10.1016/j.conbuildmat.2019.03.287 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.
response of the infilled frame, as observed in post-earthquake investigation [1,2]. However, in highly seismic regions, ignoring the IMW and frames interaction may not always be safe as the enhancement of stiffness as IMW may change the dynamic response of the frame significantly. The presence of IMW may result in unexpected force distribution and cause local damage in the frame [3]. Regarding the complexity of the interactions, extensive studies have been carried out experimentally or analytically. Since 1950s, a number of experimental tests [4–9] have been carried out to evaluate the seismic behavior of RC frames with IMW. These tests indicated that IMW could improve the
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L. Wang et al. / Construction and Building Materials 211 (2019) 756–770 Table 1 Property of test specimens. Test ID
Dimensions 2
PC PCS PCW PCWD PCWDS
Joint Trans. Rebar
Connection Type Beam-to-Beam
Column-to-Column
0.2% 0.2% 0.2% 0.2% 0.2%
Overlapping Overlapping Overlapping Overlapping Overlapping
Mechanical Mechanical Mechanical Mechanical Mechanical
2
Beam (mm )
Column (mm )
130 230 130 230 130 230 130 230 130 230
250 250 250 250 250 250 250 250 250 250
performance of RC frames significantly in accordance with the strength and energy dissipation capacity. Moreover, they were indicated that the number of bays will affect the failure mode, peak and residual capacity, and shear stress distribution of the frames significantly. Moreover, Kakaletsis and Karayannis [10,11], Mohammadi and Nikfar [12] studied the influences of openings on the seismic behavior of masonry infilled RC frames. They were found that the openings should be located as close to the edge of the panel as possible to achieve an improvement in the behavior of the infilled frames. The reduction factor of the ultimate strength of infilled frames with openings depends highly on the material of the confining frame, but the reduction factor of stiffness does not. Moretti et al. [13] and Jiang et al. [14] experimentally investigated the influence of connection between IMW and the bounding frame. The test results indicated that IMW should be connected to the frame through dowels. Including the infill walls, which were rigidly connected to the frame, could increase the strength, stiffness, and energy-dissipation capacity significantly. However, the displacement-based ductility decreased considerably. Furthermore, the possible failure modes of infilled masonry frames (corner crushing, diagonal compression failure, sliding shear failure, diagonal cracking failure, and frame failure) were summarized by tests [3,15,16]. In the past decades, researchers found simplified methods to simulate the lateral action of the IMW. The most commonly used method is diagonal strut method, which is diagonally connected the opposite compression corners of the frame [17,18]. Smith and Carter [19] indicated that the width of the strut is associated with the contact length between the frame and IMW. An analytical model is proposed to determine the equivalent width through a parameter k; which is expressed as the relative stiffness of the frame to that of infill panel and as shown in Eq. (1):
k¼
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4 ðEm t inf sin2hÞ 4Ef Icol hinf
ð1Þ
where Em ; t inf and hinf are infill’s Young’s modulus of elasticity, thickness and height, respectively; Ef Icol is the flexural rigidity of
ah ¼
Sleeve Sleeve Sleeve Sleeve Sleeve
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4Ef Icol hinf 4 2 Em tinf sin2h
p
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4Ef Ib Linf 4 aL ¼ p Em t inf sin2h
Infilled Walls
Concrete in the Connections
N/A N/A Full Wall Opened Wall Opened Wall
Normal SFRC Normal Normal SFRC
ð4Þ
ð5Þ
In the preceding two decades, researchers found that singlestrut model cannot simulate the complex behavior (bending moment or shear force) of the masonry infilled frames well, although it can fairly accurately to simulate the general behavior of infilled frames [23,24]. Therefore, multiple-strut models were proposed by researchers [16,25–32]. PC structures have advantages of high productivity, easy quality control, and cost efficiency. In the developed countries, such as U. S., Japan, Singapore, and New Zealand, PC systems with different connection types are widely implemented in the moment resisting frames. However, the PC system still has a limit in its applications in high seismic areas mainly due to several issues still unsolved. In particular, the reliable behavior of beam-column joints is a critical issue restricting the application of PC frames. The post-earthquake investigation has indicated that the collapse of PC structures is due to inadequate design of the beam-column joints [33]. So far, the design of PC system is normally required to emulate the behavior of cast-in-place RC system [34–37]. Current practice in PC frame construction indicated that the connections between precast beams are commonly located at the beam-column joint region (refer to Fig. 1). However, it should be noted that connecting the beams at the joint region is unfavorable because it disturbs the continuity of the reinforcements. Moreover, the congestion of reinforcements in the joints many bring difficulties in erection stage. Therefore, some researchers [38–40] proposed beam-to-beam connection at the midspan of the beam (refer to Fig. 2a). The connection types include 180-degree overlapping hooks, 90-degree double hooks or welding. However, as the T-shaped or cruciform precast components are very heavy and difficult to transport due to their large dimensions, applying this method in practice is
the columns; and h ¼ tan1 hLinf is the angle whose tangent is the inf
infill’s height-to-length aspect ratio. Based on experimental and analytical results, Mainstone and Weeks [20] proposed an empirical equation to determine the equivalent width of the strut as shown in Eq. (2), which was adopted by FEMA [21]. The parameter k is given in Eq. (1).
w ¼ 0:175ðkhcol Þ
0:4
ð2Þ
rinf
In CSA S340.1-04 [22], the equivalent strut width, w; could be determined using the vertical, ah ; and horizontal contact, aL ; length between the strut and the frame, h ¼ tan1 hLinf is the angle whose inf
tangent is the infill’s height-to-length aspect ratio
w¼
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi a2h þ a2L
ð3Þ
Fig. 1. Assembling method by connecting the beam reinforcements at the joint core.
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Fig. 2. Assembling configurations (a) midspan connection (Park [38]), (b) relocated the connection at a distance away from the column face, proposed by Khoo et al. [43] and used in this study.
difficult especially for the frames with long beam span [41]. To solve this problem, French et al. [42] proposed a new method, in which the connections were relocated at the beam span at a distance away from the column faces (refer to Fig. 2b). In this method, the beam-column joint with short, protruding beam stubs were cast with the column, subsequently, the beam spanning in between columns would form another precast member. Khoo et al. [43] adopted similar idea as French et al. [42] but the 180degree hooks or 90-degree double hooks were used for reinforcement lap splice. However, in these studies, the beam-column joints or beam-to-beam connections were the main concerns. The effects of column-to-column connection on seismic behavior of PC frames were ignored. In this paper, to re-exam the efficiency of proposed beam-tobeam connection proposed by Khoo et al. [43] and to evaluate the global behavior of PC frames subjected to simulated seismic load, a series of PC and RC single-storey frames were tested. Moreover, the effects of IMW with or without openings on PC frames were also investigated. 2. Experimental program 2.1. Test specimens As tabulated in Table 1, eight single-storey, single-bay, 1/2 scaled planar frame specimens were tested in this study. These geometrically identical eight specimens were categorized into two groups based on construction methods: RC-series and PC-
series. However, based on the configuration, these specimens could be further categorized into three groups: bare frame, full infilled frames, and opened infilled frame. As it shown in Fig. 3, PC frames are constructed using three-steps: 1. Constructing the beamcolumn components (refer to Fig. 3a), 2. Assembling them at the site via cast-in-place beam-to-beam and column-to-column connections (refer to Fig. 3b), and 3. Casting concrete at the cast-inplace connections (refer to Fig. 3c). Fig. 4a shows the dimensions and reinforcement details of PC specimens, in ½ scales. The reinforcement ratio and connection types examined in this study were similar to that used in Khoo et al.[43]. The height of the frame is 1515 mm, while the span of the frame is 2250 mm from the column centres. Therefore, the height/length ratio is about 1/1.5. The beam and column cross-section were 130 mm (width) 230 mm (depth) and 250 mm (width) 250 mm (depth). These frames can be categorized to frames with limited ductility in accordance with NZS 3101 [44], however, double check with ACI 318-14 [45] was also made. The transverse reinforcement at the beam and column ends was designed to be R6@70 mm while the middle zones were placed to be R6@140 mm due to limited ductile detailing required in NZS 3101 [44]. Two transverse reinforcements were installed in the beam-column joints, which is only about 30% of that required by the NSZ3101 [44] code. Similar to Khoo et al. [43], as shown in Fig. 4a, the beam-to-beam connection was composed of overlapping 180-degree hooks. To avoid the connection at the potential plastic hinge zone, the beam-to-beam connection was placed at a distance of 350 mm (1.75d) away from the column face, where d is the beam effective depth. For column-to-column
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of 240 mm 115 mm 90 mm was used for this study. The bed and head joints were 9 mm thick. Thus, the size of brick and thickness of bed and head joints actually were not scaled down. Steel fiber reinforced concrete (SFRC) was used for Specimens PCS and PCWDS at the beam-to-beam connections and column-to-column connections to investigate the enhancement efficiency of SFRC at the connections. 2.2. Material properties The selected porous brick had a mean value of compressive 0 strength (parallel to the hole), f b ; of 15.9 MPa and a void ratio of 16%. As shown in Fig. 5, compressive and shear strength of the masonry is measured to be 5.0 MPa and 0.55 MPa, respectively. The Young’s modulus of the masonry is measured to be 2.8 GPa. The compressive strength of the mortar was 6.0 MPa based on the tests on 70.3 mm cubes. As 1/2 scaled specimens were designed, the aggregate and clear cover was also scaled down. The measured properties of the concrete and reinforcements are tabulated in Tables 2 and 3, respectively. 2.3. Test setup and instrumentation Fig. 6 shows the experimental setup of a typical specimen. As shown in the figure, the lateral load was applied by a hydraulic 0 actuator. The initial axial force 0:2f c Ag was applied by a hydraulic jack through four high strength strands at the top of each column. The hydraulic jack was manually adjusted at the end of each load step to ensure the axial force in the column was almost constant during tests. As shown in Fig. 7, displacement-controlled loading procedure was applied in this study. In the initial four increments (0.1% to 0.33% drift ratio), the specimens were subjected to one fully reversed loading cycle. After that, three fully reversed loading cycles were applied at each increment. The specimen was fixed to the strong floor by foundation beams, which was fixed to the strong floor by prestressed strands with diameter of 50 mm. Brackets were specially designed to prevent out-of-plane failure of the specimens. The applied lateral load was measured by the load cell embedded in the actuator while the displacement at the loading point was measured by a linear variable displacement transducer (LVDT), which was installed at the center of the top beam. As shown in Fig. 4c, LVDTs are also utilized to measure the shear deformation of the infill panel and flexural deformation of the columns. The rigid rotation and horizontal movement of the bottom strong beam was also monitored during tests. Electric wire strain gauges (TML FLA-5-11-5LT) are installed in the beam and column longitudinal reinforcement, as shown in Fig. 4a. Fig. 3. Construction procedure of precast concrete frames: (a) precast components, (b) joint established through overlapping hooks, and (c) cast-in-place concrete at joints.
3. Test observation and results 3.1. Global response and failure modes
connections, mechanical sleeves were used for connecting the column longitudinal reinforcements, as plastic hinge can develop there. Two transverse reinforcements were also installed at the beam-to-beam connections and column-to-column connections. For RC frames, as shown in Fig. 4b, the beam and column crosssection is identical to that of PC frames. The column height and beam span of RC frames were identical to that of PC frames. Different to PC frames, longitudinal reinforcements in the beam and column were continual and the RC frames were cast-in-place. For Specimens RCW and PCW, full IMW were installed, as shown in Fig. 4c. However, as shown in Fig. 4d, for Specimens RCWD and PCWD, a door opening with size of 900 mm 500 mm was designed. A lintel with size of 750 mm 115 mm 90 mm was also placed above the door opening. Porous brick with size
3.1.1. Bare specimens Fig. 8a presents the development of crack patterns of specimen PC in critical drift ratio (DR), which is defined as the ratio of lateral displacement at the loading point to storey height. Fig. 9a shows the load-displacement curve of the specimen PC. As shown in Fig. 8a, for specimen PC, when the DR reached 0.2%, cracking first occurred at the interface between the precast and cast-in-place concrete surface of the column-to-column joints. At a DR of 0.33%, a flexural crack first formed at the beam end. When further increasing the DR to 0.5%, more flexural cracks were observed at the column-to-column joints. In the first cycle of DR of 0.67%, the hysteresis loop expressed clear signs of rebar yielding, whereby the loop flattens significantly. Increasing DR to 1.3%, horizontal
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Fig. 4. Dimensions and reinforcement details of tested specimens: (a) PC&PCS, (b) RC, (c) RCW&PCW, and (d) RCWD&PCWD&PCWDS.
cracks were observed at the beam and column ends. Moreover, diagonal cracks occurred at the column-to-column joint. However, there were fewer cracks at the beam-to-beam connection. At this stage, some pinching was observed in the hysteresis loop. At the DR of 2.0%, concrete crushing first occurred at the beam and column ends. Severe concrete spalling was observed at a DR of 2.8%. The positive peak load resisting capacity (PLRC) of 171 kN was achieved. After that large area of concrete spalled at the beam ends. The test was terminated at a DR of 5.0%, which exceeded the DR required by ASCE/SEI 7-16 [46] of 2.0%, longitudinal reinforcements buckled severely in the beam and the load resistance
dropped over 15%. In general, the hysteresis loop of specimen PC was ductile. The failure mode of PC is shown in Fig. 10a. As shown in the figure, the concrete spalling is very severe in the beam ends. Almost the whole block of concrete at the left beam end fell out and thus, the longitudinal reinforcements buckled severely. Concrete crushing was also observed at the column ends. However, compared to the beam end, it was quite mild. Little damage occurred at the beam-to-beam connections while severe flexural cracks were observed the column-to-column connections. As shown in Figs. 8b and 9b, the global behavior of PCS is very similar to that of PC in terms of crack pattern and load-displacement
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Fig. 5. Property measurements for masonry: (a) compressive strength, (b) shear strength.
Table 2 Concrete properties in different construction stages. 0
Stages
Cylindrical compressive strength f c
Tensile strength f t
Elastic modulus
Precast beam, column components Cast-in-place in connections for PC, PCW, and PCWD Cast-in-place in connections (SFRC) for PCS and PCWDS
26.8 MPa 34.5 MPa 36.8 MPa
2.4 MPa 2.9 MPa 7.8 MPa
26.7 GPa 31.1 GPa 33.0 GPa
Table 3 Properties of reinforcements. Types
Diameter
Yield Strength MPa
Ultimate Strength MPa
Elastic Modulus GPa
Elongation
R6 T12 T16
6 12 16
318 348 359
529 488 543
198 203 206
15.1% 16.3% 16.6%
Note: R and T represents plain rebar and deformed rebar, respectively.
Fig. 6. Specimen PCW ready for test.
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Fig. 7. Applied lateral displacement history.
hysteresis. Unlike that of Specimen PC, at the beginning of the test of PCS, more flexural cracks developed at the zone beyond the columnto-column connection as the SFRC was used at the cast-in-place connections. Although PCS’s load-displacement hysteresis loop was similar to that of PC before reaching the PLRC, the degradation of the load resistance beyond the PLRC in PCS was mild. As shown in Fig. 9b, the positive and negative PLRC values of PCS are 192 kN and 189 kN, respectively. Thus, the PLRC of PCS was slightly larger than that of PC. However, similar to PC, the hysteresis loop is ductile with maximum DR of 5.0%. The failure mode of PCS is illustrated in Fig. 10b. In general, the failure mode of PCS was similar to that of PC. For cast-in-place specimen RC, as shown in Figs. 8c and 9c, at the beginning stage, the first crack occurs at the bottom of the column, rather than at the interface between the precast and cast-inplace concrete surfaces. After that, the behavior of Specimen RC was similar to those of specimens PC and PCS. The failure mode of this specimen is shown in Fig. 10c. Thus, it was confirmed that the PC bare specimens with proposed construction types could achieve similar behavior of RC counterpart. 3.1.2. Infilled specimens with full IMW As shown in Fig. 8d, for specimen PCW, flexural crack occurs at the column, where is not the interface between precast and castin-place concrete surfaces. The stepped crack was also observed at IMW. At a DR of 0.5%, flexural cracks were also observed at the beam ends and more flexural cracks were formed at the column. The hysteresis loops were significantly pinched as the initial stepped crack began to spread the wall diagonally and crushing was first observed at the right upper corner of the wall. Increasing the DR to 1.3%, horizontal cracks began to develop at the beam ends. A number of diagonal cracks formed at the column top end. Slight concrete crushing occurred at the column bottom. The load resistance began to drop suddenly and thus, the hysteresis loop of this specimen was brittle. When DR reached 2.0%, concrete crushing occurred at the beam ends and severe shear deformation was observed at the column top. Crushing also occurred at the IMW. At the final portion of test, the shear failure of the column top caused the sudden drop of load resistance. The load-displacement hysteresis of this specimen is illustrated in Fig. 9d. It can be seen that the positive and negative PLRC of PCW are 424 kN and 404 kN, respectively. The ultimate deformation capacity was 40 mm in accordance with 2.8% of DR, which is larger than the required DR of 2.0% in ASCE/SEI 7-16 [46]. The failure mode of PCW is illustrated in Fig. 10d. As shown in the figure, the shear failure is concentrated on top end of the right column. Severe crushing and spalling occurred in IMW. Although crushing was also observed in the beam ends, compared to specimen PC, it was less severe than what was in PCW due to the interaction of the frame. For RC specimen RCW, as shown in Fig. 8e, flexural cracks occur at the column base with a DR of 0.2%. When the DR was increased
to 0.33%, flexural cracks occurred at the mid-height of the columns. At this DR stage, flexural cracks also formed at the beam ends and diagonal stepped cracks formed at IMW. In general, the specimen only experienced elastic response with little residual deformation after the force was released. When further increasing DR, more flexural cracks formed at the beam ends and mid-height of the columns. Two cracks were also observed at the beam-column joints. However, no new cracks noticed at the infills. When the DR reached 1.0%, the infills at the right up corner began to crush and an obvious gap was observed between the infills and surrounding frame. Obvious pinching was observed in the hysteresis loop. Diagonal cracks suddenly formed at the top of right column at a DR of 1.3%. The load resistance began to drop suddenly after this DR. Further increasing the DR, more bricks were crushed and the gap between the infills and frame widened. At a DR of 2.8%, shear failure occurred at the top of the right column. Similar to PCW, the hysteresis loop of this specimen is brittle. As illustrated in Fig. 9e, the positive and negative PLRC of RCW are 417 kN and 396 kN, respectively. The failure mode of RCW is illustrated in Fig. 10e. In general, the failure mode of RCW was very similar to that of PCW, although the wall was crushed less. 3.1.3. Infilled specimens with opened IMW As shown in Fig. 8f, for specimen PCWD, stepped cracks diagonally occurs at the left panel of wall at a DR of 0.33%. At a DR of 0.5%, a number of flexural cracks occurred at the left column. The diagonally cracks developed across the entire left panel at a DR of 0.67%. After this DR, the hysteresis loop became more flatten and pinching was observed in the hysteresis loop. Further increasing DR to 1.0% caused more crack to develop and crushing was observed at the right up corner of the panel. The hysteresis loops became more pinched and showed greater stiffness degradation. At a DR of 1.3%, horizontal cracks occurred at the beam ends and slight panel crushing was observed at there. The crushing of the panel became more severe and the brick partially collapsed. At a DR of 2.0%, more bricks collapsed and more flexural cracks developed at the columns. In general, the hysteresis loop is ductile. The concrete crushing at the beam ends was severe, which was similar to that of the corresponding bare specimen. As shown in Fig. 9f, the measured positive and negative PLRC of specimen PCWD are 246 kN and 264 kN, respectively. In addition, the failure mode of the specimen is illustrated in Fig. 10f. It can be seen that severe crushing occurred at the right column base and beam ends. However, no shear crack occurred at the beam-column joints. Wide diagonal stepped cracks occurred at the left panel and some of the bricks collapsed. For strengthened specimen PCWDS and RC specimen RCWD, a similar behavior was observed. As shown in Fig. 9g, the measured positive and negative PLRC of specimen PCWDS are 254 kN and 283 kN, respectively. Comparing to that of PCWD, the positive PLRC increased by 7.4%. For specimen RCWD, the measured positive and negative PLRC were 251 kN and 275 kN, respectively. Thus, even for infilled frame, the investigated PC construction method could achieve similar load resisting capacity as the corresponding RC specimen. The failure modes of PCWDS and RCWD are shown in Fig. 10g and h. Comparing with that of PCWD, less damage was observed in the column bottom of PCWDS. For RCWD, the crushing at the beam ends was less severe. 3.2. Stiffness degradation Fig. 11 illustrates the stiffness degradation behavior of tested specimens. It can be seen that the initial stiffness, which is defined as the ratio of load resistance to the lateral displacement at first crack stage, of PC, PCS, RC, PCW, RCW, PCWD, PCWDS, and RCWD were 28.1 kN/mm, 28.9 kN/mm, 31.4 kN/mm, 89.5 kN/mm,
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Fig. 8. Crack pattern development of the specimens: (a) PC, (b) PCS, (c) RC, (d) PCW, (e) RCW, (f) PCWD, (g) PCWDS, and (h) RCWD.
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Fig. 9. Lateral load versus displacement hysteresis curves: (a) PC, (b) PCS, (c) RC, (d) PCW, (e) RCW, (f) PCWD, (g) PCWDS, and (h) RCWD.
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Fig. 9 (continued)
95.8 kN/mm, 63.0 kN/mm, 69.3 kN/mm, and 65.7 kN/mm, respectively. Thus, the PC specimens achieved slightly lower initial stiffness compared to the corresponding RC specimens including both bare and infilled specimen. Adding steel fiber at the cast-in-place concrete of the connection could increase the initial stiffness of the specimen by 2.8% and 10.0% for bare frames and opened infilled frames, respectively. However, the degradation of the stiffness was similar for PC and RC specimens. Moreover, the IMW could increase the initial stiffness of the PC and RC specimen by 218.5% and 205.1%, respectively. However, the door opening will decrease the initial stiffness of the PC and RC specimen by 29.6% and 27.7%, respectively; this will be further discussed in later by comparing this data with that of analytical models. 3.3. Energy dissipation capacity The energy dissipation capacity was determined by the area enclosed by the lateral load–displacement hysteresis curve. Fig. 12 illustrates the comparison of the curves of cumulative energy dissipation capacity, which is calculated by summing up the energy dissipated in consecutive loops. It was found that the energy dissipation capacity of Specimen PC, PCS, RC, PCW, RCW, PCWD, PCWDS, and RCWDS were 3.1, 3.2, 3.3, 3.1, 2.8, 3.0, 3.1, and 3.0 kNm, respectively. However, it should be noted that the lower energy dissipation capacity measured in the infilled frames was mainly due to the tests terminating when the load resisting capacity dropped over 15% from the PLRC. If the energy dissipation capacity at a DR of only 2.8% was taken into consideration, the energy dissipation capacity of infilled frames would be much larger than that of the bare frame. Similarly, the infilled frames with full walls achieved the larger value than that of the frame with opened walls. Moreover, as shown in the figure, the PC specimens achieved similar energy dissipation capacity as the corresponding RC specimens for both bare and infilled frames. Furthermore, SFRC in the cast-in-place connections has little effects on the energy dissipation capacity of the specimens. 4. Discussion of the influence of design variables 4.1. Effects of PC construction Fig. 13a and b show the comparison of the envelope of hysteresis loop and Table 4 tabulated the key results. As shown in the
figure and table, the average PLRC of specimens PC and RC are 166.5 kN and 170.5 kN, respectively. Thus, specimen PC could achieve 97.7% of the strength of the corresponding RC bare frame. Moreover, the displacement-based ductility, which is defined as the ratio of ultimate deformation capacity (Du) to the nominal yield displacement (Dy) in positive direction, of PC and RC are 5.4 and 5.8, respectively. It should be noted that the nominal yield displacement is defined as shown in Fig. 14 by assuming the area of E1 equals to E2. Thus, the ductility of specimen PC is about 93.1% of that of specimen RC. Comparing their failure modes, similar failure modes with plastic hinges at the column base and beam ends were observed. However, the interface between precast and cast-inplace concrete surfaces was prone to crack early. The average PLRC of specimens PCW, RCW, PCWD, and RCWD were 414.0 kN, 406.5 kN, 255.0 kN, and 263.0 kN, respectively. Thus, for infilled frames, PCW and PCWD could achieve 101.8% and 97.0% of the strength of RCW and RCWD, respectively. Comparing their ductility, it was found that PCW only achieve 74.4% of that of RCW. 4.2. Effects of infilled masonry walls Fig. 13c shows the comparison of the envelope of hysteretic curve of PC specimens with or without IMW and Table 4 tabulated the key results. As shown in the figure and table, the average PLRC of PC, PCW, and PCWD were 166.5 kN, 414.0 kN, and 255.0 kN, respectively. Thus, the full IMW increased its PLRC by 148.6%. The walls with door opening increased the PLRC of the bare frame by 53.2%. Moreover, the displacement-based ductility of PC, PCW, and PCWD was 5.4, 2.9, and 7.2, respectively. Thus, the full IMW only achieved 53.7% of the ductility of PC bare frame. However, the openings will increase the ductility of the infilled PC frame. Comparing their failure modes, the IMW may result in the shear failure of the column due to the interaction between the walls and frames. Moreover, the openings may affect the stability of the walls. 4.3. Effects of steel fibre in cast-in-place concrete of the connection As shown in Fig. 13d and Table 4, the average PLRC of PC, PCS, PCWD, and PCWDS are 166.5 kN, 190.5 kN, 255.0 kN, and 268.5 kN, respectively. Thus, the SFRC at cast-in-place connections could increase the average PLRC of PC bare frame by 14.4%. For infilled PC frames, the SFRC could increase the average PLRC by 3.1%. This was due to the SFRC not increasing the bending moment capacity
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Fig. 10. Failure mode of the specimens: (a) PC, (b) PCS, (c) RC, (d) PCW, (e) RCW, (f) PCWD, (g) PCWDS, and (h) RCWD.
of the column-to-column connection effectively. However, as shown in the figure, the degradation of the strength in PCS and PCWDS is much milder than that of corresponding specimens PC and PCWD. Moreover, SFRC could increase the ductility of the bare and infilled PC frame by 14.8% and 5.6%, respectively. Comparing of the failure modes, it was observed that little cracks were seen in the bare and infilled frame with SFRC in the cast-in-place column connections. However, the SFRC in the beam-to-beam connection had little effects as the damage was mainly concentrated at the beam ends, which was beyond the position of beam-tobeam connections.
5. Analytical analysis Simplified analytical models were developed to predict the PLRC of tested specimens with or without infills. 5.1. Specimens PC & PCS & RC As shown in Fig. 15, for bare frame, it is assumed plastic hinges are formed at the bottom of the column, which is actually observed in specimens PC, PCS, and RC. Thus, the PLRC of the bare frame could be determined by Eqs. (6) and (7):
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Fig. 11. Comparison of the stiffness degradation of tested specimens. Fig. 12. Comparison of the energy dissipated capacity of tested specimens.
Fig. 13. Comparison of the envelope of the hysteresis curves: (a) effects of PC construction for bare frame, (b) effects of the PC construction for infilled frame, (c) effects of the opening, and (d) effects of the steel fiber reinforced concrete in the joints.
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Table 4 Comparison of the critical results. Test ID
Positive Peak Load (kN)
Negative Peak Load (kN)
Energy Dissipation at test final (kNm)
Energy Dissipation at DR* of 2.8% (kNm)
Initial Stiffness (kN/mm)
Yield Displacement (mm)
Yield Strength (kN)
Ductility
PC PCS RC PCW RCW PCWD PCWDS RCWD
171 192 175 424 417 246 254 251
162 189 166 404 396 264 283 275
3.1 3.2 3.3 3.1 2.8 3.0 3.1 3.0
1.8 1.9 2.0 3.1 2.8 2.4 2.5 2.4
28.1 28.9 31.4 89.5 98.5 63.0 69.3 65.7
10.3 9.1 9.5 9.7 7.2 7.8 7.4 9.4
134 158 140 346 341 202 209 204
5.4 6.2 5.8 2.9 3.9 7.2 7.6 6.0
Note: DR represents drift ratio.
F c hc þ D Nc ¼ 2 M pc
ð6Þ
V u ¼ 2F c
ð7Þ
where Fc is the shear force in each column; Mpc is ultimate moment strength of the column considering axial force effects; Nc is the ini-
tial axial force of the column; and D is the lateral displacement in accordance with PLRC. As tabulated in Table 5, the calculated PLRC was 165 kN, which was about 99%, 87%, and 97%, of the measured average PLRC of specimens PC, PCS, and RC, respectively. 5.2. Specimens PCW & RCW As shown in Fig. 16, for infilled frame with full walls, the IMW works like a single diagonal compression strut could help to resist the lateral load, as recommended by FEMA [21]. Thus, the PLRC of PCW and RCW could be determined as below:
F c hc þ D N c ¼ 2 M pc
ð8Þ
V u ¼ 2F c þ V W
ð9Þ
0
V W ¼ at inf f m cosh
ð10Þ 0:4
where V W is the lateral resistance from IMW; a ¼ 0:175ðkhc Þ r inf is the width of the strut; k is a factor, as shown in Eq. (1); t inf is the thickness of the infill panel and equivalent strut; r inf is the diagonal length of the infill panel; h is the angle whose tangent is the infill 0 height-to-length aspect ratio; f m is the compressive strength of infill panel; Efe is modulus of elasticity of frame material; Eme is modulus of elasticity of infill material; Icol is the moment inertial of column; hinf is the height of infill panel. As tabulated in Table 5, the calculated PLRC of PCW and RCW was 344 kN. The measured average PLRC of PCW and RCW were 414.0 kN and 406.5 kN, respectively. Thus, the calculated values were 83% and 85% of the measured one for PCW and RCW, respectively.
Fig. 14. Definition of the nominal yield strength and yield displacement.
5.3. Specimens PCWD & PCWDS & RCWD For the opened infilled frame with door opening, the layout of the struts is shown in Fig. 17a and b. Thus, similar to full infills, by using superposition principle, the negative and positive PLRC could be determined by Eqs. (11)–(13), respectively:
Fig. 15. Analytical model for bare frames.
F c hc þ D N c ¼ 2 M pc
ð11Þ
V u ¼ 2F c þ V W1 þ V W2 þ V W3
ð12Þ
V u ¼ 2F c þ V W2 þ V W3
ð13Þ
Table 5 Comparison of the calculated peak load resistance with that measured from tests. Results
Measured Calculated Calculated Measured
PC (average) (kN)
166.5 165 0.99
PCS (average) (kN)
190.5 165 0.87
RC (average) (kN)
170.5 165 0.97
PCW (average) (kN)
414 344 0.83
RCW (average) (kN)
406.5 344 0.85
PCWD (kN)
PCWDS (kN)
RCWD (kN)
Posi.
Neg.
Posi.
Neg.
Posi.
Neg.
246 269 1.09
264 306 1.16
254 269 1.06
283 306 1.08
251 269 1.07
275 306 1.11
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2.
3.
Fig. 16. Analytical model for infilled frame with full walls.
4.
5.
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148.6% and 218.5%, respectively. Moreover, the PC frames with full infilled masonry walls could increase the energy dissipation capacity by 72.2% at drift ratio of 2.8%. Similar results were observed in counterpart RC specimens. The masonry walls with opening ratio of 27.8% could increase the lateral load resistance and initial stiffness of PC frames by 63.0% and 124.2%, respectively. At drift ratio of 4.0%, the opened walls could increase the dissipation capacity by 33.3%. For RC frames, similar extent of increasing was measured. Steel fiber reinforced concrete (SFRC) used in cast-in-place connection has little effects on behavior of PC frames in terms of initial stiffness, lateral load resistance, and energy dissipation capacity. The proposed PC construction technology, connecting the column longitudinal reinforcement from a different storey by mechanical sleeves and designing the beam-to-beam connection beyond the potential beam’s plastic hinge zone (1.5d), could achieve similar behavior to that of a cast-in-place RC bare and infilled frame. Analytical models could predict the lateral load resistance of the bare and infilled frames with reasonable accuracy.
Authors’ contributions Dr. Kai Qian gives idea for the study, Dr. Lei Wang carried out experimental tests. All authors analyzed the data and were involved in writing the manuscript. Conflicts of interest The authors declare that there is no conflict of interest regarding the publication of this paper. Acknowledgements
Fig. 17. Analytical model for infilled frame with door opened walls.
For V W1 ; V W2 ; and V W3 ; they can be determined similar to how V W is determined [21]. As shown in Fig. 17a, when positive lateral load is applied, the resistance of V W3 should consider the reduction factor 0.5 as the strut was not directly connected the corner of the frame. Similarly, as shown in Fig. 17b, when negative load was applied, the resistance of V W3 should consider the reduction factor 0.5. As the strut of V W1 was not directly connected to the corner of the frame and the strut width was relatively small, the contribution of V W1 was ignored. As it is presented in Table 5, the calculated positive and negative PLRC of PCWD, PCWDS, and RCWD were 269 kN and 306 kN, respectively; the measured positive and negative PLRC of RCWD were 251 kN and 275 kN, respectively. The calculated positive and negative PLRC were 107% and 111% of the measured ones, respectively. For PCWD, the measured positive and negative PLRC were 109% and 116% of the measured ones, respectively. Similarly, for PCWDS, the calculated positive and negative PLRC were 106% and 108% of the measured ones, respectively 6. Conclusions Based on the results of this study, the following conclusions can be drawn: 1. Introducing full infilled masonry walls to PC frames could increase the lateral load resistance and initial stiffness by
The authors gratefully acknowledge the financial support provided by the Natural Science Foundation of China (Nos. 51778153, 51568004, 51478118). The high-level innovation team in colleges and universities and excellence scholar program in Guangxi (201738). Any opinions, findings and conclusions expressed in this paper do not necessary reflect the view of Natural Science Foundation of China. References [1] L.D. Decanini, A.D. Sortis, A. Goretti, L. Liberatore, F. Mollaioli, P. Bazzurro, Performance of reinforced concrete buildings during the 2002 Molise, Italy, Earthquake, Earthquake Spectra 20 (S1) (2004) 221–255. [2] B. Zhao, F. Taucer, T. Rossetto, Field investigation on the performance of building structures during the 12 May 2008 Wenchuan earthquake in China, Eng. Struct. 31 (8) (2009) 1707–1723. [3] P.G. Asteris, D.M. Cotsovos, C.Z. Chrysostomou, A. Mohebkhah, G.K. Al-Chaar, Mathematical micromodeling of infilled frames: state of the art, Eng. Struct. 56 (2013) 1905–1921. [4] M. Dolšek, P. Fajfar, The effect of masonry infills on the seismic response of a four-storey reinforced concrete frame—a deterministic assessment, Eng. Struct. 30 (7) (2008) 1991–2001. [5] B. Pantò, L. Calio, P.B. Lourenço, Seismic safety evaluation of reinforced concrete masonry infilled frames using macro modelling approach, Bull. Earthq. Eng. 15 (9) (2017) 3871–3895. [6] F. Pires, E.C. Carvalho, The behaviour of infilled reinforced concrete frames under horizontal cyclic loading, in: Proceedings of the 10th World Conference on Earthquake Engineering, 1992, pp. 3419–3422. [7] A.B. Mehrabi, P.B. Shing, M.P. Schuller, J.L. Noland, Experimental evaluation of masonry-infilled RC frames, J. Struct. Eng. 122 (3) (1996) 228–237. [8] P. Gulan, M.A. Sozen, Procedure for determining seismic vulnerability of building structures, ACI Struct. J. 96 (3) (1999) 336–342. [9] G. Al-Chaar, M. Issa, S. Sweeney, Behavior of masonry-infilled nonductile reinforced concrete frames, J. Struct. Eng. 128 (8) (2002) 1055–1063. [10] D.J. Kakaletsis, C.G. Karayannis, Experimental investigation of infilled R/C frames with eccentric openings, Struct. Eng. Mech. 26 (3) (2007) 231–250.
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