Seismic performance assessment of a precast concrete-encased CFST composite wall with twin steel tube connections

Engineering Structures 207 (2020) 110240

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Seismic performance assessment of a precast concrete-encased CFST composite wall with twin steel tube connections

T

⁎

Jing Zhoua,b, , Peng Lib, Naifu Guob a b

State Key Laboratory of Subtropical Building Science, South China University of Technology, Guangzhou 510641, China College of Civil Engineering and Mechanics, Xiangtan University, Xiangtan 411105, China

A R T I C LE I N FO

A B S T R A C T

Keywords: Precast concrete-encased high-strength concrete-filled steel tube composite shear wall Concrete-filled steel tube Vertical connection Quasi-static test Seismic performance

A new type of Precast Concrete-encased high-strength Concrete-filled Steel tube (PCCS) composite shear wall with twin steel tube connections is proposed. Six two-storey PCCS wall specimens were prepared and tested under reversed quasi-static cyclic loading. The axial load ratio, the out-of-plane load eccentricity and the connection of the outer steel tubes are considered in this investigation of the assembly integrity and seismic performance of the PCCS walls. The experimental results show that the PCCS walls exhibit good assembly integrity and satisfactory seismic performance in terms of hysteresis behaviour, ductility, stiffness degradation and energy dissipation capacity. The use of bolts to radially fasten the inner and outer steel tubes can constrain the twin steel tubes. The out-of-plane load eccentricity significantly influences the cyclic behaviour. The specimens with large load eccentricity still have good deformation and ductility capabilities, which can be attributed to the ability of the embedded high-strength concrete-filled steel tube columns to improve the out-of-plane flexural capacity of a PCCS wall. Finally, the PCCS walls are compared with other typical precast concrete walls and cast-in-place concrete-encased concrete-filled steel tube composite walls in terms of ductility capacity, and a calculation model for shear bearing capacity is proposed and evaluated.

1. Introduction A Reinforced Concrete-encased Concrete-filled Steel tube (RCCS) composite shear wall is composed of a high-strength Concrete-Filled Steel Tube (CFST) component and a Reinforced Concrete (RC) component and can have a higher bearing capacity, stiffness and ductility than a conventional RC shear wall due to the confinement provided by the embedded steel tubes. Furthermore, a RCCS wall can exhibit better durability and fire resistance than a steel-plate RC composite shear wall due to the protection provided by the outer RC encasement. Cast-inplace RCCS shear walls are known to exhibit excellent structural behaviour and are therefore increasingly used in super-high-rise buildings, especially when seismic considerations are important. With the benefits of high-strength CFST, the use of cast-in-place RCCS shear walls and RCCS columns have become more popular leading to extensive research on their mechanical properties [1–5]. Qian et al. [6] studied the seismic behaviour of RCCS walls with a circular highstrength CFST embedded in the wall boundary elements to investigate the influences of the axial load ratio and the stirrup requirements. Bai et al. [7] studied the seismic behaviour of RCCS walls with two circular high-strength CFSTs embedded in the wall boundary elements to

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replace the stirrup. Ji et al. [8] designed RCCS walls with circular or square high-strength CFST columns embedded in the boundary elements and then investigated the seismic behaviour of the composite walls. Zhao et al. [9] designed RCCS walls with circular high-strength CFSTs embedded in the wall boundary and/or centre of the composite wall to investigate the seismic behaviour. Hou et al. [10] studied the influence of the axial compression ratio on the seismic behaviour of RCCS walls. To study the mechanical properties of RCCS walls, the authors’ research team conducted a series of experimental studies [11–12], including axial compression testing, axial tensile testing, combined compression-bending testing, combined compression-shear testing, combined tension-shear testing and seismic performance testing of RCCS walls and cyclic behaviour testing of RCCS wall-RC coupling beam joints. These experimental studies [6–12] have demonstrated that RCCS shear walls have a high bearing capacity, stiffness and ductility capacity, even under high axial loads. Relevant studies based on reversed cyclic loading tests on RCCS walls with square CFSTs embedded in the boundary elements have also been conducted in other countries to alleviate the problem of reinforcement congestion in structures in those regions [13–14]. To guide the application of RCCS walls, the design requirements for RCCS walls were officially included in China’s

Corresponding author. E-mail address: [email protected] (J. Zhou).

https://doi.org/10.1016/j.engstruct.2020.110240 Received 1 September 2019; Received in revised form 8 January 2020; Accepted 13 January 2020 0141-0296/ © 2020 Elsevier Ltd. All rights reserved.

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steel tube. The outer steel tubes first embedded in the PC wall can be welded or connected by flange bolts or cannot be connected to the adjacent upper and lower storeys. The CFST core column has a continuous through-core bearing capacity and stiffness, which changes the force transfer mechanism and failure mechanism, significantly improves the shear and tensile capacity of the composite wall, and reduces the dependence on the number of vertical reinforcement connections and construction measures. Therefore, the assembly integrity and outof-plane compression stability of the PCCS wall is significantly improved, ensuring that the PCCS wall has excellent plastic deformation and ductility capacities and fully exploiting the advantage of the high load bearing capacity of the PCCS shear wall, making it possible to construct taller high-rise PC buildings. Based on current technical standards for PC buildings [22], design requirements (such as the axial load ratio and height-to-thickness ratio of the PC wall, the construction height, and the potential of building PC wall buildings in seismically active regions) can be adequately relaxed to lessen the structural weight and reduce the seismic response. The object of the current paper is, therefore, to present an experimental study of PCCS composite wall specimens under high axial force and lateral cyclic loading. The influences of the axial load ratio, out-of-plane load eccentricity, CFST crosssection ratio and the connection type of the outer steel tubes on the seismic behaviour are investigated, and the existing formulae for shear capacity calculations are evaluated.

technical specification for steel tube-reinforced concrete column structures [2]. Precast reinforced Concrete (PC) shear walls are widely used due to the construction advantages of using durable and rapidly prefabricated members to create cost-effective and high-quality structures. Currently, PC shear walls are too reliant on the jointed connection measures of vertical reinforcement to ensure construction quality. The construction of connecting elements is complex because of many issues; for example, controlling the sleeve connection of a traditional vertical reinforcement is difficult. Consequently, there are strict requirements on the use category, construction height and application regions of PC buildings. The connections between PC walls and those between the walls and the foundation require special attention to ensure good seismic performance. Therefore, a variety of PC walls with innovative jointed connections have been proposed and investigated in recent years and they have been increasingly used in seismic regions [15]. Such connections include grouted dowel connections [16], improved assembly horizontal connections [17], unidirectional Glass Fibre-Reinforced Polymer (GFRP) composite connections [18], single-row grout-filled sleeves connections [19], grouted metal duct base connections [20], bolt connections [21], etc. These innovations in precast concrete research have improved the seismic performance of PC shear walls, thereby promoting the widespread use of PC shear walls in low-rise to high-rise buildings today; however, the use of PC shear walls in super-high-rise buildings remains very limited. Because the risk of failure increases significantly due to the assembly and super-high-rise requirements, super-high-rise PC structures inevitably require a performance equivalent to that of cast-in-place structures. However, existing jointed connection technology makes it difficult for the performance of an assembled PC shear wall to match that of a cast-in-place wall, and PC shear walls have not yet met the assembly integrity and seismic ductility requirements of super-high-rise PC structures subject to strong earthquake loads or frequent wind loads. The current technical specifications for PC structures [22] are strictly limited by the seismic design of PC shear walls; for example, the axial compression ratio, construction height, and seismic available regions of the PC shear wall are more limiting than those of the cast-in-place RC shear wall. To solve these problems effectively, Chen et al. [23] proposed a type of precast concrete-encased CFST shear wall with square CFST columns embedded in the boundary element and investigated the cyclic behaviour of twostorey composite wall specimens considering variable cross-sectional shapes. In this connection system, the steel tube and wall reinforcement were cut and spliced by grout-filled coupling sleeves filled with highstrength mortar. Wu et al. [24] also studied the seismic performance of a precast concrete-encased CFST composite shear wall by entirely replacing the longitudinal reinforcement of conventional RC shear walls with a circular CFST column in the boundary element. However, the jointed connection measures of the vertical reinforcement and steel tubes are still vital to ensure construction quality in this connection system. In this paper, a new type of Precast Concrete-encased high-strength Concrete-filled Steel tube (PCCS) composite shear wall with twin steel tube connections is proposed based on existing research [23–24]. The PCCS shear wall adopts the composite mode of a precast component and a cast-in-place component and is significantly different from existing PC shear walls in terms of the wall construction, vertical connections of the main elements and bearing properties, as shown in Fig. 1. The encased high-strength CFST core column is a partially castin-place component and is used as the through-core component to connect the precast components in the adjacent upper and lower storeys. This CFST core column is not only a structural element that helps with the installation and positioning of the PC wall component but is also a part of the composite wall, providing good mechanical resistance to compression, shear and flexure. The use of bolts to radially fasten the inner and outer steel tubes avoids relative axial and circumferential movements between the outer steel tube and the inner

2. Experimental programme 2.1. Specimen design A total of six two-storey PCCS wall specimens with a scale of 1:3 to an actual structure were labelled SW1 to SW6. The six specimens had an identical overall geometry (b × h × H = 120 × 720 × 1000 mm3). The reinforcement and steel tube parameters are as follows: HRB335 reinforcing bar (HRB335 denotes the category of reinforcement, and it is marked with ; Number 335 denotes the nominal yield strength of reinforcement, namely, fy = 335 MPa), horizontal web reinforcement 8@100 (diameter = 8 mm), vertical web reinforcement 8@95, stirrups in the wall boundary element 6@100 (diameter = 6 mm), tie reinforcement 6@100 and Q345 seamless steel tube (nominal yield strength fy = 345 MPa). The concretes inside and outside the steel tubes were strength grades C70 and C50, respectively. For specimen SW5, only two circles of radial through bolts (6 6) were installed to restrain the inner and outer steel tubes. For the other specimens, the outer steel tubes were welded, and two circles of radial fastening bolts (4 6) were placed above and below the butt joint of the outer steel tubes to prevent relative axial and radial sliding between the two tubes. The vertical reinforcements in the boundary element were double fillet welded, and the vertical web reinforcements in the post-casting concrete layer were connected by strapping. The outer wall of the outer steel tube was fillet welded with 6@300 circular shear reinforcement. The material properties are shown in Tables 1 and 2. Tensile coupon tests were carried out on the steel tubes and reinforcing bars according to GB/T228 [25], and concrete tests were performed on testing standard cubes (150 × 150 × 150 mm3) after 28 days of curing according to GB 50,010 [26]. A shear span ratio of 2.8 was calculated based on the following formula.

λ=

(F2 × 2300 + F1 × 1100) = 2.8 (F2 + F1) h w

(1)

where λ denotes the shear span ratio; F1 and F2 denote the horizontal loads (F1:F2 = 1:2) applied to two loading beams in the first and second storeys, respectively; and hw denotes the effective length of the wall, hw = h − 40 mm. Details of the specimens are shown in Fig. 2a and b and Table 3. The axial load ratio, out-of-plane load eccentricity, CFST cross-section ratio and outer steel tube connection are considered. The contributions of the 2

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Fig. 1. Configuration of precast composite shear walls.

steel tubes and reinforcements are considered in determining the axial load ratios [6], which can reflect the high bearing capacity of PCCS shear walls. A large out-of-plane bending moment in the shear wall is highly likely, and the shear wall is thereby subjected to eccentric loading under the combined actions of compression, shear and bending in the in-plane and out-of-plane directions, resulting in poor PC shear wall performance. Therefore, the out-of-plane load eccentricity should be considered when investigating the role of the CFST columns in the out-of-plane stability and assembly integrity of the two-storey PCCS specimens. The CFST cross-section ratio and whether the outer steel tubes are welded should be considered mainly in terms of the economic efficiency of assembling the PCCS wall and the convenience of construction. As shown in Table 3, SW1 is the reference specimen for the axial load and the outer steel tube welding scenario, SW2 and SW1 are compared to investigate the axial load ratio, SW4 and SW3 are compared to investigate the high-strength CFST cross-section ratio (steel tube cross-section ratio), SW5 and SW3 are compared to investigate the effects of outer steel tube welding, and SW3 and SW6 are compared with SW1 to investigate the out-of-plane load eccentricity. The rest of the design parameters were identical among the different specimens. The axial load ratios were calculated based on the following formula.

nd =

Nd fc,d [A − (Acc + Ast )] + fcc,d Acc + fyt,d Ast

nt =

Nt fc [A − (Acc + Ast )] + fcc Acc + fyt Ast

(2b)

where n denotes the axial load ratio; N denotes the axial load applied to the specimen; subscripts d and t denote the design value and test load, respectively; fc and fcc denote the nominal values of the axial compressive strength of the precast concrete outer steel tube and highstrength concrete inside the steel tube, respectively; fyt denotes the yield strength of the steel tube; A denotes the gross cross-sectional area of the PCCS wall; Acc denotes the cross-sectional area of the highstrength concrete inside the steel tubes; and Ast denotes the cross-sectional area of the steel tubes. Considering both the load factor and the material strength reduction factor in the design, the ratio of the designed axial load ratio (nd) to the test axial load ratio (nt) is calculated to be 1.9 or approximately 2.0 [6]. 2.2. Experimental setup The experimental setup for the cyclic loading tests is shown in Fig. 3. In each test, the specimen was fixed to the workbench via the anchor bolts on the ground beam. Limit jacks were placed at the two ends of the ground beam to prevent sliding of the specimen in the loading direction. The out-of-plane drift of the loading beam at the top of the second storey was constrained by the sliding bar and the sliding groove in the vertical plane of the specimen. Two hydraulic servo

(2a)

Table 1 Measured mechanical properties of the concrete. Concrete

Location

Cube compressive strength (MPa)

Nominal compressive strength (MPa)

Elastic modulus (GPa)

C50 C70

Outside steel tube Precast Inside steel tubeCast-in-place

55.2 76.3

34.7 48.2

34.2 37.4

3

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Table 2 Measured mechanical properties of the steel. Steel shape

Thickness(mm)

Outer diameter (mm)

Yield strength (MPa)

Ultimate strength (MPa)

Elastic modulus (GPa)

Elongation(%)

Yield strain (%)

Steel Steel Steel Steel Steel

4.0 3.5 3.5 – –

76 63 51 8 6

355 361 366 380 371

465 470 473 492 485

1.93 1.95 1.94 1.98 2.02

22.6 22.9 22.6 22.6 18.9

0.184 0.185 0.189 0.192 0.184

tube tube tube bar bar

top of the wall and remained unchanged according to the axial load ratio of the test. The master actuator on the second storey increased the horizontal displacement at increments of 2 mm until the specimen yielded, with one cycle per displacement level. At the test site, the yielding of the steel reinforcement or steel tube in tension on the outer side of the bottom of the first-storey wall was taken as the criterion to determine the yielding of the specimen. After the specimen yielded, the displacements were increased to an increment equivalent to the yield displacement, with 3 cycles per displacement increment. Each cycle was followed by a few minutes of stoppage to record the crack development in the specimens. The loading scheme continued until the specimen could no longer sustain the load or until the load dropped to 85% of the maximum value. The loading history is shown in Fig. 4. To accommodate practical conditions, the test methodology stipulated in the standard for the testing of concrete structures [26] was followed.

actuators were used to apply horizontal cyclic loads. A hydraulic actuator applied the axial force on top of the wall through a distribution beam, and the axial load was measured by a pressure sensor. The position of the distribution beam was adjusted to achieve eccentric loading. The horizontal cyclic load was applied according to F1:F2 = 1:2 at two loading points. The electro-hydraulic servo actuator of the loading beam in the second storey was selected as the master actuator, for which the displacement control mode was used. The horizontal load of the master actuator was calculated in real time by the control system based on the displacement value; then, 50% of the horizontal load on the second storey was transferred to the actuator on the first storey to apply the first-storey horizontal load, for which the force control mode was used. The entire test process was conducted using the above loading mode to ensure that the loadings at the two points were synchronized and always maintained in proportion. Before applying horizontal cyclic loads, vertical forces were first applied at the 100120100 200 50

1200 950 200

2900

6@50

Loading beam

The second PC wall

4 14 Horizontal web bar 8@100

Post-casting layer

Vertical web bar 8@95

Welding connection

6@50

Loading beam

4 14 50 Tie reinforcement 6@100

The first PC wall Post-casting layer

300

50

12 20

500 500

4 14 8@50(4)

Ground beam

450

720 1620

Steel plate 8 mm

500

450

Shear reinforcement 6@300

400

1200 950

50

Outer steel tube D76 (63)

(a) Strapping connection

Stirrup 6@100

Tie reinforcement 6@100

Vertical web bar Horizontal web bar 8@100 8@95 D76/63

120

Welding connection

65

195

D63/51 200

195

65

720

(b)

(c)

Fig. 2. Dimensions and reinforcement details of the specimens (unit: mm): (a) vertical dimensions and reinforcements; (b) cross-section; (c) semi-finished specimen and local construction. 4

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Table 3 Design parameters of the specimens. Specimens

b×h×H

Horizontal web reinforcement

Vertical web reinforcement

Stirrup reinforcement

Tie reinforcement

Concrete types inside/outside the steel tubes

SW1

120 × 720 × 1000

8@100

8@95

6@100

6@100

C70/C50

SW2

120 × 720 × 1000

8@100

8@95

6@100

6@100

C70/C50

SW3

120 × 720 × 1000

8@100

8@95

6@100

6@100

C70/C50

SW4

120 × 720 × 1000

8@100

8@95

6@100

6@100

C70/C50

SW5

120 × 720 × 1000

8@100

8@95

6@100

6@100

C70/C50

SW6

120 × 720 × 1000

8@100

8@95

6@100

6@100

C70/C50

Specimens

ξ

nt

nd

ρ (%)

e (mm)

Whether the outer steel tubes are welded The number of radial through bolts in each circle

SW1

0.6

0.31

0.210

7.22

0

Yes, 4 6

SW2

0.8

0.41

0.210

7.22

0

Yes, 4 6

SW3

0.6

0.31

0.210

7.22

15

Yes, 4 6

SW4

0.6

0.31

0.144

5.45

15

Yes, 4 6

SW5

0.6

0.31

0.210

7.22

15

No, 6 6

SW6

0.6

0.31

0.210

7.22

40

Yes, 4 6

Note: ξ denotes the area ratio of CFST, ρ denotes the area ratio of the steel tube, and e denotes the out-of-plane load eccentricity.

V8

S1 Steel column

G2

Reaction beam

T7

T8

S2 H2

Actuator

Sliding support

S3

Steel girder

Sliding groove

V7

T6

S6

Reaction wall

Specimen

E

V3

V6 T4

Actuator

S

V4

S4 S5

W

T5

S8

T3

S9

Limit jack

H1 V2

V5 T2

S7

Ground beam

T1

G1 V1

Workbench

Fig. 5. Monitoring point layout.

Fig. 3. Test setup.

W (+) 3 Cycles

3ǻy 2ǻy

3 Cycles 3 Cycles

Top displacement

ǻy

Cycles

2×n mm Switch to cyclic loading when yielding of steel bar or steel tube

Loading scheme continued until failure occurred

E (-) Fig. 4. Loading history. 5

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Fig. 6. Comparing the cracks between specimen SW1 (Fig. 6a) and specimen SW2 (Fig. 6b), it can be seen that the cracks in specimen SW2 appeared later than those in SW1 despite SW2 exhibiting a higher axial load ratio, as shown in Table 3. Furthermore, comparison of these two specimens reveals that the concrete in specimen SW2 experienced rapid crushing, resulting in serious damage and significant buckling of the outer steel tube (Fig. 7a). Compared with specimen SW3 (Fig. 6c), specimen SW4 (Fig. 6d) had a small CFST cross-section ratio. As a result, under the same displacement loading level, numerous cracks developed in specimen SW4 due to the thick concrete outside the steel tube, and a large crushing zone formed in the late stage of the test. The diagonal cracks in specimen SW3 developed relatively slowly, as its large CFST column significantly increased the shear resistance. Unlike specimen SW3 (Fig. 6c), specimen SW5 did not have a welded outer steel tube (Fig. 6e). There were fewer horizontal and diagonal cracks in the PC of specimen SW5 than in specimen SW3, and the diagonal cracks were short; however, there was not many differences in the cracking and failure modes between specimen SW3 and specimen SW5, indicating that the radial bolts played a major role in combining the inner and outer steel tubes into a single overall component capable of resisting internal forces. For specimens SW1 (Fig. 6a), SW3 (Fig. 6c) and SW6 (Fig. 6f), as the out-of-plane eccentricity increased, the horizontal cracking increased, while the diagonal shear cracking slowed on the south and north surfaces. In specimen SW3, there were many long horizontal cracks, and the initiation points of the diagonal cracks were closer to the central axis of the wall. In specimen SW6, horizontal cracks appeared earlier on the south surface, developed more extensively and were long, and the crushing of the concrete was concentrated at the bottom of the north surface of the wall under eccentric loading. Specimen SW6 was essentially subjected to a complex stress state under compression, shear and out-of-plane bending, correspondingly decreasing the energy dissipation and deformation of this specimen. However, no out-of-plane instability failure occurred, and the assembly integrity was still good. In specimens SW3 (Fig. 6c) and SW6 (Fig. 6f), which were subjected to out-of-plane eccentric loading, horizontal cracks appeared on the south surface in the second storey (Fig. 7b). There was no crushing damage to the concrete on the north surface, and the strain of the steel tube and reinforcement was still in the elastic range. No significant cracks or damage were observed in the second-storey walls of the other specimens.

The loads, displacements and strains at key locations were measured. The locations of the measuring points were identical in the six specimens. There were a total of 9 displacement gauges (Si), among which 6 were placed in the plane of the wall (2 at the end of the loading beam and 4 at one-third and two-thirds of the wall height) and 3 were placed on the ground beam (1 horizontal and 2 vertical) to monitor the torsion of the ground beam. A total of 20 strain gauges were placed 100 mm above the post-casting concrete layer in each storey of the specimen to measure the longitudinal strain and circumferential strain (Ti) of the outer steel tube, the strain (Vi) of the vertical web reinforcement, the strain (Hi) of the horizontal web reinforcement and the strain of the stirrups in the boundary elements (Gi). The layout of the strain gauges on the reinforcement and steel tubes in the first storey of the specimen is shown in Fig. 5. 3. Test results and analysis 3.1. Experimental observations For clarity, the front surface and back surface of the PCCS wall under observation are implied to as the south surface and north surface, respectively, whereas the left side and the right side as observed from the south are denoted as west side and east side, respectively. The damage in each specimen mainly occurred in the first storey. The first storey successively experienced initial cracking on the east and west sides, the development of horizontal cracks and diagonal shear cracks on the south surface and north surface, the yielding of the reinforcement and steel tubes under tension and compression and the crushing of the concrete in the boundary elements. These processes caused the bearing capacity of the specimen to gradually decline. The failure mode of the overall specimen was combined flexural-shear failure. Diagonal shear cracks occurred after the development of horizontal cracks, and certain cracks in the precast wall were well developed and widely distributed. There was no in-plane or out-of-plane relative movement between the PC wall and the post-casting concrete layer of the specimen, and the CFSTs had significant connecting and restraining effects. Three types of cracks developed in the first storey of the wall, namely, horizontal flexural cracks, diagonal shear cracks and vertical splitting cracks of the concrete cover in the boundary element. Generally, at displacement levels 2 or 3 (4–6 mm) of the loading of the master actuator, horizontal flexural cracks first appeared on the east side and west side of the post-casting concrete layer at the bottom of specimens. As the displacement increased, the cracks gradually developed upward and became exposed in the PC wall. The flexural cracks developed approximately horizontally towards the central axis of the wall on the south surface and north surface, and there were significant differences in the lengths of the horizontal cracks among the different specimens. Subsequently, diagonal shear cracks formed, which developed from horizontal cracks and extended at an angle of approximately 45° towards the central axis of the wall. As the displacement increased, certain diagonal cracks intersected close to the central axis of the wall. However, the concrete-encased CFST inhibited the development of the principal diagonal cracks on the wall surface and thus avoided failing via the compression-shear or tension-shear failure modes. After the reinforcing bars and outer steel tubes in the boundary elements yielded under tension or compression, vertical splitting cracks appeared in the concrete at the toe of the wall. Before the crushing of the concrete, the cracks in the concrete cover of the post-casting concrete layer were sparsely distributed and short. When the specimen reached the peak load, its bearing capacity began to gradually decrease, and the concrete in the wall toe was crushed and collapsed. Both the concrete in the postcasting concrete layer and the concrete in the PC wall were damaged, but there were differences among the specimens. After the tests, the crushed concrete was cleared, and the outer steel tube and the reinforcing bars in the boundary elements were observed to be significantly buckled. The first-storey failure of the specimens is shown in

3.2. Load–displacement curves The hysteresis curve and skeleton curve of the first storey are shown in Figs. 8 and 9, respectively, where F is the sum of the horizontal loads of the two storeys (F = F1 + F2). Before the cracking of the specimens, the load–displacement relationship was basically linear. When the average drift ratio was approximately 1/300, the stiffness decreased, the bearing capacity increased, and there was residual deformation after unloading. When the average drift ratio was approximately 1/140, the horizontal load reached its peak, and the residual deformation was significant. The six specimens all exhibited different degrees of pinching, mainly due to bond slip after the formation of diagonal shear cracks and to the large axial load ratio adopted in the test. Specimens SW1 and SW2 had overall arch-shaped hysteresis curves with relatively full loops and experienced little bond slip. Specimens SW3 to SW6 had overall inverted S-shaped hysteresis curves and significant bond slip. Except for the hysteresis curve of specimen SW2, which had the highest axial load ratio, the hysteresis curves of the other five specimens had a long segment involving a gradual decrease in the load bearing capacity after the peak load, indicating that the CFST effectively withstood the horizontal force and that the PCCS composite walls had a stable mechanical performance and large deformation capacity in the later 6

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Fig. 6. Failure modes of the specimens: (a) SW1; (b) SW2; (c) SW3; (d) SW4; (e) SW5; (f) SW6.

the hysteresis curve. The main difference was manifested in the magnitude and attenuation rate of the bearing capacity. Compared with specimen SW4, specimen SW3 had a large peak load due to its large CFST column, which significantly increased the horizontal load bearing capacity.

stages. The axial load ratio significantly influenced the cyclic behaviour. Compared with specimen SW1, specimen SW2 had a higher axial load ratio and a higher peak load. However, after the peak load, the bearing capacity of specimen SW2 rapidly declined, the concrete in the compression zone was rapidly crushed, and the number of loading cycles decreased. Due to the effect of the embedded CFST, the deformation of specimen SW2 in the late stage of the test did not significantly decrease, indicating that even at a very high axial load ratio, the specimen still maintained excellent deformation capability. Other types of PC shear walls had little advantage in terms of deformation capability. The outof-plane load eccentricity impacted the hysteretic behaviour. For specimens SW1, SW3 and SW6, as the eccentricity gradually increased, the area of the hysteresis loop gradually decreased, and the pinching effect significantly increased, but the ultimate deformation did not significantly decrease. Therefore, the force resistance characteristics of the embedded CFST subjected to out-of-plane compression-flexure were fully exploited, and the out-of-plane stability and assembly integrity of the PCCS wall were good. The CFST cross-section ratio and whether the outer steel tube was welded did not significantly influence the shape of

3.3. Strain of the outer steel tubes The representative strain hysteresis curves of six specimens were obtained from the outer steel tube of the boundary element at the bottom of the wall on the first storey (100 mm above the interface of the post-casting concrete layer), as shown in Fig. 10. The vertical strain in each specimen reached or exceeded the yield strain of steel (0.184–0.192%), indicating that the outer steel tube had good strain development under cyclic loading and that the vertical connections of the twin steel tubes and the PC wall were reliable. In specimen SW5, for which only radial bolts were used to fasten the inner and outer steel tubes, the maximum tensile strain reached 0.263%, and the strain hysteresis curve was as full as that of SW3, as shown in Fig. 10g, indicating that the inner and outer steel tubes constrained and connected 7

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Fig. 7. Local damage of the specimens: (a) buckling of the reinforcing bar and steel tube of SW2; (b) cracking of the second-storey wall of SW6.

the twin steel tubes and worked jointly to resist the vertical and horizontal internal forces. The weld-free method may be considered for the outer steel tube to simplify the assembly process of the PCCS wall and reduce the workload of field welding. The specimens entered the plastic deformation stage after yielding; the vertical compressive strain (negative value in Fig. 10) of the outer steel tube of the specimens was generally larger than the vertical tensile strain, and the hysteresis loop of the tensile strain was not as large as that of the compressive strain. Thus, vertical shortening of the PCCS walls occurred due to the buckling of the steel tube and compression of the concrete under an eccentric load or a high axial load. Specimen SW1 had a relatively complete strain hysteresis curve, and the tensile strain was slightly smaller than the compressive strain, indicating that during the plastic deformation stage in slightly unfavourable working conditions, the steel tube could withstand internal in-plane bending and tensile forces; additionally, the PCCS wall had good resistance to bending. Specimen SW2 had a large axial load ratio, and the circumferential tensile strain of the outer steel tube was the largest due to the radial push from the inner steel tube subjected to compression, leading to significant buckling of the outer steel tube. Specimen SW6 had a large out-of-plane load eccentricity, and thus, the circumferential tensile strain of the outer steel tube was relatively large due to the inner steel tube pushing the outer steel tube in the radial direction.

axial compression ratio, had a large characteristic point load. Specimen SW4, with the smallest CFST cross-section ratio, had a small characteristic point load. The mean θ1y, θ1p, θ1u and μ of all the specimens were approximately 1/300, 1/140, 1/70 and 4.5, respectively. The seismic design code for buildings in China [27] requires the plastic drift ratio of castin-place RC shear walls in areas with rare earthquakes (corresponding to an earthquake exceedance probability of 2–3% in a 50-year period) to be limited to 1/120. The PCCS composite walls in the present study exceed the code requirements. The specimens achieved a performance equivalent to that of a cast-in-place wall and exhibited good deformation capability. The encased cast-in-place CFST core column played a significant role in improving the integrity of the composite wall. Specimen SW6 had the smallest ductility factor of 3.75, and the out-of-plane load eccentricity influenced its displacement ductility. Specimen SW2 had a high ductility factor of 4.95, indicating that the CFST significantly improved the compression bearing capacity and seismic integrity of the PCCS wall. Although specimen SW4, with a small CFST cross-section ratio, had a low bearing capacity, its ductility factor was 5.15, which was the highest of the specimens, indicating that the CFST cross-section ratio of the PCCS shear wall needs to be optimized. The axial and lateral bearing capacities, deformation, ductility, and vertical connection integrity can be comprehensively considered in future in-depth studies.

3.4. Bearing capacity and deformation

3.5. Stiffness degradation

The values of the horizontal loads and displacements at the characteristic points are presented in Table 4, including the yield load (Fy), yield displacement (Δ1y), peak load (Fp), peak displacement (Δ1p), ultimate load (Fu), ultimate displacement (Δ1u) and ductility factor (μ) for each specimen in the first storey. The yield point was calculated using the method shown in Fig. 11. The ultimate point was determined when the horizontal load decreased to 85% of the peak load. The drift ratio was θ1 = Δ1/H1 (H1 is the calculated height of the first storey of the specimen and was taken as 1100 mm). The displacement ductility factor was μ = Δ1u/Δ1y. The yield load of each specimen was approximately 0.78 times the peak load. Specimen SW2, with the highest

The stiffness degradation curves of the specimens by relative rigidity (Kj/Kin) as a function of relative displacement (Δ/Δ1y). Kj is expressed as n

Kj =

n

∑ P ij ∑ uij i=1

i=1

(3)

where Kin is the initial rigidity corresponding to the first loading cycle; P ij and uij are the maximum load and horizontal displacement, respectively, under the ith loading cycle when relative displacement (Δ/Δy) equals j; and n is the number of loading cycles at each force or displacement level. 8

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Fig. 8. Hysteretic loops of lateral force versus displacement in the first storey of the specimens: (a) SW1; (b) SW2; (c) SW3; (d) SW4; (e) SW5; (f) SW6.

stiffness. For a PC wall, this value indicates a relatively high residual stiffness, which enhances the collapse resistance of the overall structure. Specimen SW2 had the largest axial compressive ratio and highest initial stiffness and exhibited fast stiffness degradation before yielding. Specimen SW4 had a small CFST cross-section ratio and exhibited relatively fast stiffness degradation before yielding.

As shown in Fig. 12, the specimens had different initial stiffnesses, and their rates of stiffness degradation also varied. In the later stages, the ultimate stiffness of the six specimens converged. This result mainly occurred because the lateral stiffness was predominantly provided by the reinforcement bars and CFSTs after the concrete outside the steel tube and the concrete in the post-casting concrete layer cracked, failed and largely ceased to function. Taking 1/70 as the average ultimate drift ratio of the specimens, the corresponding average ultimate stiffness of the six specimens was approximately 14.6% of the initial 9

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the cracking and spalling of concrete outside the steel tube, the encased high-strength CFST component can still bear the horizontal shear force in the joint connection, which ensures that the PCCS composite wall exhibits excellent assembly integrity and good energy dissipation capacity and has a performance that is equivalent to that of cast-in-place walls. These combination effects are significantly affected by factors including the steel ratio of the CFST component, the area ratio of the CFST component and the out-of-plane load eccentricity. The advantages of the PCCS composite wall over existing PC shear walls are clearly highlighted. The innovation of the PCCS composite wall is the transition from cast-in-place construction to partially preassembled production and partially cast-in-place construction in the creation of an RCCS composite shear wall. The PCCS composite shear wall has a significant advantage in terms of wall construction, vertical connections, bearing properties and seismic performance. The encased high-strength CFST column is not only a structural component that helps with the installation and positioning of the PC wall component but is also a part of the composite wall, providing good mechanical resistance to compression, shear and flexure. The encased CFST column changes the force transfer mechanism and failure mechanism, significantly improves the shear and tensile capacity of the PC wall, and reduces the dependence on the number of vertical reinforcement connections. Therefore, the assembly integrity and out-of-plane compression stability of the PCCS wall is significantly improved, ensuring that the PCCS wall has excellent ductility capacities, making it possible to construct high-rise PC buildings.

Fig. 9. Envelope curves of lateral force versus displacement in the first storey of the specimens.

3.6. Energy dissipation The dissipated hysteretic energy of the specimen was obtained by calculating the area enclosed by the hysteresis curve [6,28]. The cumulative dissipated hysteretic energy Ed corresponding to the ultimate displacement Δ1u was calculated, as shown in Fig. 13. Specimen SW1, without out-of-plane load eccentricity, exhibited the highest cumulative dissipated energy. Specimen SW2 had a larger axial compression ratio, which decreased the number of loading cycles and reduced the energy consumption. Specimen SW3, with relatively small out-of-plane load eccentricity, also exhibited a large cumulative dissipated energy. The deformation of specimen SW4, with a small CFST cross-section ratio, was relatively large, but the area of the hysteresis loop was reduced, and the dissipated energy was less than that of SW3. Specimen SW5 exhibited significant pinching and dissipated less energy than SW3. Specimen SW6, with a large out-of-plane load eccentricity, had the lowest cumulative dissipated energy because the cracking damage occurred earlier.

5. Calculation of shear strength As previously mentioned, the studied PCCS composite walls have a performance equivalent to cast-in-place performance; therefore, the calculation of the bearing capacity of elements subjected to combined axial forces, bending and shear is identical to that for RCCS composite walls. However, particular attention should be paid to shear resistance because the major concern when designing such a PC composite member is the prevention of shear failure. The shear resistance of a composite section built from three or more encased CFSTs is not yet covered by design codes. The present study aims to develop a design method to calculate the shear strength. The encased CFSTs are considered reinforcing bars, and the shear strengths of the PCCS composite walls are calculated based on the superposition method proposed in GB 50010 [26]. For simplicity reasons, only four encased CFSTs with the same cross-sectional size are considered here. The shear strength of the PCCS composite walls can be calculated by adding the shear strength of the PC component and that of the CFST component (Fig. 15), as follows:

4. Comparison with previous research Here, the ductility factors of the PCCS composite wall in the present study are compared with the ductility factors from previous studies. Chen et al. [23] investigated the cyclic behaviour of a two-storey PCCS composite wall with CFSTs embedded in the boundary element and presented the ductility factors of two composite wall specimens. Typical cast-in-place RCCS composite walls, based on ductility factor test data presented by Qian et al. [6], Bai et al. [7], Ji et al. [8], Zhao et al. [9] and Dan et al. [13], are also selected for comparison. The detailed data are not illustrated in the current paper due to space limitations. The basic design parameters (e.g., number, shape and location of CFSTs; the testing axial load ratio; and the shear span ratio) and the average ductility factor are illustrated in Table 5. Flexural-dominated failure was observed in the previous and current research. For different types of walls, a general conclusion can be drawn. The average ductility factors in the present study are much higher than those in the study of Chen et al. [23]. The average ductility factors of the PCCS specimens in the present study and of the cast-inplace RCCS specimens in the previous studies [6–9,13] are 3.75–5.15 and 3.3–6.9, respectively. The overall average ductility factors of the PCCS specimens in the present study and of the cast-in-place RCCS specimens in the previous studies are 4.54 and 4.65, respectively, as shown in Fig. 14. The PCCS composite wall has a ductility factor similar to that of the RCCS composite wall due to the combination of the encased high-strength CFST core component and the PC component. After

V = VPC + VCFST = Vc + Vsb + Vcc + Vst

(4)

where V is the shear strength of the composite section, VPC is the shear strength of the PC component, VCFST is the shear strength of the CFST component, Vc is the shear strength provided by the precast concrete in the shear-compression zone, Vsb is the shear strength provided by the horizontal reinforcement bars, and Vst is the shear strength provided by the steel tubes. 5.1. Shear strength of the PC component Many factors are involved in the shear failure of a PC shear wall: the damage pattern is complex, and the understanding of the shear mechanism is insufficient. There is no perfect theoretical system for determining the shear strength of a PC wall; therefore, the calculation methods for the bearing capacity of a PC wall differ among different national codes. In the present approach, based on the experimental results and Chinese code, there are four calculation assumptions: (i) The condition for forming a flexural-shear crack is that the tensile stress of the tensile side at the height of hw/2 above the monitoring cross-section 10

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Fig. 10. Hysteretic loops of the lateral force versus outer steel tube strain in the specimens: (a) SW1; (b) SW2; (c) SW3; (d) SW4; (e) SW5; (f) SW6; (g) comparisons between SW3 and SW5.

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Table 4 Test results of specimens at characteristic points. Specimen

SW1 SW2 SW3 SW4 SW5 SW6

Loading direction

+ − + − + − + − + − + −

Yield point

Peak point

Ultimate point

Ductility factor

Fy (kN)

Δ1y (mm)

θ1y

Fp (kN)

Δ1p (mm)

θ1p

Fu (kN)

Δ1u (mm)

θ1u

μ

mean

338.5 320.3 388.8 329.1 313.9 333.5 286.4 294.3 289.6 337.6 331.1 343.4

4.50 4.28 3.21 3.16 3.27 3.76 3.32 3.67 3.21 4.04 3.39 3.67

1/244 1/257 1/343 1/348 1/336 1/293 1/331 1/300 1/343 1/272 1/324 1/300

417.3 403.3 457.6 403.2 406.9 426.7 380.7 390.1 389.3 430.7 414.4 424.2

9.50 7.50 6.80 6.40 9.23 8.77 7.51 11.18 7.87 6.84 7.47 9.29

1/116 1/147 1/162 1/172 1/119 1/125 1/146 1/98 1/140 1/161 1/147 1/118

354.6 342.3 388.5 342.7 345.8 362.6 323.5 332.0 330.9 365.7 352.2 360.4

16.73 17.99 15.02 16.50 16.35 15.97 15.86 20.16 17.54 17.16 14.38 12.23

1/66 1/61 1/73 1/67 1/67 1/69 1/69 1/55 1/63 1/64 1/76 1/90

3.7 4.2 4.7 5.2 5.0 4.2 4.8 5.5 5.5 4.2 4.2 3.3

3.95 4.95 4.60 5.15 4.85 3.75

shear crack forms due to the resulting shear force. According to the principle of material mechanics and assumption (i), the bending stress at point B (as shown in Fig. 16) can be expressed as

MB N M − Vc (h w 2) N − t = A − t Wc Ac Wc Ac

σB =

(5)

The shear span ratio λ, section resistance moment Wc of the PC wall and calculation assumption (iii) are respectively given by

λ=

MA Vc h w

(6)

be h w2 6

(7)

σB = 0.3fc = 3f tc

(8)

Wc =

Combining Eqs. (5), (6), (7) and (8), one obtains

Fig. 11. Diagram of the yielding point and ultimate point.

Vc =

1 N 1 N ⎛0.05fc be h w + t ⎞ = ⎛0.5f tc be h w + t ⎞ λ − 0.5 ⎝ 6⎠ λ − 0.5 ⎝ 6⎠

be = b −

mπd 2 4h

(9) (10)

where σB and MB are the bending stress and bending moment at point B, respectively, and be is the equivalent width of the PC wall (as shown in Fig. 15). λ is the shear span ratio. For cantilever wall specimens, the shear span ratio is equal to the wall aspect ratio. The lower and upper bounds of λ for walls that undergo flexural-dominated failure are limited to 1.5 and 2.2 (i.e., values greater than 2.2 are assumed to be equal to 2.2) in the Chinese code [29]. Additionally, m is the number of CFSTs, and d is the diameter of the outer steel tube. The experimental results in a previous study [11] demonstrated that the factors affecting the shear strength of an RCCS wall include the strength of the concrete, the shear span ratio, the ratio of the horizontal reinforcement and the axial tension force. Eq. (9) fully reflects the influence of the above factors and disregards the favourable effect of stirrups on the shear strength of concrete due to the beneficial effect of CFSTs. Out-of-plane load eccentricity has an effect on the hysteretic behaviour of the PCCS shear wall, but the effect on the shear capacity is not large, as shown in Fig. 9. Out-of-plane load eccentricity is not taken into account in the calculation model of shear strength. The axial compression force is beneficial for maintaining the shear strength of eccentrically compressed shear walls. In the present analysis, a 1/6 axial force is considered to increase the shear strength of the PC component. However, after the compression is increased beyond a certain extent, the beneficial effect with respect to shear decreases; thus, the axial force needs to be limited in Chinese code [29] (i.e., values greater than 0.2fcbh are assumed to be equal to 0.2fcbh). The actual compression section is also different from the assumed one; thus, Eq. (9) is adjusted and rewritten as

Fig. 12. Stiffness degradation curves of the specimens.

is 0.3fc (fc is the compressive prism strength of the outer steel tube). (ii) The condition for forming a web shear crack is that the main tensile stress in the web reaches the concrete tensile strength ftc (ftc = 0.1fc). (iii) The shear strength of the PC shear wall is the sum of the shear strength of the concrete and the shear strength of the horizontally distributed steel bars calculated according to the 45° strut-and-tie model. (iv) Under the repeated loading of tensile and compressive forces, the shear strength of a PCCS composite wall decreases by 20%. According to the strut-and-tie model, the shear force transfer in the composite wall is shown in Fig. 16. Under combined compressive and shear loading, the wall specimen at the tensile side mainly withstands the bending moment when the shear span ratio λ > 1.5, but a flexural12

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8

Ref.[6-9, 13]

7

6.9

Dutility factor

6

Current study 5.15

5

4.65

4.54 4 3.75

Ref. [23] 3.2

3 Fig. 13. Energy dissipation curves of the specimens.

2 Vc =

1 ⎛0.5f c be h w + Nt be ⎞ t λ − 0.5 ⎝ 6b ⎠

(11)

Ash A Z cot θ = fy sh h w s s

2.3

Precast

Precast

Cast-in-place

Fig. 14. Comparisons of ductility factors with previous studies.

The contribution of horizontal steel bars in a shear wall to the shear strength has been well studied. Referring to the calculation method of an ordinary RC shear wall, the shear strength associated with the horizontal web steel bars calculated according to the 45° strut-and-tie model is as follows:

Vsb = fy

3.3

5.2. Shear strength of the CFST component If the encased CFSTs are considered steel bars, the shear strength can be evaluated using the strut-and-tie model. However, it is obvious that the shear strength provided by CFSTs cannot be neglected in such composite sections. The test results in the present study demonstrate that the CFST cross-sectional area affects the shear strength of the PCCS wall specimen. To conduct a statistical analysis, the PCCS wall is treated as an integrated part of the entire shear wall because there was no shear failure at the interface between the CFST and the concrete. In this case, full interaction between the concrete and encased outer steel tube is ensured by shear bar connectors, and full interaction between the inner and outer steel tubes is ensured by radial through bolts. The contribution to the shear strength is the sum of the embedded CFST shear

(12)

where Ash is the cross-sectional area of horizontal reinforcements in the same cross-section, fy is the yield strength of the reinforcement bars, s is the spacing between the horizontal reinforcements, Z is the calculated section height, and θ is the 45° strut-and-tie model.

Table 5 Comparisons with previous studies. Type of wall

Specimen

Number, shape and location of CFSTs

Testing axial load ratio

Shear span ratio

Ductility factor

Reference

Precast

SW1 SW2 SW3 SW4 SW5 SW6

4○, 4○, 4○, 4○, 4○, 4○,

0.31 0.41 0.31 0.31 0.31 0.31

2.8 2.8 2.8 2.8 2.8 2.8

3.95 4.95 4.60 5.15 4.85 3.75

Current study

Precast

SW1 SW2

2×□, boundary 2×□, boundary

0 0

1.6 1.6

2.3 3.2

Chen et al. [23]

Cast-in-place

SW2 SW3 SW4 SW5 SW6

2×○, boundary 2×○, boundary 2×○, boundary 2×○, boundary 2 × 2○, boundary

0.26 0.31 0.38 0.32 0.31

2.0 2.0 2.0 2.0 2.0

4.2 4.3 3.3 3.4 4.4

Qian et al. [6]

Cast-in-place

STHW2 STHW3

2 × 2○, boundary 2 × 2○, boundary

0.16 0.19

2.1 2.1

4.7 4.0

Bai et al. [7]

Cast-in-place

SRCW4 SRCW5

2×□, boundary 2×○, boundary

0.34 0.32

2.3 2.3

4.6 6.1

Ji et al. [8]

Cast-in-place

SW2 SW3 SW4 SW5 SW6 SW7 SW8

2 × 2○, boundary 2 × 2○, boundary 2 × 3○, boundary 2 × 3○, boundary 2○+2○+2○, boundary and middle 2○+2○+2○, boundary and middle 6○, uniform distribution

0.18 0.26 0.22 0.28 0.20 0.25 0.21

2.1 2.1 2.1 2.1 2.1 2.1 2.1

6.9 3.5 5.4 4.2 6.2 4.0 4.7

Zhao et al. [9]

Cast-in-place

CSRCW1

2×○, boundary

0.02

2.9

5.1

Dan et al. [13]

uniform uniform uniform uniform uniform uniform

distribution distribution distribution distribution distribution distribution

Note: □ indicates a rectangular steel tube, and ○ indicates a circular steel tube. 13

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be

600

300

200

100

V CFST=95.3

V CFST=127.1

400

Estimating values Testing values SW2 + SW3 SW5 SW6 SW1 + + - SW4 - + + + -

V PC=305.1

h

Shear strength (kN)

500

V PC=297.6

b

0

Specimen V

V

PC

Fig. 17. Comparison between the theoretical and measured results of the shear strength.

V

CFST

Fig. 15. Decomposition of a composite section into a PC section and CFST section.

The shear strength of the encased CFSTs can be established according to experimental data, as shown in Eq. (15):

VCFST = Vtp − Vc − Vsb

strengths. The pure shear bearing capacity of the CFST with twin steel tubes when the shear span ratio equals zero is calculated as follows:

VCFST = f tcc Acc + (1 − α ) fyt Ast

where χ is the coefficient to be determined and V is the peak load from the test. Considering the shear strength of a specimen as the dependent variable, the coefficient is obtained from Eqs. (14) and (15). Along with calculation assumption (iv), the shear strength of the PCCS composite wall is incorporated into Eqs. (11), (12) and (14), which are simplified as follows:

(13)

where α is the strength reduction factor considering the influence of the transverse shear in combined bending and compression (generally α = 0.4 in the Chinese code) and ftcc is the tensile strength of the concrete in the steel tube. Referring to the provision for steel RC shear walls in the code for the design of composite structures [29], to consider the effect of shear span ratio, the contribution of the encased CFSTs to the shear strength is calculated as follows:

VCFST = Vcc + Vst = χ

Vt =

f tcc Acc + 0.6fyt Ast λ

(15)

(14)

1 ⎛0.4f be h w + 2Nt be ⎞ + 0.8fy 15b ⎠ λ − 0.5 ⎝ t 0.2(f tcc Acc + 0.6fyt Ast ) Ash hw + λ s

Fig. 16. Strut-and-tie method for shear force transfer in the composite wall. 14

(16)

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The materials strengths measured in the test (i.e., the nominal compressive strength of the concrete and the yield strengths of the reinforcement bars and the steel tube) are used as the representative material strengths. Introducing the strength parameters into Eq. (16), the predicted values are compared with the experimental values, as shown in Fig. 17. For specimens SW1 ~ SW3, SW5 and SW6 in Table 3, Vpc = Vsb + Vc = 89.8 + 207.8 = 297.7 kN, VCFST = 127.1 kN and Ves = 424.7 kN. For specimen SW4 in Table 3, Vpc = Vsb + Vc = 97.3 + 207.8 = 305.1 kN, VCFST = 95.3 kN and Ves = 400.4 kN. Ves is the theoretical bearing capacity calculated using Eq. (16). The Ves/Vtp ratio varies between 0.93 and 1.09, with a mean value of 1.02 and a standard deviation of 0.042. Thus, the proposed calculation model provides fairly consistent shear strengths of the PCCS composite wall. The contribution of the shear strength of the small CFST area ratio (SW4, 23.8%) is lower than that of the large CFST area ratio (SW1 ~ SW3, SW5 and SW6, 29.9%).

(7) The calculation model provides a fair estimation for the shear strength of a PCCS wall; this model takes into account the influence of transverse shear on the combined axial force and bending moment.

6. Conclusions

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.

In the present study, steel tubes and vertical reinforcement bars were welded in PCCS composite wall specimens, requiring an enormous amount of welding. Future research will focus on weld-free or flange connections for the outer steel tubes, pressed sleeve connections for the vertical reinforcing bars, optimization of the CFST cross-section ratio and steel tube wall thickness and overall stability performance of the PCCS composite wall. Further experimental and analytical studies need to extend the parametric range in the assessment of the proposed calculation model. Declaration of Competing Interest

This paper presents an innovative composite wall: a PCCS composite shear wall with twin steel tube connections. Six PCCS composite wall specimens were tested under combined vertical loads and horizontal cyclic loads. The major findings and conclusions obtained from this study are summarized as follows:

Acknowledgements The authors gratefully acknowledge the financial support provided by the Natural Science Foundation of Guangdong Province, China (grant no. 2018A030313688), the Natural Science Foundation of Hunan Province, China (grant no. 2018JJ2401) and the Foundation of the State Key Laboratory of Subtropical Building Science, South China University of Technology (grant no. 2018ZB28).

(1) The PCCS composite shear wall has advantages in terms of wall section configuration, vertical connection capability, vertical and lateral load bearing capacity, seismic hysteretic behaviour and assembly integrity, and it is suitable for high-rise PC buildings. (2) The overall failure mode of the tested specimens was combined flexural-shear failure, as manifested by the tensile yielding of the reinforcing bars and outer steel tube, the crushing of the concrete in the boundary elements, the buckling of the outer steel tube and the significant diagonal shear cracking. Cracks developed and were distributed in both the post-casting concrete layer and the PC wall. The CFSTs played significant connecting and constraining roles. (3) The hysteresis curves of the specimens exhibited obvious pinching. After the peak load, there was a long segment with a gradual decline in the bearing capacity. The specimens showed good plastic deformation and ductility capabilities. The average ultimate drift ratio was approximately 1/70, and the average ductility factor was 4.5, indicating that the PCCS walls had good assembly integrity. Even the specimen with a high axial load ratio of 0.41 exhibited good deformation and ductility capabilities during the later stages of testing, although its bearing capacity declined and its stiffness rapidly decreased. (4) There was no significant difference in the performances of the specimens for which only bolts were used to radially fasten the inner and outer steel tubes and the specimens for which the outer steel tubes were welded and bolts were used to radially fasten the inner and outer steel tubes. The use of only bolts to radially fasten the inner and outer steel tubes can effectively constrain and connect the twin steel tubes. (5) The out-of-plane load eccentricity had an important influence on the hysteretic performance of the specimens. The specimens with relatively large load eccentricity still had good deformation and ductility capacities. The CFST core column reliably played constraining and connecting roles and improved the out-of-plane compression-flexural resistance and assembly integrity of the precast wall. (6) The CFST cross-section ratio influenced the performance of the specimens, and the specimen with the smallest cross-section ratio had the highest displacement ductility. The parameters of the PC composite wall should be optimized by comprehensively considering the compression and lateral force bearing capacities, deformation, ductility and assembly integrity of the specimens.

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