Cyclic testing of steel beam blind bolted to CFST column composite frames with SBTD concrete slabs

Engineering Structures 148 (2017) 293–311

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Cyclic testing of steel beam blind bolted to CFST column composite frames with SBTD concrete slabs Jingfeng Wang a,b,⇑, Beibei Li a, Donghua Wang a, Chunfeng Zhao a a b

School of Civil Engineering, Hefei University of Technology, Anhui Province 230009, China Anhui Civil Engineering Structures and Materials Laboratory, Anhui Province 230009, China

a r t i c l e

i n f o

Article history: Received 15 March 2017 Revised 23 May 2017 Accepted 25 June 2017

Keywords: Concrete-filled steel tubular (CFST) composite frame Blind bolt End plate connections Steel-bars truss deck (SBTD) Seismic behavior

a b s t r a c t This paper aims to investigate the seismic behavior and failure mechanism of novel type of blind bolted concrete-filled steel tubular (CFST) composite frames under seismic loading. The composite frames are composed of square or circular CFST columns and steel-concrete composite beams with steel-bars truss deck (SBTD); the steel beams and the CFST columns are assembled by blind bolts and end plates. The failure modes, hysteretic behavior, strength and rigidity degradation, ductility and energy-dissipating performance of all specimens were analyzed and discussed. The influence of composite action of the floor slab on the performance of blind bolted CFST composite frames were also explored under seismic loading. The experimental results showed that the blind bolted CFST composite frames with the SBTD concrete slabs exhibited good hysteretic behavior, large ductility and high energy dissipation. The stiffness and strength of the frame increased greatly with the presence of the SBTD concrete slab. These test studies could promote the engineering design and application of the blind bolted CFST composite frames. Ó 2017 Elsevier Ltd. All rights reserved.

1. Introduction The concrete filled steel tubular (CFST) composite structures have been extensively used in multi- and high-rise buildings in China owing to their obvious advantages such as high load bearing capacity, easy construction, overall cost-effectiveness and excellent earthquake resistance. The CFST columns are usually assembled with the steel beams in practical frame structures, and the concrete slabs are often connected to the steel beam flanges by shear connectors. Traditionally, composite beam-to-column joints are designed either as notionally pinned or rigid connections. However, the actual performance of the composite joints lies between these two idealized conditions, namely the semi-rigid connections [1]. Previous experimental studies [2–4] on CFST composite joints with blind fasteners and end plates indicated that the semi-rigid characteristic of beam-to-column joints and composite action of slabs should be properly considered during the design process of CFST composite frames. The static and cyclic behaviors of CFST columns were investigated by many scholars [5–11]. Furthermore, the existing steel beam to CFST column connections are mostly welding, such as the internal or external diaphragm plates, and additional fittings. ⇑ Corresponding author at: School of Civil Engineering, Hefei University of Technology, Anhui Province 230009, China. E-mail address: [email protected] (J. Wang). http://dx.doi.org/10.1016/j.engstruct.2017.06.065 0141-0296/Ó 2017 Elsevier Ltd. All rights reserved.

However, the Northridge and the Kobe earthquakes confirmed that steel beam-to-column joints welding suffered from cracks and brittle fractures. In order to avoid extensive welding and improve the ductility and energy-dissipating capacity of joints, a growing number of researchers focus on the blind bolted connections with the development and promotion of construction industrialization. The static behavior of blind bolted HSS or CFST column joints has been studied by some scholars, such as Korol et al. [12], France et al. [13], Lee et al. [14], Wang et al. [15–17], Yao et al. [18], Wang et al. [19] and Wang et al. [20]. Other scholars explored the hysteretic behavior of blind bolted joint to CFST columns, such as Goldsworthy and Gardner [21], Elghazouli et al. [22] and Wang et al. [23,24]. Owing to the existence of the concrete slabs and the composite action between the slabs and the steel beams by shear connectors, the lateral torsion buckling was prevented and the moment carrying capacity of steel beams was improved in the CFST column to composite beam rigid connections [25–27]. However, few scholars paid attention to the blind bolted moment CFST joints with concrete slabs. Moreover, previous studies [2–4,28–31] mainly focused on the blind bolted flush end plate composite joints, which concrete slab with or without profiled steel sheets (PSSs). Considering different levels of shear connection and reinforcement ratios, Loh et al. [2] completed experiments on five flush end plate composite joints which the concrete slabs were supported by PSSs.

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2000

A

SBTD concrete slab

400

A

M20 Grade10.9 Blind bolts

Secondary beam HN 250×125×6×9 Primary beam HN 300×150×6.5×9

M20 Grade10.9 Blind bolts

SBTD concrete slab

1550

Flush end plate

Secondary beam HN 250×125×6×9

1475

Extended end plate

Primary beam HN 300×150×6.5×9

Square CFST column 200×8

(a) Specimen SCF1 2000

A

SBTD concrete slab

400

A

M20 Grade10.9 Blind bolts

Secondary beam HN 250×125×6×9 Primary beam HN 300×150×6.5×9

1550

Flush end plate

M20 Grade10.9 Blind bolts

SBTD concrete slab

Secondary beam HN 250×125×6×9

1475

Extended end plate

Primary beam HN 300×150×6.5×9

Circular CFST column 200×8

(b) Specimen CCF1 Fig. 1. Dimension of specimens (unit: mm).

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J. Wang et al. / Engineering Structures 148 (2017) 293–311 Table 1 Information of the test specimens. Specimen

Story

Height (mm)

Colum section B
t (mm)

Beam section hb
bf
tw
tf (mm)

End plate type

SCF1

1 2

1475 1550

h 200
8 h 200
8

H300
150
6.5
9 H300
150
6.5
9

Extended end plate Flush end plate

CCF1

1 2

1475 1550

s 200
8 s 200
8

H300
150
6.5
9 H300
150
6.5
9

Extended end plate Flush end plate

600 100

200

200

70

Top chord rebar (TC rebar)

100 Web rebar

15

Bottom chord rebar (BC rebar) Thin-walled steel sheeting

Fig. 2. Steel-bars truss deck.

Ataei et al. developed moment-rotation models of flush end plate composite joints [3] and conducted four blind bolted flush end plate composite connections with precast concrete slabs [28]. Mirza and Uy [29] and Thai and Uy [30] reported experimental and numerical analysis results of flush end plate composite connections. Wang et al. investigated the cyclic performance of blind bolted flush or extended end plate composite joints between CFST columns and composite beams by quasi-static tests [4] and pseudo-dynamic tests [31]. The blind bolted moment composite frames consist of CFST columns and steel-concrete composite beams with steel-bars truss deck (SBTD). The composite beams are formed with the concrete slab casting on SBTD and connected to the steel beam top flange by the shear studs. The steel beams were assembled to the square or circular CFST columns by the blind bolts and the end plates. The blind bolted end plate joints is a good way to overcome extensive welding work and improve the ductility and energy dissipation of joints. The type of beam-to-column connections provide a reasonable degree of continuity and optimization of the moment distribution in composite frames. Meanwhile, the use of SBTD could leave out the bottom templates of slabs and significantly reduce the amount of work site reinforcement colligation. Nevertheless, few literatures reported the structural performance of the type of blind bolted moment CFST composite frames under seismic loading. Presently, some scholars studied the seismic performance of CFST frames with rigid beam-to-column connections, such as Han et al. [32,33], Wang et al. [34], Herrera et al. [35], Nie et al. [36], and He et al. [37]. Currently, only Wang et al. [38] studied the cyclic performance of the blind bolted moment square CFST frames without concrete slabs; the experiment results made known that the type of the blind bolted CFST frame exhibited excellent earthquake resistance performance. However, Ref. [38] don’t consider the composite action of the floor slabs and the circular CFST column composite joints. Thus, those highlight the need for an extensive research on the seismic behavior of blind bolted CFST composite frames with SBTD concrete slabs. Against above background, a series of experimental and analysis studies were carried out to explore the cyclic behavior and failure

mechanism of blind bolted end plate composite frames between square or circular CFST columns and SBTD concrete composite beams. Two single-bay, two-story specimens of the type of blind bolted CFST composite frames were conducted under low cyclic loading. The failure modes and slab crack pattern were recorded and analyzed. The hysteretic behavior, strength and stiffness degradation, ductility and energy-dissipating performance of all specimens were analyzed from test data. The influence of composite action of the floor slabs on the performance of blind bolted CFST composite frames were also explored under seismic loading. The experimental analysis results would be useful in the design of blind bolted CFST composite frames with SBTD concrete slabs. 2. Experimental program 2.1. Test specimens The objective of this paper aims to insight into the influence of the column section type and the end plate type on the anti-seismic behavior and failure modes of the blind bolted CFST composite structures. Two 1/3 scale models of single-bay, two-story blind bolted CFST columns to steel beams composite frames were designed and manufactured. The configuration and dimension of both specimens were shown in Fig. 1 and Table 1. The first and second story height of each frame are respectively 1475 mm and 1550 mm, and the beam span is 2000 mm. The columns of specimen SCF1 are square CFST with a cross-section of 200
200
10 mm, while the columns of specimen CCF1 are circular CFST with a cross-section of 200
10 mm. The primary beams are commercial H-shape steel beam with a cross-section of HN 300
150
6.5
9 mm and the secondary beams with a cross-section of HN 250
125
6
9 mm. The width and thickness of cast-in-situ composite slabs are respectively 1200 mm and 100 mm in each story. The steel-concrete composite beams are formed with the concrete slab casting on SBTD and connected to the steel beam flange with the shear studs. The composite beams were designed to form a full shear connection between the steel beam and the SBTD con-

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(a) A-A (specimen SCF1)

(b) A-A (specimen CCF1) Fig. 3. Details of concrete slab (unit: mm). Note: The symbol ‘‘ ” represents the yield strength of the reinforcing bar is of grade 335 N/mm2.

crete slab. The SBTD consists of steel-bars truss and thin-walled steel sheeting. The diameter of top chord rebar (TC rebar) and bottom chord rebar (BC rebar) is 8 mm and the diameter of web bar is 4.5 mm. The spacing and height of the steel-bars truss are respectively 200 mm and 70 mm (seen in Fig. 2). The arrangement of transverse distributed rebar (TD rebar) and the end constructional rebar (EC rebar) are illustrated in Fig. 3. Two rows of shear studs which with the diameter of 16 mm and height of 90 mm were welded through the thin-walled steel sheeting to the top flanges of the primary beams, while single row of shear studs were welded to the top flanges of the secondary beams. It meets the full shear connection criteria in accordance with specification GB50017 [39]. The layout of shear studs is shown in Fig. 4. The flat end plates were applied to the square CFST composite frame (specimen SCF1), while the curved end plates were used to the circular CFST composite frame (specimen CCF1). The design of the type of connections presented in this paper was referred to EC4 [40] and other scholar’s research findings [2–4,28–31]. For

each specimen, two pairs of 12 mm thick extended and flush end plates were respectively welded to the first and second story primary beam end by fillet welds. Details of the end plates are shown in Figs. 5and 6. The primary steel beams with end plates were fastened to steel tubes by the high strength blind bolts with hooked extensions into the concrete core, as shown in Fig. 7. Previous test studies [2–4,28–31] showed that the typed blind bolted composite joints exhibited semi-rigid characteristics. The hook-typed extension to the bolt is high strength reinforcing rebar with 20 mm in diameter, 70 mm in horizontal length and 35 mm in hooked length and the yield strength of the extensions is of grade 335 N/mm2. The test results [41] verified that the blind bolts with the hook-typed extensions enhanced the strength and initial rigidity of the type of joints. The blind bolt applied in both specimens is Grade 10.9 M20. All bolts for the beam-to-column joints were finally tightened to a torque of 442 Nm recommended in code GB50017 [39]. The blind bolted CFST composite frames were installed by laboratory personnel. Firstly, the end plates were welded to the pri-

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Fig. 4. Details of composite joints.

45 110 45

D

12

D

340

65 70 70 70 65

D

60 129

D

Diameter φ 22

540

162 129 60

Diameter φ 22

45 110 45

45 110 45 200

200

(a) Extended end plate

(b) Flush end plate Fig. 5. Flat end plates.

mary steel beam end and then assembled to CFST columns by means of high strength blind bolts. While the flange-welded web-bolted connections were applied on the secondary beam-tocolumn connections (illustrated in Fig. 4). Then the SBTD was installed on the surface of beam top flanges by shear studs which

were welded through the thin-walled steel sheeting. Thereafter, the transverse and end construction rebars were orderly colligated manually in the laboratory. Meanwhile, the edge banding plates were fixed at two sides of CFST columns to avoid concrete flow everywhere. Finally, ordinary concrete was poured on SBTD in each

J. Wang et al. / Engineering Structures 148 (2017) 293–311

φ 22

E

110 °

540

E

0 12

65 70 70 70 65

E

60 129

E

Diameter φ 22

10

162 129 60

Diameter φ 22

340

298

45 110 45 200

45 110 45 200

(a) Extended end plate

(b) Flush end plate Fig. 6. Curved end plates.

Steel tube wall M20 Grade 10.9 Blind bolt

35

Hooked rebar extension

70

End plate

(a) Diagram of the blind bolt

(b) Photo of the blind bolt Fig. 7. Blind bolts with extension.

Table 2 Material properties of steel. Specimen

Steel wall thickness (mm)

Yield stress (N/mm2)

Ultimate stress (N/mm2)

Young’s modulus (N/mm2)

Elongation at fracture (%)

Circular steel tube Square steel tube Steel beam flange Steel beam web End plate Web rebar TC/BC rebar TD/EC rebar

8 8 9 6.5 12 4.5 8 12

335.7 338.3 381.2 358.1 363.8 392.3 386.5 364.3

478.2 485.7 498.5 485.2 473.9 472.6 438.2 417.4

2.03
105 1.97
105 2.01
105 2.14
105 2.08
105 2.02
105 2.06
105 2.06
105

21.0 20.1 20.3 21.5 20.8 19.3 22.5 20.6

Table 3 Material properties of core concrete. Specimen

Specimen dimension (mm)

Age (day)

fcu (N/mm2)

1-1 2-1 3-1 Average 1-2 2-2 3-2 Average

150
150
150 150
150
150 150
150
150

28 28 28

50.12 52.74 57.13 53.62

100
100
300 100
100
300 100
100
300

28 28 28

Ec (N/mm2)

32456.5 30754.7 32648.2 34657.0

299

J. Wang et al. / Engineering Structures 148 (2017) 293–311

story after the steel tubular columns filled with self-consolidating concrete (SCC). 2.2. Material properties For the steel properties, steel coupons were cut from steel tubes, sheets and rebars and conducted to determine the tensile strength, modulus of elasticity as well as breaking elongation. Table 2 showed the results of steel material tests. Besides, the yield stress and ultimate stress of M20 Grade 10.9 blind bolts are respectively 932 N/mm2 and 1023 N/mm2. The steel tubes of both specimens were casted with commercial SCC mix. Four group tests were undertaken to get the material properties of SCC and slab concrete. Each group had three concrete cube with the size of 150
150
150 mm for the cube compressive strength and the size of 100
100
300 mm for the elastic modulus. The cube compression strength of the SCC and slab concrete was determined by the standard cubic compression test. The average cube compressive strength of core concrete in the steel

tubes is 53.62 N/mm2 at 28 days and the modulus of elasticity is 34,657 N/mm2, as shown in Table. 3. The average cube compressive strength of ordinary concrete used in the concrete slabs is 29.43 N/mm2 at 28 days, and the modulus of elasticity is 28,702 N/mm2, as listed in Table 4. 2.3. Test setup The test setup is illustrated in Fig. 8 and photograph of test site is shown in Fig. 9. Each specimen was tested under constant vertical load and low-cyclic lateral load. A rigid steel reaction frame was constructed to apply constant vertical load on the top of column ends and accommodate large lateral displacements of the actuator to the specimens. The constant vertical load executed on the top of columns by two hydraulic jacks, and the gliding devices were installed on jacks to reduce friction between jacks and the rigid reaction frame. The ratio of axial compression of CFST column is 0.3 during test process. The low-cyclic lateral load applied to the CFST column wall by two servo controlled hydraulic actuators.

Table 4 Material properties of slab concrete. Specimen

Specimen dimension (mm)

Age (day)

fcu (N/mm2)

C31 C32 C33 Average C41 C42 C43 Average

150
150
150 150
150
150 150
150
150

28 28 28

30.04 28.79 29.46 29.43

100
100
300 100
100
300 100
100
300

28 28 28

27674 29445 28801 28702

Steel reaction frame Hydraulic jack SBTD concrete slab Secondary beam

MTS actuator B

Steel rod

Primary beam

CFST column SBTD concrete slab Secondary beam

MTS actuator A

Steel rod

Primary beam

RC reaction wall Rigid foundation Ground beam

Concrete beam

Ec (N/mm2)

Steel rod

Ground anchor

Fig. 8. Experimental diagram.

Hydraulic jack

300

J. Wang et al. / Engineering Structures 148 (2017) 293–311

The column end horizontal load (P) and displacement (D) were automatically captured by the hydraulic servo system. Furthermore, there were eleven linear variable displacement transducers (LVDTs) to obtain the connection rotation, deflection of primary beams, sliding displacement at the column base and lateral interstory displacement of specimens. The arrangement of the LVDTs is depicted in Fig. 11. Strain gauges were placed on critical points of steel beam, steel tube, end plate, reinforcement rebar and shear stud to determine the stress distribution. A total of seventy-two strain gauges were respectively mounted on specimen SCF1 and CCF1 (seen in Fig. 12).

3. Experimental results 3.1. Failure modes

Fig. 9. Experimental setup photograph.

8 6 4 2 0 -2 -4 -6 -8

0

4

8

12

16

20

24

28

32

Number of cycles (n) Fig. 10. Loading history.

The column bottom was constrained by steel rigid foundations to form fixed condition through ground anchors and long steel rods. Preloading procedure was conducted to ensure the equipments ran normally. The displacement loading protocol was applied during formal loading process in accordance with the ATC-24 [42] guidelines, as illustrated in Fig. 10. The displacement of specimen corresponding to the yield load (0.7Pmax, and Pmax is assumed as theoretical value of maximum load of specimens) is defined as the theoretical yield displacement (Dy). Two cycles were employed to specimen at each displacement level of 0.125Dy, 0.25Dy, 0.5Dy and 0.7Dy; three cycles were employed to specimen at each displacement level of 1Dy, 1.5Dy and 2Dy; two cycles were exerted at each displacement level of 2.5Dy, 3Dy, 4Dy, 5Dy, 6Dy. According to the inverted triangular load distribution, the loading displacement ratio of the first and second stories of the tested specimens were calculated as 0.63. The theoretical yield displacement (Dy) of the first and second stories are respectively 12.6 mm and 20 mm.

During the loading cycles of 0.7Dy and 1.0Dy (namely that the maximum lateral displacement of the first inter-story was respectively 8.82 mm and 12.6 mm), the first crack which located in the first story slabs was observed around the left side CFST column and subsequently appeared near the opposite side CFST column. The cracks were visible around perimeter of the column region up to the loading cycles at 2.0Dy (namely that the maximum lateral displacement of the first inter-story was 25.2 mm). New diagonal cracks appeared and original diagonal cracks continuously propagated in longth and width with increasing of inter-story dispalcement. Some transverse cracks were developed and formed finally. Meanwhile, the concrete around CFST column in the first and second story slabs began to cursh and peel off. Seperation between the CFST columns and the slabs was also apparent at latter loading cycles. The sizzling sound of the concrete compression and frizzle sound of the composite frame deforamtion could be heard during loading process. The failure modes of cracks and concrete spalling of slabs were shown in Figs. 13(a), (c) and 14(b). There was bending deformation on the extended or flush end plates and the deformation of extended end plates was larger than that of flush end plates. The maximum deformation of extended end plate of specimen SCF1 and CCF1 was respectively about 15 mm and 14 mm, whilst the maximum deformation of flush end plate of specimen SCF1 and CCF1 was respectively about 5 mm and 3 mm, as seen in Figs. 13(a) and 14(a). The slight local buckling on the left-hand beam flange was also observed. The end plates and steel tubular walls of the CFST column were in touch with each other around the blind bolts during the loding process, although the end plates buckled between the gap of two rows of blind bolts (Figs. 13(a) and 14(a)). The inner-side stiffener was not welded to each CFST column in the manufacturing factory owing to negligence of the workers. The test was terminated as the lateral displacement increased to 6Dy (namely that the maximum lateral displacement of the first and second inter-stories were respectively 75.6 mm and 44.4 mm), weld fracture occurred at the bottom end of the CFST columns (Figs. 13(d) and 14(d)) of specimen SCF1 and CCF1, and weld fracture subsequently also appeared at the primary beam-to-extended end plate connection of specimen CCF1 (Fig. 14 (c)). This experimental phenomenon demonstrated that the setting of stiffeners or other construction methods for CFST columns was very necessary during the practical design of CFST columns.

3.2. Crack pattern of slab The failure patterns of slab cracks in each specimen were shown in Figs. 15 and 16. On the basis of careful observation and recording, it was found that:

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J. Wang et al. / Engineering Structures 148 (2017) 293–311

100 40

40 100

400

400

1720

300

D3

D8

D9

D7

1250

1550

D11

300

D2

D5

D6

D4

1175

1275

D10

200

D1

Fig. 11. Layout of LVDT’s.

Specimen SCF1: (1) The cracks in the first story slab expanded along the inner side of the square CFST column to both side edges of the slab, and some diagonal cracks gradually formed transverse cracks across the slab width; (2) the cracks in the first story slab spread from two side edges of square CFST column to mid-span of the slab and the extended range of cracks lies between 1/3 and 4/15 of beam span away from the central axis of the square CFST columns; (3) the crack distribution of the second story slab were similar to that of the first story slab, while the quantity and distribution of cracks in the second story slab were obviously lessen than those in the first story slab. Specimen CCF1: (1) The cracks was in radial pattern to expand from the circular CFST column to the surrounding slab and some full-length cracks gradually formed in a certain extent time; (2) the cracks spread from side edges of circular CFST column to mid-span of the slab and the extended range of cracks lies between 10/27 and 10/23 of beam span away from the central axis of circular CFST columns; (3) the quantity and distribution of cracks in the second story slab were obviously lessen than those in the first story slab. In general, the SBTD concrete composite slab exhibited good structural performance during the seismic loading process. Only a few number of cracks occurred in the side slab; the thin-walled steel sheeting located at the bottom of slab did not buckle and

no overall crushing of the floor slab was observed. The maximal crack of specimen SCF1 and CCF1 were respectively 2 mm and 3 mm. The slab and steel beam remained intact in the process of testing. The test results indicated that the steel-concrete composite beams worked in full shear interaction. 3.3. Load-displacement hysteretic behavior The hysteretic curves of inter-story shear (Ps)-drift (Ds) relation in each story were illustrated in Figs. 17 and 18. The test results demonstrated that the blind bolted CFST composite frames with SBTD concrete slabs exhibited good hysteresis performance and ductility. Some conclusions could be drawn from Figs. 17 and 18. The inter-story shear of both specimens increased with increasing of lateral displacement and the slope of curves began to decrease as the specimen was in inelastic stage. The strength and stiffness degradation of specimen SCF1 were more obvious than those of specimen CCF1 at each same level loading, especially as the specimens stepped into elastic-plastic and plastic stages. On the one hand, there was almost no pinching effect on hysteretic curves of specimen CCF1 during the whole loading, while little pinching effect appeared on hysteretic curves of specimen SCF1 at the inelastic loading stage. It indicated that the anti-seismic perfor-

J. Wang et al. / Engineering Structures 148 (2017) 293–311

100 100

33

34

100 200 200 250

60 59 58 57

56 55 54 53

43

44

41

42

(b) Strain gauges of square column section

32

29

30

350

40

39

100

925

31

(c) Strain gauges of circular column section

52 51 50 49

100

48 47 46 45

200

100

14

300

13

100

1300

200

100

36

400

28

27

35

100 100

1600

100

302

120 30 150

200

38

37

(d) Strain gauges of beam section

190 200

700

400

9 10 5 6 1 2

190 200

10

(e) Strain gauges of extended end plate in 1st story

7 8 3 4 Steel-bars truss top chord Steel-bars truss bottom chord

(g) Strain gauges of SBTD in 1st story

10

700

23

11 12

65 340

105

72 190 200

65

105

71

10

(f) Strain gauges of flush end plate in 2nd story 400

1100

340

105

210 190 200

10

60

60

66

70

105

68 69

63

65

210 540

540

65

210

62

67

65

64

210

61

60

60

(a) Strain gauges in frame

24 19 20 15 16

1100

25 26 21 22 17 18 Steel-bars truss top chord Steel-bars truss bottom chord

(h) Strain gauges of SBTD in 2nd story

Fig. 12. Layout of strain gauges.

J. Wang et al. / Engineering Structures 148 (2017) 293–311

Bending deformation

Concrete crushing

(a) Deformation of extended end plate

(b) Slab crack pattern in 1st story

Concrete crushing

Weld fracture

(c) Slab crack pattern in 2nd story

(d) Weld fracture at bottom of CFST column

303

Fig. 13. Failure modes of specimen SCF1.

mance and ductility of the blind bolted circular CFST composite frame was superior to the square CFST composite frame at the same width and steel ratio of column section. However, the lateral resistant load and stiffness of specimen SCF1 were more than those of specimen CCF1 at the same lateral displacement. Because of space limitations, the results of the underlying mechanics and analytical models analysis were detailed in Ref. [43] and were not presented in this paper. 4. Analysis of seismic behavior 4.1. Load-displacement envelope curve The horizontal load-displacement skeleton curve of tested specimens were constructed by connecting maximum load point at each displacement level in accordance with the loaddisplacement hysteretic curves. Fig. 19 showed the effect of column section type on the strength capacity and rigidity of the blind bolted end plate CFST composite frames. It showed that the maximum strength capacity and elastic stiffness of blind bolted square CFST composite frame were larger than those of the circular CFST composite frame. To obtain characteristic points from the load-displacement envelope curves to evaluate quantificationally earthquake resistance of the blind bolted CFST composite frames, three typical characteristic points were introduced in accordance with specification JGJ/T 101 [44], as shown in Fig. 20. Point 1 represents the yield point. Point 2 shows the maximum load (Pm,t) and corresponding displacement (Dm,t) of the composite frame when the lateral load

attained Pm,t. Point 3 is the failure point as the lateral load falls to 85% of Pm,t. Table 5 summaries the three characteristic points of the specimens based on the inter-story envelop curves. Moreover, the test measured maximum inter-story shear and elastic inter-story stiffness of the first and second story for the CFST composite frames were illustrated in Figs. 21 and 22. For the reason that the inertia moment of square CFST column is larger than that of circular CFST column at the same width and steel ratio of column section. The first inter-story maximum strength and elastic rigidity of specimen SCF1 increased by 44.8–52.4% and 27.0–32.0%, compared to specimen CCF1. The test results were also found in the second inter-story shear-drift envelope curves. Meanwhile, the inter-story elastic rigidity of the second story of specimen SCF1 and CCF1 were respectively decreased by 0.2–49.5% and 9.7– 29.4%, compared to the first inter-story elastic rigidity of both specimens. In order to examine the composite action of SBTD concrete slab on the strength and rigidity of the blind bolted CFST composite frames, the pure square CFST frame from Ref. [38] was compared with the specimen SCF1. Comparing with the pure square CFST frame, the first inter-story maximum strength and elastic rigidity of square CFST composite frame considering SBTD concrete slab effect were respectively markedly enhanced by 91.7–158.1% and 101.3–109.1%. These effect were also validated from the second inter-story shear-drift envelope curves. It verified that the presence of SBTD concrete slab is of significant influence on seismic performance and could improve the bearing capacity and elastic rigidity of the CFST composite frame dramatically.

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Bending deformation

Concrete crushing

(a) Deformation of extended end plate

(b) Slab crack pattern in 1st story

Weld fracture

Weld fracture

(c) Weld fracture between primary beam flange and end plate

(d) Weld fracture at bottom of CFST column

Fig. 14. Failure modes of specimen CCF1.

4.2. Strength degradation The strength degradation factor (ki) at the same level load can be used to analyze the strength degradation of the composite frame in accordance with specification JGJ/T 101 [44]. It is defined as follows:

ki ¼ Pij =P 1j

ð1Þ

In which Pij and P1j are the maximum lateral loads at the ith and 1st level loads, respectively, as the relative displacement (D/Dy) reaches to j. The relationship between strength degradation factor (ki) at each same level load and relative column end displacement (D/ Dy) was shown in Fig. 23. ‘PD’ and ‘ND’ respectively mean ‘Positive Direction’ and ‘Negative Direction’ in Fig. 23. There was nonexistence of strength degradation at the level load of 6Dy on account of that the test was terminated at the first cycle of 6Dy. It was found that slight strength degradation tendency appeared as specimens were in elastic stage, while the strength degradation factor at each same load began to step away from 1.0 but no obvious strength degradation when specimens were in elastic-plastic and plastic stage. The strength degradation factor (kj) at overall loads was also introduced to explore the global strength degradation characteristics of tested specimens during the whole loading process. It is expressed as follows:

kj ¼ Pj =P max

ð2Þ

In which Pj is the maximum horizontal load under the jth loading cycle when the relative column end displacement (D/Dy) equals j; and Pmax is the maximum horizontal load during the whole loading process. The relationship between strength degradation factor (kj) at overall loads and the relative horizontal displacement (D/Dy) were illustrated in Fig. 24. For the blind bolted CFST composite frames, each inter-story strength degradation factor at the overall loads raised gradually with the increase of D/Dy; when exceeded the maximum carrying capacity, kj reduced gradually with the increase of D/Dy. The kj reached approximately 0.85 at the failure state. The results illustrated that at ultimate limit state, the strength of blind bolted CFST composite frame was not reduced remarkably and could withstand a larger deformation. 4.3. Stiffness y degradation The stiffness degradation of the test specimens could be measured by the rigidity degradation factor accordance to specification JGJ 101 [44]. The rigidity degradation factor (Kj) of the composite frames is expressed as follows: n X Kj ¼ P ij i¼1

,

n X uij

ð3Þ

i¼1

In which Pij is the maximum horizontal loads at the ith level loads as the relative horizontal displacement (D/Dy) reaches to j. lij is the maximum horizontal displacements at the ith level loads as the rel-

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2000 mm 626 mm

663 mm

3.0

1.0

711 mm

2.0

2.0

1.5 1.0

1.5

2.0

2.5 1.0 3.0

4.0

4.0

1.5

1.5 2.5

1.5 3.0

3.0

2.5 1.0 4.0

3.0 2.0

1.0 2.5

3.0

(a) Slab crack pattern in 1st story

(b) Slab crack pattern in 2nd story Fig. 15. Slab crack pattern of specimen SCF1. Note: the Arabic numerals in figures represent the relative displacement levels (D/Dy) of the loading history in this paper.

ative horizontal displacement (D/Dy) reaches to j. n is the number of loading cycles for each level load. Fig. 25 showed the relationship of stiffness degradation factor of specimens versus relative displacement (D/Dy). ‘PD’ and ‘ND’ mean ‘Positive Direction’ and ‘Negative Direction’, respectively, in Fig. 25. The column section type and end plate type was found to impact on the rigidity of the blind bolted moment CFST composite frame. The first inter-story average rigidity degradation factor of specimen SCF1 at the elastic and failure stages were respectively about 1.27–1.32 and 1.37–1.63 of the specimen CCF1. The results were also could be obtained in the second story. It indicated that the elastic and failure rigidity degradation factor of blind bolted square CFST composite frame were larger than those of circular CFST composite frame at the same width and steel ratio of column section. The second inter-story elastic rigidity degradation factor of specimen SCF1 and CCF1 were respectively about 1.01–1.49 and 1.09– 1.29 of the first story. Nevertheless, the first inter-story failure rigidity degradation factor of specimen SCF1 and CCF1 were respectively about 1.03–1.14 and 1.01–1.06 of the second story. In addition, the failure rigidity degradation factor of the blind bolted square or circular CFST composite frame were respectively about 0.19–0.27 and0.19–0.30 of elastic rigidity degradation factor.

It showed that the blind bolted CFST frame could still withstand lateral load at failure stage. 4.4. Ductility High ductile structure has good plastic deformation capacity and could avoid the occurrence of brittle failure to leave people enough time to escape under the case of earthquake and other disasters. So the ductility is of great significance during the process of structural seismic design. The displacement ductility factor (l) is adopted to study the overall ductility of the type of composite structures. The displacement ductility factor (l) can be calculated as the ratio of failure state displacement (Df,t) to yield state displacement (Dy,t). The formula is shown as follows:

l ¼ Df;t =Dy;t

ð4Þ

In Table 6, hf is the inter-story drift angle as structure reaches to failure state and could be defined as hf = Df,t/H; hy is the interstory drift angle as structure justly reaches to yield state and could be defined as hy = Dy,t/H; H is the height of the columns.

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J. Wang et al. / Engineering Structures 148 (2017) 293–311

2000 mm 858 mm

663 mm

479 mm

3.0

2.5 2.5 3.0 2.0 3.0

3.0

4.0 5.0 2.5

4.0

(a) Slab crack pattern in 1st story 2000 mm 858 mm

663 mm

479 mm

3.0

2.5 2.5 3.0 2.0 3.0

3.0

4.0 5.0 2.5

4.0

(b) Slab crack pattern in 2nd story Fig. 16. Slab crack pattern of specimen CCF1.

800

800 SCF1-1st story

600

400

400

200

200

Ps (kN)

Ps (kN)

600

0 -200

0 -200

-400

-400

-600

-600

-800 -80 -60

-40 -20

0

20

40

60

80

CCF1-1st story

-800 -80 -60 -40 -20

0

20

40

Δs(mm)

Δs(mm)

(a) Specimen SCF1

(b) Specimen CCF1

60

80

Fig. 17. Hysteresis curves of inter-story shear (Ps)-drift (Ds) in 1st story. Note: Ps andDs represent the inter-story shear and corresponding inter-story drift of specimens, respectively.

The first and second inter-story displacement ductility factors (l) of specimen CCF1 were respectively increased by 22.9% and 10.0%, compared with specimen SCF1. The phenomena explained

that the blind bolted circular CFST composite frame exhibited better ductility than that of the square CFST composite frame at the same width and steel ratio of column section. In addition, the duc-

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J. Wang et al. / Engineering Structures 148 (2017) 293–311

600

600 400

SCF1-2nd story

400

Ps (kN)

Ps (kN)

200 0

200 0

-200

-200

-400

-400

-600 -80 -60 -40 -20

0

20

40

60

CCF1-2nd story

-600 -80 -60 -40 -20

80

0

20

40

Δs(mm)

Δs(mm)

(a) Specimen SCF1

(b) Specimen CCF1

60

80

800

800

600

600

400

400

200

200

Ps (kN)

Ps (kN)

Fig. 18. Hysteresis curves of inter-story shear (Ps)-drift (Ds) in 2nd story.

0 -200 SCF1-1st story CCF1-1st story

-400

50

-800 -150

-400 -600 -800 -150

-100

-50

0

0 -200

100

150

SCF1-2nd story CCF1-2nd story

-600 -100

-50

0

50

Δs(mm)

Δs(mm)

(a) 1st story

(b) 2nd story

100

150

Fig. 19. Skeleton curves of inter-story shear (Ps)- drift (Ds).

tility factor of the first inter-story was less than that of the second inter-story for all specimens. Table 6 summarizes the ductility coefficient of the test CFST composite frame. It showed that the inter-story displacement ductility factor (l) of the specimen SCF1 and CCF1 is l = 3.03–4.52. The Chinese building seismic design code GB 50011 [45] has detail ductility regulation for the multi- and high-rise steel building structures: the elastic inter-story drift [he] = 1/250, and the elastic-plastic inter-story drift [hp] = 1/50. However, presently detailed ductility regulations of the blind bolted end plate CFST composite frame is lacking for building structures. Based on the specification GB 50011 [45], the elastic and the elastic-plastic inter-story drift of specimen SCF1 are respectively hy = (1.58– 3.91)[he] and hf = (1.17–2.50)[hp]; the elastic and the elasticplastic inter-story drift of specimen CCF1 are respectively hy = (1.52–3.15)[he] and hf = (1.34–2.55)[hp]. The test results showed

P 2

Pm,t

3

Pf,t Py,t

1

O

Δy,t

Δm,t

Δf,t

Δ

Fig. 20. Characteristic points of tested specimens.

Table 5 Characteristic values on envelop curves. Specimen

SCF1

Yielding point

1st inter-story 2nd inter-story

CCF1

1st inter-story 2nd inter-story

Maximum point

Failure point

Dy,t (mm)

Py,t (kN)

Dm,t (mm)

Pm,t (kN)

Df,t (mm)

Pf,t (kN)

(+) () (+) ()

23.06 21.75 10.74 9.77

440.41 448.06 306.60 201.75

51.09 63.41 29.3 36.95

758.88 704.36 447.41 433.06

69.96 73.89 36.19 44.19

645.07 600.12 379.64 368.11

(+) () (+) ()

18.6 17.94 10.09 9.43

269.22 291.21 188.88 167.84

50.44 50.44 30.26 28.95

497.79 486.31 302.69 291.21

75.20 69.30 43.23 41.5

426.06 414.58 254.87 248.18

Note: ‘(+)’ and ‘()’ mean ‘Positive Direction’ and ‘Negative Direction’, respectively.

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900 800

500

758.88

700

433.06

400

600

497.79

486.31

Ps (kN)

Ps (kN)

447.41

704.36

500 400 300 200

302.69

291.21

300 200 100

100 0

0

SCF1+

CCF1+

SCF1-

CCF1-

SCF1+

CCF1+

(a) 1st story

SCF1-

CCF1-

(b) 2nd story

Fig. 21. Comparison on maximum inter-story shear of composite frames.

35

25

20.60

19.10

28.55

30

16.23

14.47

K (kN/mm)

K (kN/mm)

20 15 10

25

20.65 18.72

20

17.8

15 10

5

5

0

0

SCF1+

CCF1+

SCF1-

CCF1-

SCF1+

CCF1+

(a) 1st story

SCF1-

CCF1-

(b) 2nd story

Fig. 22. Comparison on elastic inter-story stiffness of composite frames.

1.2

1.2

0.8

0.8

0.4

0.4

SCF1-PD SCF1-ND CCF1-PD CCF1-ND

0.0

0.0

-0.4

-0.4

-0.8

-0.8

-1.2

0

1

2

3

4

5

SCF1-PD SCF1-ND CCF1-PD CCF1-ND

6

-1.2

0

1

(a) 1st story

2

3

4

5

6

(b) 2nd story

Fig. 23. Strength degradation factor at the same loads. Note: ‘PD’ and ‘ND’ respectively mean ‘Positive Direction’ and ‘Negative Direction’.

1.2

1.2

0.8

0.8

0.4

0.4

SCF1-PD SCF1-ND CCF1-PD CCF1-ND

0.0

0.0

-0.4

-0.4

-0.8

-0.8

-1.2

0

1

2

3

4

(a) 1st story

5

6

SCF1-PD SCF1-ND CCF1-PD CCF1-ND

7

-1.2

0

1

2

3

4

(b) 2nd story

Fig. 24. Strength degradation factor at the total loads.

5

6

7

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J. Wang et al. / Engineering Structures 148 (2017) 293–311

40000

45000 SCF1-PD SCF1-ND CCF1-PD CCF1-ND

20000 10000 0

SCF1-PD SCF1-ND CCF1-PD CCF1-ND

36000

Kj(kN/m)

Kj(kN/m)

30000

27000 18000 9000

0

1

2

3

4

5

6

7

0

0

1

(a) 1st story

2

3

4

5

6

7

(b) 2nd story

Fig. 25. Rigidity degradation factor.

Table 6 Ductility coefficient of specimens.

Dy,t (mm)

Df,t (mm)

hy

hf

l

(+) () (+) ()

23.06 21.75 10.74 9.77

69.96 73.89 36.19 44.19

1/64 1/68 1/144 1/159

1/21 1/20 1/43 1/35

3.03 3.40 3.37 4.52

(+) () (+) ()

18.60 17.94 10.09 9.43

75.20 69.30 43.23 41.50

1/79 1/82 1/154 1/164

1/20 1/21 1/36 1/37

4.04 3.86 4.28 4.40

Specimen SCF1

1st inter-story 2nd inter-story

CCF1

1st inter-story 2nd inter-story

curves at each displacement level for both specimens. The energy dissipation parameters of specimens at ultimate state were listed in Table 7. The energy-dissipating capacity (Ee) shown in Table 7 is defined as:

B

Ee ¼ 2pne

F A O

C E

D Fig. 26. Idealized P-D hysteretic relationship.

that all of them are higher than those of the inter-story drift angle limitation of steel structural building specified in GB 50011 [45], and declared that the blind bolted CFST composite structures displayed better ductility and could satisfy the seismic design requirements of structures. 4.5. Dissipated energy The structural hysteretic loops can reflect the energy dissipation capacity. The more structure dissipates energy, the more safe structure and the less likely to be damaged. The equivalent damping factor (ne) was defined by Eq. (5), which expressed in specification JGJ/T 101 [44]. The parameters in Eq. (5) are illustrated in Fig. 26. SABC and SCDA are respectively areas of the hysteresis curve of ABC and CDA, and SOBE and SODF are respectively triangle areas of OBE and ODF.

ne ¼

1 SABC þ SCDA 2p SOBE þ SODF

ð5Þ

Fig. 27 showed the equivalent damping factor (ne) versus relative horizontal displacement (D/Dy) relationship of all specimens. Fig. 28 showed the comparison of energy dissipation of hysteretic

ð6Þ

Fig. 27 indicated that the cumulative equivalent damping factor (ne) of the blind bolted end plate CFST composite frames increased with increasing relative displacement (D/Dy). The first and second inter-story cumulative equivalent damping factor (ne) of the blind bolted square CFST composite frame were slightly larger than those of the circular CFST composite frame when 0 < D/Dy  1.5, but the cumulative equivalent damping factor (ne) in opposition when D/Dy > 1.5. However, Fig. 28 showed that each inter-story energy dissipation of hysteretic loops of the blind bolted square CFST composite frame was larger than that of the circular CFST composite frame except loading cycles of 5.0 D/Dy and 6.0 D/Dy. At the failure stage, more sharp level of reduction on the stiffness and strength occurred for the square CFST composite frame. Thus, the energy dissipation of hysteretic loops of blind bolted square CFST composite frame was lower than that of the circular CFST composite frame at the loading cycles of 5.0 D/Dy and 6.0 D/Dy. Table 7 provided the total dissipation energy capacity of both specimens. It showed that the total dissipation energy (Wtotal) of the square CFST composite frame was more than that of the circular CFST composite frame, while the equivalent damping factor (ne) and the energy-dissipating capacity (Ee) of the circular CFST composite frame were larger than those of the square CFST composite frame. Although the second inter-story displacement was lower than the first inter-story displacement which resulted in lower total dissipation energy (Wtotal) of the second story, the equivalent damping factor (ne) and the energy-dissipating capacity (Ee) of the second inter-story were larger than those of the first inter-story. The blind bolted end plate CFST composite frames exhibited good seismic behavior. At the same width and steel ratio of column section, energy dissipation analysis indicated that the blind bolted square CFST composite frame possessed larger stiffness and

J. Wang et al. / Engineering Structures 148 (2017) 293–311

0.20

0.20

0.15

0.15

ξe

ξe

310

0.10 0.05 0.00

0.05

SCF1-1st story CCF1-1st story

0

1

2

3

4

5

6

0.10

0.00

7

SCF1-2nd story CCF1-2nd story

0

1

(a) 1st story

2

3

4

5

6

7

(b) 2nd story

Fig. 27. ne versus D/Dy relationship.

30000

80000

SCF1-1st story CCF1-1st-story

SCF1-2nd story CCF1-2nd-story

W (kN.mm)

W (kN.mm)

60000

40000

20000

10000

20000

0 0.125 0.25 0.5 0.7 1.0 1.5 2.0 2.5 3.0 4.0 5.0 6.0

0 0.125 0.25 0.5 0.7 1.0 1.5 2.0 2.5 3.0 4.0 5.0 6.0

(a) 1st story

(b) 2nd story

Fig. 28. Comparison of energy dissipation of hysteretic loops in the test specimens.

Table 7 Energy dissipation parameters of specimens at ultimate state. Specimen SCF1 CCF1

1st inter-story 2nd inter-story 1st inter-story 2nd inter-story

D/Dy

Wtoal (kNmm)

ne

Ee

4 4 4 4

94971.31 37879.67 81937.91 35002.09

0.103 0.125 0.127 0.143

0.652 0.790 0.801 0.901

strength than that of the circular CFST composite frame and dissipated more hysteretic energy at each loading level. Whereas the blind bolted circular CFST composite frame behaved better seismic performance in terms of ductility, equivalent damping factor (ne) and energy-dissipating capacity (Ee). 5. Conclusions This paper reported a series of cyclic loading test of the seismic behavior of blind bolted CFST composite frame between square or circular CFST columns and SBTD concrete beams. The test findings obtained may be summarized within the limitation of the study: (1) For the blind bolted CFST composite frames between square or circular CFST columns and SBTD composite beams, the failure modes include: deformation of end plate; fracture at the junction between the steel beam and the end plate; local buckling at top and bottom flanges of steel beam in the joints; fracture at the bottom end of CFST column; and cracking and local crushing of concrete slab.

(2) The strength and rigidity of the blind bolted CFST composite frames with SBTD concrete slabs are affected by the column section type. The strength and rigidity of square CFST composite frame were higher than those of the circular CFST composite frame under the same width and steel ratio of column section. It is also found that the presence of SBTD concrete slab could dramatically improve the strength and rigidity of the blind bolted CFST column to steel beam frames. (3) All blind bolted CFST composite frames exhibited favorable hysteretic performance and large lateral deformation capacity. The elastic and the elastic-plastic inter-story drift of the blind bolted CFST composite frames is respectively hy = (1.52–3.91)/50 and hf = (1.17–2.55)/250. It satisfied the ductility requirement in some aseismic region. (4) The blind bolted CFST composite frames displayed large energy dissipation capacity. The total dissipation energy (Wtotal) of square CFST composite frame were larger than that of the circular CFST composite frame, while the equiva-

J. Wang et al. / Engineering Structures 148 (2017) 293–311

lent damping factor (ne) and the energy-dissipating capacity (Ee) of square CFST composite frame were lower than those of the circular CFST composite frame.

Acknowledgements This work described in paper is supported by the National Natural Science Foundation of China (Project 51478158 and Project 51178156) and the New Century Excellent Talents in University (Project NCET-12-0838), which is greatly appreciated. The authors would also like to acknowledge the assistance of Bo Wang, Zheng Sun and Xuezhou Wu of Hefei University of Technology, who helped to conduct the experiments. References [1] Chen WF, Lui EM. Stability design of steel frames. CRC Press Inc., 2000 Corporate Blvd., N. W., Boca Raton, Florida; 1991. [2] Loh HY, Uy B, Bradford MA. The effects of partial shear connection in composite flush end plate joints Part I-experimental study. J Constr Steel Res 2006;62 (4):378–90. [3] Ataei A, Bradford MA, Valipour HR. Experimental study of flush end plate beam-to-CFST column composite joints with deconstructable bolted shear connectors. Eng Struct 2015;99:616–30. [4] Wang JF, Zhang HJ, Jiang Z. Seismic behavior of blind bolted end plate composite joints to CFTST columns. Thin-Walled Struct 2016;108:256–69. [5] Choi IR, Chung KS, Kim CS. Experimental study on rectangular CFT columns with different steel grades and thicknesses. J Constr Steel Res 2017;130:109–19. [6] Elchalakani M, Karrech A, Hassanein MF, Yang B. Plastic and yield slenderness limits for circular concrete filled tubes subjected to static pure bending. ThinWalled Struct 2016;109:50–64. [7] Kim D, Jeon C, Shim C. Cyclic and static behaviors of CFT modular bridge pier with enhanced bracings. Steel Compos Struct 2016;20(6):1221–36. [8] Montuori R, Piluso V. Analysis and modelling of CFT members: moment curvature analysis. Thin-Walled Struct 2015;86:157–66. [9] Chacón R. Circular Concrete-Filled Tubular Columns: state of the art oriented to the vulnerability sssessment. Open Civil Eng J; 2015: 9(1–M4); 249–259. [10] Goto Y, Ebisawa T, Lu X. Local buckling restraining behavior of thin-walled circular CFT columns under seismic loads. J Struct Eng 2014;140(5):04013105. [11] Skalomenos KA, Hatzigeorgiou GD, Beskos DE. Parameter identification of three hysteretic models for the simulation of the response of CFT columns to cyclic loading. Eng Struct 2014;61(1):44–60. [12] Korol RM, Ghobarah A, Mourad S. Blind bolting W-shape beams to HSS columns. J Struct Eng 1993;119(12):3463–81. [13] France JE, Davison JB, Kirby PA. Strength and rotational stiffness of simple connections to tubular columns using flow drill connectors. J Constr Steel Res 1999;50(1):15–34. [14] Lee J, Goldsworthy HM, Gad EF. Blind bolted moment connection to unfilled hollow section columns using extended T-stub with back face support. Eng Struct 2011;33(5):1710–22. [15] Wang JF, Spencer BF. Experimental and analytical behavior of blind bolted moment connections. J Constr Steel Res 2013;82(82):33–47. [16] Wang JF, Zhang N, Guo SP. Experimental and numerical analysis of blind bolted moment joints to CFTST columns. Thin-Walled Struct 2016;109:185–201. [17] Wang JF, Zhang N. Performance of circular CFST column to steel beam joints with blind bolts. J Constr Steel Res 2017:36–52. [18] Yao H, Goldsworthy H, Gad E. Experimental and numerical investigation of the tensile behavior of blind-bolted T-stub connections to concrete-filled circular columns. J Struct Eng 2014;134(2):198–208.

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