Investigation on square concrete filled double-skin steel tube (CFDST) subjected to local bearing force: Experiments

Thin-Walled Structures 94 (2015) 394–409

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Investigation on square concrete filled double-skin steel tube (CFDST) subjected to local bearing force: Experiments You-Fu Yang a, Chao Hou b, Chun-Yuan Meng a, Lin-Hai Han b,n a b

State Key Laboratory of Coastal and Offshore Engineering, Dalian University of Technology, Dalian, 116024, China Department of Civil Engineering, Tsinghua University, Beijing, 100084, China

art ic l e i nf o

a b s t r a c t

Article history: Received 9 April 2015 Received in revised form 22 April 2015 Accepted 22 April 2015 Available online 21 May 2015

Experimental studies on the behaviour of concrete filled double-skin steel tube (CFDST) subjected to local bearing forces are presented in this paper. Sixteen specimens were prepared and tested with the included angle between bearing member (BM) and compression member of 451 and 901, whilst both the inner and outer steel tubes of the CFDST specimens are square hollow sections (SHS). The main parameters in the tests were: 1) outside width ratio between BM and compression member: from 0.4 to 0.6; 2) hollow ratio of CFDST: from 0 to 0.6; 3) wall thickness of outer steel tube: 3.05 mm and 3.95 mm; and 4) cross-section of BM: solid and hollow. The failure pattern, load versus deformation curve, bearing capacity and corresponding deformation at bearing capacity of the tested specimens are presented and analyzed. The experimental results show that, while subjected to local bearing forces, CFDST specimens have a high bearing capacity and a good deformation-resistant ability. The calculated bearing capacities of CFDST under local bearing forces using the proposed formulae in the paper are evaluated by comparison with the experimental results. & 2015 Elsevier Ltd. All rights reserved.

Keywords: Concrete filled double-skin steel tube (CFDST) Concrete filled steel tube (CFST) Square section Local bearing force Bearing capacity

1. Introduction Concrete filled double-skin steel tube (CFDST) has the similar performance as traditional concrete filled steel tube (CFST) under the same dimensions of outer steel tube and strength of materials, while having a lighter weight due to the absence of concrete filling inside the inner tube [1]. As a result, CFDST has the potential advantages in some engineering applications, such as bridge piers over deep valleys, offshore platform support columns, base connecting segments of offshore wind turbine, large diameter columns of building structures and other towering structures or column members. The experimental and theoretical studies of CFDST have been widely conducted, and much of the existing research has been summarized in [2–4]. Recently, some new studies were carried out on the experimental behaviour and nonlinear simulation of CFDST stub columns subjected to axial loading [5–8], CFDST beams [9], CFDST members subject to pure torsion [10], partially loaded CFDST columns [4], CFDST columns under long-term sustained loading [11] and CFDST columns under fire [12].

n

Corresponding author. Tel.: /fax: þ86 10 62797067. E-mail addresses: [email protected], [email protected] (L.-H. Han).

http://dx.doi.org/10.1016/j.tws.2015.04.026 0263-8231/& 2015 Elsevier Ltd. All rights reserved.

For tubular structures, grouting of joints can be adopted to obtain improved strength and stiffness [13], and the chord would become a CFST or a CFDST by grouting whole hollow section or annulus between the chord and an internal tube, respectively. In practice, the chords of a tubular structure may be subjected to local bearing forces transmitted by the brace members, and the performance of chord under local bearing forces is a very important issue for ensuring the structural safety. Hou et al. [14–16] introduced the tests and finite element analysis of circular CFST and CFDST under local bearing forces. The test results of square CFST under local bearing forces were presented in Yang et al. [17]. Feng and Young [18–19] carried out the experimental and theoretical studies of concrete filled stainless steel tubular joints with square and rectangular brace and chord members subjected to compression. The performance and design calculations of rectangular CFST under local bearing forces were reported by Packer [20]. Till now, however, there is a lack of knowledge regarding the behaviour of square CFDST members under local bearing forces. From the aforementioned information, it is clear that the research on performance of square CFDST under local bearing forces needs to be conducted to guide the engineering design. The aim of this study is thus to investigate the experimental behaviour of square CFDST under local bearing forces with the loads applied by steel bearing members (BM), both solid and

Y.-F. Yang et al. / Thin-Walled Structures 94 (2015) 394–409

β

Nomenclature bb bi bo f c' L0 Lb P P uc P ue tb ti to

outside width of bearing member outside width of inner tube outside width of outer tube cylinder compressive strength of concrete length of the specimen length of bearing member load calculated bearing capacity experimental bearing capacity wall thickness of hollow bearing member wall thickness of inner tube wall thickness of outer tube

χ δc Δl Δl;ue Δv Δv;ue ε εue εy θ

width ratio between bearing member and compression member hollow ratio protrusion of concrete lateral displacement lateral displacement corresponding to bearing capacity vertical displacement vertical displacement corresponding to bearing capacity strain strain corresponding to bearing capacity yield strain of outer steel tube included angle between bearing member and compression member

 Cross-section of BM: solid and hollow.

hollow BM are adopted in the tests. The tests of 8 specimens with the included angle between BM and compression member (θ) of 901 and 8 specimens with θ of 451 were carried out to evaluate the effects of outside width ratio between BM and compression member, hollow ratio and wall thickness of outer steel tube on the performance of square CFDST under local bearing forces. The experimental results are also used to evaluate the accuracy of the calculated bearing capacity of square CFDST under local bearing forces using the proposed formulae in this paper.

where, bb is the outside width of BM. The hollow ratio (χ) is defined as:

2. Experimental program

χ¼

2.1. Test specimens Sixteen CFDST specimens with inner and outer steel tubes of square hollow sections (SHS) were prepared, and the length of the specimens (L0 ) was 800 mm, consistent with that in Hou et al. [16]. The cross-section of CFDST adopted in the tests is shown in Fig. 1, where bo and bi are the outside width of outer and inner tube, respectively, and t o and t i are the wall thickness of outer and inner tube, respectively. All specimens were transversely compressed by square steel bearing member (BM) to failure with the included angle between BM and compression member (θ) of 451 and 901, whilst no external forces were applied in the longitudinal direction. The experiments were carried out to study the effects of the following parameters on the performance of square CFDST under local bearing forces:

 Width ratio between BM and compression member, β: from 0.4 to 0.6;

 Hollow ratio, χ: from 0 to 0.6;  Wall thickness of outer steel tube, t o : 3.05 mm and 3.95 mm; and Outer tube

to

bi

ti

Concrete Inner tube bo Fig. 1. Cross-section of CFDST.

395

For CFDST under local bearing forces, the width ratio between BM and compression member (β) is given by: β¼

bb bo

bi bo  2t o

ð1Þ

ð2Þ

The summary of the tested specimens is presented in Table 1, where t b is the wall thickness of hollow BM, Lb is the length of BM for specimens with θ of 901 and is the nominal length of BM equals to (80þbb/2) mm for specimens with θ of 451, P ue is the experimental bearing capacity, and Δv;ue and Δl;ue are the measured vertical and lateral displacement corresponding to P ue , respectively. In Table 1, the first part of specimen label represents the information of compression member, where ‘T4’ and ‘T3’ refer to the wall thickness of outer steel tube of 3.95 mm and 3.05 mm, respectively, and the numbers in the brackets denote the hollow ratio. The second part of specimen label represents the message of BM, where ‘S’ and ‘H’ refer to solid and hollow BM, respectively, and the numbers denote the included angle between BM and compression member (θ). The last part of specimen label represents the outside width ratio between BM and compression member (β). The outer and inner SHS of the specimens were welded from two U-shape channels produced by mild steel plate, and there was no endplate at the ends of the specimens to allow the protrusion of concrete under local bearing forces. The hollow BM was cold-formed square steel tube. The properties of steel were obtained by testing three tensile coupons, and the measured yield strength, tensile strength, elastic modulus, Poisson’s ratio and elongation after fracture are given in Table 2. For all specimens, the annular space between outer and inner tubes was filled using the same self-consolidating concrete (SCC) batch for consistency. The mix proportions of SCC were: Cement: 380 kg/m3; Fly ash: 170 kg/m3; Coarse aggregate: 840 kg/m3; Sand: 840 kg/m3; Water: 173 kg/m3; and Water reducing agent: 11.0 kg/m3. The fresh SCC had the slump and spreading of 250 mm and 530 mm, respectively. The cube compressive strength and elastic modulus of SCC were obtained by testing cubes with a side length of 150 mm and prisms of size 150 mm  150 mm  300 mm, respectively. The measured cube compressive strength at 28 days and when loading tests conducted were 52.4 and 59.4 N/mm2, respectively. The elastic modulus of SCC was 34,500 N/mm2.

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Table 1 Summary of the tested specimens. No.

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16

Specimen label

T4(0.6)-S90-0.6 T4(0.6)-S90-0.5 T4(0.6)-S90-0.4 T3(0.6)-S90-0.6 T4(0.6)-H90-0.6 T4(0.45)-S90-0.6 T4(0.3)-S90-0.6 T4(0)-S90-0.6 T4(0.6)-S45-0.6 T4(0.6)-S45-0.5 T4(0.6)-S45-0.4 T3(0.6)-S45-0.6 T4(0.6)-H45-0.6 T4(0.45)-S45-0.6 T4(0.3)-S45-0.6 T4(0)-S45-0.6

Compression member

Bearing member (BM)

bo (mm)

to (mm)

bi (mm)

ti (mm)

L0 (mm)

bb (mm)

tb (mm)

Lb (mm)

240

3.95 3.95 3.95 3.05 3.95 3.95 3.95 3.95 3.95 3.95 3.95 3.05 3.95 3.95 3.95 3.95

140 140 140 140 140 105 70 – 140 140 140 140 140 105 70 –

3.05 3.05 3.05 3.05 3.05 3.05 3.05 – 3.05 3.05 3.05 3.05 3.05 3.05 3.05 –

800

145 120 95 145 145 145 145 145 145 120 95 145 145 145 145 145

– – – – 4.8 – – – – – – – 4.8 – – –

80 80 80 80 80 80 80 80 152.5 140 127.5 152.5 152.5 152.5 152.5 152.5

Table 2 Properties of steel. Thickness (mm)

Tensile Yield strength (N/ strength (N/ 2 mm2) mm )

Elastic modulus (N/ mm2)

Possion's ratio

Elongation (%)

3.95 3.05 4.8

442.4 390.5 379.4

1.91  105 1.84  105 1.97  105

0.276 0.269 0.284

21.8 23.1 24.4

515.4 533.1 521.4

2.2. Test setup Experiments were performed on a 10000 kN capacity servo control testing machine. The test setup was similar to that used in the tests of square CFST specimens under local bearing forces [17], as shown in Fig. 2, where P is the load applied to the top of BM. The vertical and lateral displacements (Δv and Δl ) were measured by the displacement transducer (DT). Two vertical DTs located symmetrically on both sides of the specimen were placed on the symmetric axis of the bottom platen of the testing machine to measure the vertical displacement, and two more lateral DTs were placed on the center of the side walls of CFDST specimens to measure the corresponding lateral displacement. Furthermore, the strain gauges (SGs) were pasted on sixteen positions of outer tube to monitor the strain variation during loading process, and there were one longitudinal SG and one transverse SG at each position, as shown in Fig. 3, where only one position was numbered when there were two or more symmetric positions. The displacements and strains were collected by a data acquisition system in the laboratory.

χ

β

θ (1)

Pue (kN)

Δv,ue (mm)

Δl,ue (mm)

0.6 0.6 0.6 0.6 0.6 0.45 0.3 0 0.6 0.6 0.6 0.6 0.6 0.45 0.3 0

0.6 0.5 0.4 0.6 0.6 0.6 0.6 0.6 0.6 0.5 0.4 0.6 0.6 0.6 0.6 0.6

90 90 90 90 90 90 90 90 45 45 45 45 45 45 45 45

898.4 619.4 507.0 844.7 865.9 1223.7 2127.5 2944.9 1709.6 1471.2 1026.1 1787.5 1176.7 2556.9 3411.3 4366.1

4.18 3.07 11.87 3.56 5.80 3.82 3.86 5.97 5.97 11.55 11.73 7.01 6.4 7.46 9.86 12.4

1.38 1.24 0.56 1.24 1.26 1.91 1.32 1.69 0.54 2.13 1.08 0.68 0.08 0.69 NA 0.91

different from that of others, i.e. there is no evident web buckling of outer tube and the top flange of outer tube is completely sheared to failure. This may be due to the fact that, corner rotation of outer tube does not obviously appear for specimen T4(0.6)-S900.4 due to a large confinement of outer tube and non-loaded concrete to the directly loaded concrete, simultaneously, a large downward displacement of concrete beneath the BM is produced owing to the premature failure of top flange of inner tube. For most of specimens with θ of 451, local buckling of webs and top flange of outer tube close to the bottom side of BM can be observed; however, for specimen T4(0.6)-S45-0.4, local buckling of webs of outer tube does not obviously happen yet, which is similar to the observed phenomena of specimen T4(0.6)-S90-0.4. For CFDST specimens with θ of 451, there is no crack of top flange of outer tube, which is different from that of CFST specimens in [17]. This can be explained that the tensile strain of top flange of outer tube close to the top side of BM becomes smaller due to the existence of inner tube. It can also be seen from Fig. 4(b3) that, for specimen with θ of 451 loaded by hollow BM, there is no obvious damage to the specimen except for local buckling of hollow BM like a square steel tube under concentric compression, which means that the bearing capacity of this specimen is far larger than the local buckling strength of hollow BM. In general, for all specimens, the buckling range and peak lateral displacement of outer tube web increases with the increasing β, owing to the enlarged loading area of concrete beneath the BM, whilst t o and χ have no obviously effect on the overall failure pattern of the tested specimens. Moreover, similar to CFST specimens in [17], protrusion of the sandwiched concrete occurs as there is no endplate at the ends of the tested specimens.

3. Experimental results and discussions 3.1. Failure pattern 3.1.1. Overall CFDSTs Fig. 4 shows the observed failure pattern of the tested specimens. It is demonstrated that, the overall failure pattern of CFDST specimens is similar to that of CFST specimens (χ ¼ 0) in this paper and in [17]. For specimens with θ of 901, in general, the top flange of outer tube was sheared to failure along four sides of BM and the indentation of tube beneath the BM was also caused; meanwhile, the local buckling of webs of outer tube happened due to the rotation of tube corners and the sideward expanding of the sandwiched concrete. However, the failure pattern of one specimen with a relative small β (i.e. specimen T4(0.6)-S90-0.4) is

3.1.2. Steel tubes Fig. 5 demonstrates the failure pattern of inner tube of the tested specimens. It can be seen that, generally, the failure pattern of inner tube is similar to steel SHS subjected to local bearing forces with β greater than 0.8 [21], i.e. the top flange beneath the BM deforms downwards locally and the webs buckle inward due to the existence of the sandwiched concrete. Moreover, the length of indentation of inner tube is larger than the width of BM due to the diffusion of load from the sandwiched concrete. These mean that the forces applied to the BM during the loading process can be directly dispersed to the inner tube. However, owing to the local buckling failure of the hollow BM, the failure of inner tube of specimen T4(0.6)-S45-0.4 did not happen. Generally, the failure range of inner tube increases with

Y.-F. Yang et al. / Thin-Walled Structures 94 (2015) 394–409

P

P BM

Specimen

BM

Lb

Specimen

397

bb

DT bo

DT

DT

DT L0/2

L0/2

bo

P BM

P

BM bb

bo

Lb

Weld

Specimen

Weld Specimen

DT

DT L0/2

Steel roller

bo Loading pedestal L0/2 DT Loading pedestal

Steel roller

DT

DT

1.0a

1.0a

Fig. 2. Schematic diagram of the test setup. (a) θ=901. (b) θ=451.

a=10 mm

1

Longitudinal

6

1.0a

5

1.0a

3

a=10 mm

4 Transverse

2.5a

2 1.0a

240 mm

240 mm

240 mm

(2) Side view

5

1

4 5.0a

240 mm

a=10 mm 6

a=10 mm

7 9

8

10

1.0a

3

2 1.5a

240 mm

(1) Top view

240 mm

2.5a

1.0a

1.0a

(1) Top view

240 mm

240 mm

(2) Side view

Fig. 3. Arrangement of strain gauges. (a) θ ¼ 90o . (b) θ ¼ 45o .

increasing β and χ due to the enlarged loaded area of inner tube. Moreover, the inner tube of specimens with θ of 451 has a larger indentation area than that of specimens with θ of 901, because of a larger contact area between BM and compression member.

It was found that the failure of the outer and inner tubes was generally focused on the middle half length of CFDST specimens, as shown in Fig. 6. The typical failure pattern of the outer tubes included the indentation of top flange beneath the BM and the sideward

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Y.-F. Yang et al. / Thin-Walled Structures 94 (2015) 394–409

T4(0.6)-S90-0.6

T4(0.6)-S90-0.5

T4(0.6)-S90-0.4

T4(0.6)-S90-0.6

T3(0.6)-S90-0.6 Protrusion

Protrusion

Protrusion Buckling

Buckling

Buckling

Shear failure

Shear failure

Shear failure

(1)

T4(0.6)-S90-0.6

T4(0.6)-H90-0.6 Protrusion

(2)

T4(0.6)-S90-0.6 T4(0.45)-S90-0.6 T4(0.3)-S90-0.6 T4(0)-S90-0.6 Protrusion Protrusion Protrusion

Buckling

Buckling

Buckling

Shear failure

Shear failure

Shear failure

(3)

Shear failure

(4)

T4(0.6)-S45-0.4 T4(0.6)-S45-0.6 T4(0.6)-S45-0.5 Protrusion Protrusion

T4(0.6)-S45-0.6

Buckling

(1)

T4(0.6)-S45-0.6 T4(0.6)-H45-0.6 Protrusion

T3(0.6)-S45-0.6 Protrusion

Buckling

Buckling

Buckling

Buckling

(2)

T4(0.6)-S45-0.6 T4(0.45)-S45-0.6 T4(0)-S45-0.6 T4(0.3)-S45-0.6 Protrusion Protrusion Protrusion Buckling

(3)

Buckling

Buckling

(4) Fig. 4. Failure pattern of the tested specimens. (a) θ ¼ 901. (b) θ ¼451.

buckling of webs, which were similar to those of CFST specimens [17], and the shear failure appeared if there was no weld between BM and outer tube, as shown in Fig. 6(a). Furthermore, the buckling mode and

buckling range of outer tubes with θ of 451 were different from those with θ of 901. The typical failure pattern of inner tubes is similar to that of SHS under local bearing forces [21], including the indentation of top

Y.-F. Yang et al. / Thin-Walled Structures 94 (2015) 394–409

T4(0.6)-S90-0.6

T4(0.6)-S90-0.5

T4(0.6)-S90-0.4

399

T4(0.6)-S90-0.6

T3(0.6)-S90-0.6

(1) T4(0.6)-S90-0.6

(2) T4(0.6)-H90-0.6

T4(0.6)-S90-0.6

T4(0.45)-S90-0.6

(3)

T4(0.6)-S45-0.6

(4)

T4(0.6)-S45-0.5

T4(0.6)-S45-0.4

T4(0.6)-S45-0.6

(1) T4(0.6)-S45-0.6

T4(0.3)-S90-0.6

T3(0.6)-S45-0.6

(2)

T4(0.6)-H45-0.6

T4(0.6)-S45-0.6

T4(0.45)-S45-0.6

(3)

T4(0.3)-S45-0.6

(4) Fig. 5. Failure pattern of inner tube. (a) θ ¼ 901. (b) θ ¼ 451.

flange under load diffused from the sandwiched concrete, the buckling of two top corners and the inward local buckling of webs due to the existence of the sandwiched concrete, as shown in Fig. 6(b).

3.1.3. Sandwiched concrete The failure pattern of the sandwiched concrete is shown in Fig. 7. It can be seen that, for specimens with θ of 901, the tensile cracks near the

BM can be found in the length direction and the sandwiched concrete is generally crushed within the buckling range of webs of outer tubes except for one specimen with a relative small β (i.e. specimen T4(0.6)S90-0.4). For specimens with θ of 451, the sandwiched concrete is generally crushed within the buckling range of outer tube webs and bottom side of BM, whilst the tensile cracks near the BM can be found at the top side and corners. For specimen T4(0.6)-H45-0.6 (see Fig. 7 (b3)), concrete almost does not fail except for a rectangular trace as local

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Buckling

Top view

Top view

Buckling

Shear failure Indentation

Indentation

Side view

Side view

Buckling

Buckling

Buckling

Buckling Top view

Top view Indentation

Indentation

Side view

Side view Buckling

Buckling

Fig. 6. Typical failure pattern of outer and inner tubes. (a) Outer tubes. (b) Inner tubes.

buckling failure of hollow BM happens before the bearing capacity of CFDST is reached. In general, the hollow ratio (χ) has a moderate influence on the failure pattern of the sandwiched concrete. 3.1.4. Force transfer mechanism Comparisons of force transfer mechanism between square CFDST under local bearing forces in this paper and other composite specimens under local bearing or locally axial loading, i.e., in [3,16–18], are demonstrated in Fig. 8. It was shown that the load transfer mainly depended on the stress diffusion and generally the dispersed bearing area was larger than the bearing area. However, due to the difference in the profile and loading condition of the members, partially loaded square CFDST columns (Fig. 8(d)) had the smallest diffusion angle and the longest transfer path [4]. It can be found from Fig. 8 (a)-(c) that, while subjected to local bearing forces, CFDST might have a larger diffusion angle and a shorter transfer path compared with CFST as the stress diffusion in length direction was stopped by the inner tube before reaching the side walls. Moreover, compared with circular CFDST under local bearing forces, square CFDST had a larger diffusion angle and a longer transfer path owing to the weaker confinement of square steel tube to the sandwiched concrete. 3.2. Load versus deformation curves The load (P) versus displacement (Δv and Δl ) curves of the tested specimens are shown in Fig. 9, where P Δl curve of specimen T4(0.3)S45-0.6 is absent due to the signal loss of data acquisition channel. The vertical displacements (Δv ) and lateral displacements (Δl ) in Fig. 9 are considered as positive and negative, respectively. It can be seen that,

similar to square CFST specimens under local bearing forces [17], Δv is larger than Δl under the same load, and the resultant vertical displacement of specimens with θ of 451 is larger than the component vertical displacement of specimens with θ of 901. In general, the initial slope of load versus displacement curve increases with decreasing hollow ratio (χ), except for P  Δv curve of CFST specimens (χ ¼0), as well as increasing width ratio (β). It can also be seen from Fig. 9(a) that, there are the second peak load for specimens with the relatively small β, including specimens T4(0.6)-S90-0.5, T4(0.6)-S90-0.4 and T4(0.6)-S450.4. This may be explained that, for these specimens, only the early failure of inner tube happens, meanwhile, the concrete beneath the BM does not reach its strength at the first peak load. Then, the second peak load attains when concrete reaches its strength due to the larger confinement of outer tube and non-loaded concrete, whilst the capacity of inner tube generally keeps stable at the same time. However, for specimens with large β, the inner tube and concrete beneath the BM reach their strength almost simultaneously. In this case square CFDST specimens have the similar failure process with square CFST specimens, as the inner tube and the concrete close to connection area can undertake the transverse bearing forces together. Generally, square CFDST specimens under local bearing forces also have good structural performance and deformation-resistant ability. It was observed that the form of load ratio (P/P ue ) versus displacement (Δv and Δl ) curves of CFDST under local bearing forces was mainly determined by width ratio (β), as shown in Fig. 10. For members with larger β, there were three stages in the P/P ue versus Δv (Δl ) curve due to the weaker confinement of nonloaded concrete and outer tube to the directly loaded concrete, as well as the support of inner tube, i.e. elastic stage (OA), elastoplastic stage (AB) and stage after the peak load (BC). However, the stage after the peak load of members with smaller β were different

Y.-F. Yang et al. / Thin-Walled Structures 94 (2015) 394–409

T4(0.6)-S90-0.6

T4(0.6)-S90-0.5

Cracks

T4(0.6)-S90-0.6

T4(0.6)-S90-0.4

Crushing

Crushing

Crushing

T3(0.6)-S90-0.6 Crushing

Crushing

Cracks

Cracks

401

Cracks

Cracks

(1)

(2)

T4(0.6)-H90-0.6

T4(0.6)-S90-0.6 Crushing

T4(0.6)-S90-0.6

Crushing

Crushing

Cracks Cracks

T4(0.3)-S90-0.6

T4(0.45)-S90-0.6

Crushing

Cracks

Cracks

Cracks

Cracks

(3)

(4)

T4(0.6)-S45-0.6 Cracks

Crushing

T4(0.6)-S45-0.5

T4(0.6)-S45-0.4

Cracks

Cracks

T4(0.6)-S45-0.6

Crushing

Crushing

T3(0.6)-S45-0.6 Cracks

Cracks

Crushing

Crushing

(1) T4(0.6)-S45-0.6

T4(0)-S90-0.6

Crushing

Crushing

(2)

T4(0.6)-H45-0.6

T4(0.6)-S45-0.6 Cracks

Cracks

Crushing

T4(0.45)-S45-0.6

T4(0.3)-S45-0.6

T4(0)-S45-0.6

Cracks

Cracks

Cracks

Crushing

Crushing

Crushing

(3)

Crushing

(4) Fig. 7. Failure pattern of the sandwiched concrete. (a) θ ¼901. (b) θ ¼ 451.

from those of member with larger β as stronger confinement of non-loaded concrete and outer tube and support of inner tube to the directly loaded concrete were produced. The members with smaller β had a shorter stage after the first peak load (BD), then the load versus displacement curves exhibited the strengthening properties to attain the second peak load (DE), and finally the load decreased more rapidly after the second peak load than the first declining stage. In general, the members with larger β had a higher initial slope of P/P ue versus Δv (Δl ) curve. Fig. 11 shows the typical load ratio (P/P ue ) versus strain ratio (εL =εy and εT =εy ) curves at different positions and the strains at the symmetric locations equal to the average value of all symmetric positions (see Fig. 3), where εL and εT represent the longitudinal and transverse strains, respectively, while εy is the yield strain of outer tube. It can be seen that, in general, the shorter the distance away from the BM, the more sufficient the strain develops. Meanwhile, specimens with θ of 451 have larger strain values than those with θ of 901. Moreover, the strain values corresponding to P ue are generally less than εy except for the longitudinal strain at position 1 and transverse strain at position 7 of specimens with θ of 451, which is

different from square CFST specimens under local bearing forces [17]. This indicates that, when subjected to local bearing forces, the stress of outer tube of CFDST decreases due to the deformations of the corresponding inner tube. The effects of parameters on P  ε=εy curves at position 1 of the tested specimens are shown in Fig. 12, where the solid and dashed lines respectively represent the longitudinal and transverse strains. It can be seen that the development of longitudinal strains is more sufficient than that of transverse strains. In general, the width ratio (β) not only affects the peak load but also determines the strains corresponding to P ue , meanwhile, β has a significant influence on the form of P  ε=εy curve of specimens with θ of 901 and load versus transverse strain curve of specimens with θ of 451. It can also be seen from Fig. 12(b) that, generally, the initial slope of P  ε=εy curves and peak loads increase with decrease of hollow ratio (χ) except for CFST specimen with θ of 901 (i.e. specimen T4 (0)-S90-0.6). This can be explained that, the volume of concrete under compression increases with decreasing χ, resulting in the increasing confinement of outer tube to the sandwiched concrete. However, for CFST specimens with θ of 901, the absence of inner

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1

3

3 1

2 2

1

2 3

1 3

2

Fig. 8. Schematic diagram of force transfer mechanisms. 1-Bearing area; 2-Dispersed bearing area; 3-Transfer path.

tube reduces the constraint to the concrete deformations in length direction, which decreases the confinement of the outer tube to the concrete. 3.3. Bearing capacity In this paper, the maximum load is defined as the bearing capacity (P ue ), as listed in Table 1. Fig. 13 shows the variation of bearing capacity (P ue ) of the tested specimens. It can be seen that, similar to square CFST specimens under local bearing forces [17], square CFDST specimens under local bearing forces also have a high bearing capacity and P ue of specimens with θ of 451 is larger than that of specimens with θ of 901 considered that P ue of specimens with θ of 451 is the resultant result. The type of BM has little influence on P ue of specimens with θ of 901; however, P ue of specimen with θ of 451 loaded by hollow BM is obviously lower than that of specimen loaded by solid BM because the local buckling failure of hollow BM happens before the specimens reaches its ultimate strength. The effecting rules of t o on P ue of CFDST specimens are just opposite to those of CFST specimens found in [17], i.e. P ue increases and decreases with increasing t o for specimens with θ of 901 and 451, respectively. This may be explained by the fact that, the confinement of thicker outer tube to the sandwiched concrete of specimen with θ of 901 is enhanced as the failure of inner tube takes place no later than achieving the ultimate strength; however, the confinement of thicker outer tube to the sandwiched concrete of specimen with θ of 451 is decreased

due to the increasing compressive area of inner tube, which leads to a early failure of inner tube before reaching the ultimate strength. It can also be seen from Fig. 13 that, P ue increases with increasing β and decreasing χ owing to the area increase of directly loaded concrete. 3.4. Deformations corresponding to bearing capacity The displacements corresponding to bearing capacity of the tested specimens (Δv;ue and Δl;ue ) are also presented in Table 1. The variation of Δv;ue and Δl;ue is demonstrated in Fig. 14. It can be seen that, in general, specimens with θ of 451 have a larger resultant Δv;ue and a smaller Δl;ue than specimens with θ of 901. This can be explained that, for specimens with θ of 451, the actual location of outer tube buckling and the maximum lateral displacement are different from the measured position, as shown in Fig. 4(b). The parameters considered in the tests generally have moderate influencing on Δl;ue of the tested specimens whilst the ones loaded by solid BM have a smaller Δv;ue as solid BM has a smaller vertical displacement than hollow BM. It can also be seen from Fig. 14 that, in general, Δv;ue decreases with increase of β and χ, and t o has an opposite effect on Δv;ue of specimens with θ of 451 and 901. This is due to the fact that the constraint of steel tubes and non-loaded concrete to the loaded concrete decreases with increase of β and χ, which leads to a smaller Δv;ue . Moreover, for specimens with θ of 451, the increasing to may result in the early failure of the sandwiched concrete and the inner tube due to their complex loading conditions, as a result, Δv;ue decreases with increasing to.

Y.-F. Yang et al. / Thin-Walled Structures 94 (2015) 394–409

1000

403

2000 T4(0.6)-S90-0.6 T4(0.6)-S90-0.5 T4(0.6)-S90-0.4

1500 P (kN)

P (kN)

750 500 250

1000 500

T4(0.6)-S45-0.6 T4(0.6)-S45-0.5 T4(0.6)-S45-0.4

0 -15

0

4000

30

10

20

T4(0.6)-S45-0.6 T4(0.45)-S45-0.6 T4(0.3)-S45-0.6

P (kN)

4800

2000 1000 0 -15

0

6400

T4(0.6)-S90-0.6 T4(0.45)-S90-0.6 T4(0.3)-S90-0.6 T4(0)-S90-0.6

3000 P (kN)

15

0 -10

T4(0)-S45-0.6

3200 1600

0

15

30

0 -10

0

10

20

Fig. 9. Load versus displacement curves.

Fig. 10. Typical load versus displacement relationship.

After completing the tests, the protrusion of concrete (δc ) at the ends of the tested specimens was measured by a caliper. The variation of δc is demonstrated in Fig. 15, where δc is the average value of twenty measuring results. It can be seen that, except for one specimen with θ of 451 loaded by hollow BM (i.e. specimen T4(0.6)H45-0.6) fails due to the local buckling of hollow BM, δc of specimens with θ of 451 is larger than that of specimens with θ of 901, this is because the concrete protrusion of specimens with θ of 451 can only develop towards the top end with the restricted bottom end (see Fig. 2), which is similar to the findings of square CFST specimens under local bearing forces [17]. Specimens loaded by hollow BM have a smaller δc than specimens loaded by solid BM as the top flange of outer tube and concrete within the hollow BM may deform upward. It can also be seen that, δc increases with increasing β and t o and decreasing χ. This is due to the fact that, for specimens with a

larger β and a smaller χ, a larger volume of concrete beneath the BM will be forced to protrude along the length direction. Moreover, the thicker the outer tube the larger the δc , because for specimen with a larger t o , a stronger confinement to concrete and a smaller friction force between steel tubes and the sandwiched concrete were produced under the same bo . The variation of strain ratios (εue =εy ) of the tested specimens is shown in Fig. 16, where the solid and dashed lines represent the longitudinal and transverse strains, respectively, and εue is the strain corresponding to P ue . It can be seen that, the variation range of εue =εy of specimens with θ of 451 is larger than that of specimens with θ of 901. This indicates that the composite action between steel tubes and the sandwiched concrete of specimens with θ of 451 under the resultant vertical forces is better than that of specimens with θ of 901 under the component vertical forces. For specimens with θ of 901, the largest longitudinal and transverse εue generally appear at position 1 and position 5, respectively, whilst for specimens with θ of 451 the largest longitudinal and transverse εue generally appear at position 1 (or 4) and position 7, respectively. This further indicates that, a shorter distance away from the BM produces a more sufficient strain development. It can also be found that, generally, the type of BM and t o have a moderate effect on εue =εy except for specimen T4(0.6)S45-0.6. In general, εue =εy of the specimens increases with decreasing β and χ as the confinement of outer tube and non-loaded concrete to the directly loaded concrete increases.

4. Prediction of the bearing capacity The design method for circular tubular joints with chord ovalization failure has been suggested in [22], in which the bearing capacity can be calculated by substituting the effective thickness of tube into the design equations for tubular joints with circular hollow section

Y.-F. Yang et al. / Thin-Walled Structures 94 (2015) 394–409

1.2

1

1

0.8

0.8 P /Pue

1.2

0.6 Point 1 Point 2 Point 3 Point 4 Point 5 Point 6

0.4 0.2 0

-3

-2

-1

0

1

1.2

1.2

1

1

0.4 0.2 0

-8

-4

-2

-1

0

1

0.8

Point 1 Point 2 Point 3 Point 4 Point 5 Point 6 Point 7 Point 8 Point 9 Point 10

0.6

Point 1 Point 2 Point 3 Point 4 Point 5 Point 6

0.2 0

-4

0.8 P /Pue

0.6 0.4

P /Pue

P /Pue

404

2

3

Point 1 Point 2 Point 3

0.6

Point 4 Point 5

0.4

Point 6 Point 7 Point 8

0.2

Point 9 Point 10

0

4

0 -1

0

1

2

3

4

Fig. 11. Typical load ratio (P/Pue) versus strain ratio curves. (a) Specimen T4(0.6)-S90-0.6. (b) Specimen T4(0.6)-S45-0.6.

Fig. 12. P  ε=εy curves at position 1.

2000

2000

1500

1500 P ue (kN)

P ue (kN)

Y.-F. Yang et al. / Thin-Walled Structures 94 (2015) 394–409

1000 500

1000 500

0.4

0.5

0.6

0

0.7

2000

6000

1500

4500 P ue (kN)

P ue (kN)

0 0.3

405

1000 500

2

3

4

5

0

0.2

0.4

0.6

3000 1500

0 Hollow BM

Solid BM

0

Fig. 13. Variation of bearing capacity of the tested specimens.

Fig. 14. Variation of Δv;ue and Δl;ue .

406

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Fig. 15. Protrusion of concrete.

17:25 MPa r f c r110 MPa

ð3Þ

member: 8 < 2f yoðiÞ U t oðiÞ Uðbb þ 5r ext Þ Uαc pffiffiffiffiffiffiffiffiffiffiffi P soðiÞ ¼ : f yoðiÞ U t 2oðiÞ U ð2C þ 4 1 C Þ=ð1  CÞ

bo r 80 to

ð4Þ

αc ¼

bi r 40 ti

ð5Þ

(CHS). At the same time, the following limitations should be observed when using the aforementioned method: 0

(

7r

2ðbo  2t o Þ r 45 ðbo  2t o  bi Þ

ð6Þ

0

where, f c is the cylinder compressive strength of concrete. Given that the inner tube geometry of ten CFDST specimens and the concrete annulus geometry of all CFDST specimens exceed the limitation in Eqs. (5) and (6), the existed method cannot be applied directly to predicting the bearing capacity of square CFDST under local bearing forces in present paper. As can be found from Fig. 7 that, under local bearing forces, outer tube, the sandwiched concrete and inner tube can generally work well together, and there is no apparent separation between them. Therefore, similar to the method for dealing with the bearing capacity of CFDST under axial compression [1], the bearing capacity of CFDST under local bearing forces (P u ) can be determined based on the superposition method: Pu ¼

P so þP c þ P si sin θ

ð7Þ

where, P so and P si are the bearing capacity of outer and inner tube respectively, which can be temporarily determined using the method in [21] considering their similar failure pattern as SHS under transverse bearing forces transmitted from the welded web

ðβoðiÞ Z 0:8Þ ðβoðiÞ o 0:8Þ

0:7

ðβoðiÞ ¼ 1:0Þ

0:529  0:0054ðboðiÞ  2r ext Þ=t oðiÞ

ð0:8 r βoðiÞ r0:9Þ

ð8Þ

ð9Þ

in which, the subscripts o and i respectively represent outer and inner steel tube, f yoðiÞ is the yield strength of flat part of outer (inner) steel tube, r ext is the external corner radius, αc is the stability factor and the linear interpolation is needed to obtain αc when βoðiÞ is greater than 0.9 and less than 1.0, βoðiÞ is the width ratio between BM and outer (inner) tube, and C is the intermediate variable and equals to ðbb þ2t oðiÞ Þ=boðiÞ . The bearing capacity of the sandwiched concrete (P c ) can be obtained based on the formula for core concrete in CFST under local bearing forces [19] with consideration of effect of hollow ratio (χ):  0:5 A1 P c ¼ kh Uf c' UA0 U ð10Þ A0 kh ¼ 1:0  0:1χ  1:6χ 2

ð11Þ

in which, A0 is the bearing area over which the transverse load is applied, A1 is the dispersed bearing area, and kh is a factor considering the effect of hollow ratio (χ). Eq. (11) is obtained by regression of the experimental data. In this paper, the method in [19] is adopted to calculate the dispersed bearing area of the sandwiched concrete (A1 ), i.e. the bearing stress is assumed to disperse in the longitudinal and lateral directions at a slope of 1:1 until the edge of concrete or the top surface of inner tube is reached, as shown in Fig. 17. The calculated bearing capacities (P uc ) were compared with the results obtained from the experimental investigation (P ue ) except for

Y.-F. Yang et al. / Thin-Walled Structures 94 (2015) 394–409

407

Fig. 16. Variation of strain ratios (εue =εy ). (a) θ ¼ 90o . (b) θ ¼ 45o .

specimens loaded by hollow BM, as shown in Fig. 18, where μ and σ represent the mean value and the corresponding standard deviation of P uc =P ue , respectively. It is shown from the comparison that the calculated bearing capacities are generally lower than the experimental results and

the data of P uc =P ue are relatively discrete. This means that finite element analysis is needed to evaluate the mechanism of square CFDST under local bearing forces and further to propose the rational design method based on the parametric analysis.

408

Y.-F. Yang et al. / Thin-Walled Structures 94 (2015) 394–409

bo bb

bb

BM Outer tube

A1

A1

Outer tube

1

1

bi

Inner tube

bo

1

bi

1

BM

Inner tube Concrete Concrete Fig. 17. Calculation diagram of the sandwiched concrete. (a) Side section view. (b) Axial section view.

more sufficient near the connection area of specimen and BM, whilst specimens with θ of 451 have more sufficient strain developments than the ones with θ of 901. (4) Within the scope of this research, a proposed design method is used to predict the bearing capacity of square CFDST under local bearing forces, and the calculated bearing capacities are generally in good agreement with the experimental results and tend to be safe. The suggested formulae for the bearing capacity of CFDST subjected to local bearing forces could be referred to during the structural design of the relevant issues.

4000

P uc (kN)

3000 2000 1000 0

0

1000

2000 3000 P ue (kN)

4000

Fig. 18. Comparison between the calculated and experimental bearing capacities.

5. Conclusions The experimental behaviour of square concrete filled doubleskin steel tube (CFDST) under local bearing forces has been presented. The following conclusions can be made from the observations and analytical results obtained in this study: (1) In general, square CFDST specimens under local bearing forces have a high bearing capacity and a good deformation-resistant ability. The failure patterns of outer tube and the sandwiched concrete of CFDST specimens are similar to those of concrete filled steel tube (CFST) specimens except for one CFDST specimen with relative small width ratio between bearing member (BM) and compression member (β). The failure pattern of inner tube is similar to that of steel square hollow section (SHS) under transverse bearing forces. (2) For all specimens, vertical displacement and longitudinal strain are larger than lateral displacement and transverse strain, respectively, whilst load versus deformation curves are mainly affected by β and hollow ratio (χ). Meanwhile, the experimental bearing capacity (P ue ) increases with increasing β and decreasing χ, whilst type of BM and thickness of outer tube (t o ) have moderate effects on P ue , without considering the local buckling of hollow BM. (3) Generally, the parameters considered in this tests have moderate effects on lateral displacement corresponding to P ue while specimens with larger β and χ have smaller vertical displacement corresponding to P ue . Protrusion of concrete (δc ) of specimens with included angle between BM and compression member (θ) of 451 are larger than that of specimens with θ of 901, whilst δc inncreases with increasing β and t o , as well as decreasing χ. Moreover, strain of steel tube tends to develop

For better understanding, the experimental observations on the buckling mode of tubes and protrusion of the sandwiched concrete should be reproduced by a finite element analysis (FEA) study, through which the load-transfer mechanism should be investigated and the range of parameters should be extended. A design approach for truss with square CFDST chords, which takes into account the properties of CFDST under local bearing forces, should be further provided.

Acknowledgements The authors would like to acknowledge the Science Fund for Creative Research Groups of the National Natural Science Foundation of China (51421064), Tsinghua Initiative Scientific Research Program (No. 20131089347) and the Natural Science Foundation of Liaoning Province (2013020125), which collectively funded this project. The authors are also grateful to Mr. Lei Zhang for his assistance in the tests. References [1] Zhao XL, Han LH. Double skin composite construction. Progress in Structural Engineering and Materials 2006;8(3):93–102. [2] Han LH, Huang H, Tao Z, Zhao XL. Concrete-filled double skin steel tubular (CFDST) beam-columns subjected to cyclic bending. Engineering Structures 2006;28(12):1698–714. [3] Han LH, Li W, Bjorhovde R. Developments and advanced applications of concrete-filled steel tubular (CFST) structures: Members. Journal of Constructional Steel Research 2014;100:211–28. [4] Yang YF, Han LH, Sun BH. Experimental behaviour of partially loaded concrete filled double-skin steel tube (CFDST) sections. Journal of Constructional Steel Research 2012;71:63–73. [5] Li W, Ren QX, Han LH, Zhao XL. Behaviour of tapered concrete-filled double skin steel tubular (CFDST) stub columns. Thin-Walled Structures 2012;57: 37–48. [6] Li W, Han LH, Chan TM. Tensile behaviour of concrete-filled double-skin steel tubular members. Journal of Constructional Steel Research 2014;99:35–46. [7] Hassanein MF, Kharoob OF, Liang QQ. Circular concrete-filled double skin tubular short columns with external stainless steel tubes under axial compression. Thin-Walled Structures 2013;73:252–63.

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