Compressive resistance of steel-concrete-steel sandwich composite walls with J-hook connectors

Journal of Constructional Steel Research 124 (2016) 142–162

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Journal of Constructional Steel Research

Compressive resistance of steel-concrete-steel sandwich composite walls with J-hook connectors Zhenyu Huang a,⁎⁎, J.Y. Richard Liew a,b,⁎ a b

Department of Civil and Environmental Engineering, National University of Singapore, E1A-07-03, 1 Engineering Drive 2, Singapore 117576 College of Civil Engineering, Nanjing Tech University, Nanjing, Jiangsu 211816, China

a r t i c l e

i n f o

Article history: Received 20 July 2015 Received in revised form 4 May 2016 Accepted 4 May 2016 Available online xxxx Keywords: composite wall J-hook connector sandwich wall steel-concrete-steel ultra lightweight cement composite

a b s t r a c t This paper investigates the structural behaviour of Steel-Concrete-Steel (SCS) sandwich wall which consists of two external steel plates infilled with ultra-lightweight cementitious composite material. A series of compression tests consists of a wide range of parameters have been carried out on the SCS sandwich walls of different heights forming short and slender wall. The test results show that the SCS sandwich walls with J-hook connectors exhibit comparable behaviour in compressive resistance and post-peak unloading behaviour to the ones with the overlapped headed studs. The interlocking J-hook connectors play an important role in providing composite action between the steel plates and the cementitious core, and preventing or delaying the local buckling of the external steel plates. The test results are compared against the predictions by Eurocode 4 and AISC 360 methods for composite columns. It is found that the Eurocode 4 and AISC 360 methods could over-predict the compressive resistance of sandwich wall subjected to compression. A modified method is then proposed, which takes into account the effect of interlocking J-hook connectors in providing lateral restraints to the external steel plates. The predictions show a reasonable correlation with the test results. Nonlinear finite element model has been established to predict the load-displacement curves, maximum resistance and failure modes of the sandwich walls. Both the experimental and finite element results confirm that the proposed analytical formulae are conservative for design of SCS sandwich composite walls with J-hook connectors. © 2016 Elsevier Ltd. All rights reserved.

1. Introduction Steel-Concrete-Steel (SCS) sandwich is a special form of structural unit comprising a central concrete core sandwiched by two steel face plates whose behaviour is greatly influenced by the interface bonding between the two different materials. An effective way to achieve composite action is to use mechanical shear connectors to lock the concrete core and the steel plates together. For example, Bi-Steel connector [1,2], bi-directional corrugated-strip-core system [3], binding bars and tie bars [4,5] were proposed by former researchers, as shown in Fig. 1(a)(f) to achieve composite action between the concrete core and the face plates. Novel shear connectors such as the cable and U-shaped connectors, U connector-Steel bar-U connector, cable and U-shaped connectors have been proposed by Sohel et al. [6]. These connectors have their own merits in term of ease of installation and ability to withstand extreme loads without loss of structural integrity. ⁎ Correspondence to: J.Y.R. Liew, Department of Civil and Environmental Engineering, National University of Singapore, E1A-07-03, 1 Engineering Drive 2, Singapore 117576. ⁎⁎ Corresponding author. E-mail address: [email protected] (J.Y.R. Liew).

http://dx.doi.org/10.1016/j.jcsr.2016.05.001 0143-974X/© 2016 Elsevier Ltd. All rights reserved.

The SCS sandwich composite plates exhibit significant structural and economic advantages over the conventional reinforced concrete (RC) structures in terms of higher flexural stiffness and higher resistance to withstand extreme environmental and accidental loads. The external steel plates serve as a permanent formwork during concreting, promoting construction efficiency and reducing the site handling costs and time. The waterproof feature inherently provided by external steel plates reduces the surface area that needs corrosive coating and makes it easy for inspection and maintenance. Pre-fabrication of large panels in the factory and rapid installation on-site translates into further time and cost savings. As a result, SCS sandwich composite structures can be adopted as heavy duty and protective layers such as ship hulls, tunnels, military shelters and nuclear power plant walls that require resistance against impact and blast [7–9]. More recently, SCS sandwich structures are being considered for the outer shell of Arctic offshore structures as ice resistant wall to resist impact loads due to ice floes [8]. Fig. 2(a) shows the offshore platform structure using SCS sandwich structure while Fig. 2(b) shows the schematic view of the internal structure. In the case of an offshore oil platform, the SCS sandwich component is subjected to transverse as well as axial compressive load.

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Nomenclature Ac As Ase beff d Es Ec EIeff fy fu fut fck fst H hc k Ncr,Pe Npl,Rk NEC4 NAISC NuI,NuII Ntest NFE nJ PJ R s ts b νc,νs ρw λ δ0.85 δu δ μ ϕc,ϕ2 ϕa,ϕ1 χ σc,σf σcr σh

cross-sectional area of concrete core cross-sectional area of steel face plate cross-sectional area of shear connector effective width of wall section diameter of connector modulus of elasticity of structural steel secant modulus of elasticity of concrete effective stiffness yield strength of steel tensile strength of steel tensile strength of hook bar cylinder compressive strength of concrete splitting tensile strength of concrete wall height concrete thickness plate buckling coefficient elastic critical buckling force characteristic plastic resistance of composite section to compressive normal force prediction load by Eurocode 4 (encased section) prediction load by AISC 360-10 prediction loads by the proposed models failure load by test failure load by finite element analysis number of J-hook shear connectors shear resistance of a J-hook connector within a SCS composite structure width-thickness ratio parameter connector spacing thickness of steel face plate width of sandwich wall section Poisson ratio of concrete and structural steel density of ULCC non-dimensional slenderness axial shortening corresponding to 85% ultimate resistance axial shortening δu at peak load axial shortening of wall displacement ductility factor reduction factor applied to concrete strength reduction factor for steel strength resistance reduction factor for slender wall confinement stress and failure stress of concrete elastic critical plate buckling stress transverse tensile stress of steel face plate

Extensive research work has been experimentally based on steelconcrete composite structures and the majority of the work focuses on verifying the effectiveness of theoretical calculations. Past researches and application of steel-concrete composite walls in nuclear power plant introduce angle connectors and headed shear studs welded on the steel plate. However, the compressive tests show that premature local buckling may occur before steel plate yielding [10,11]. Wright et al. [12–14] studied the axial and bending behaviour of composite walls using profiled steel sheet and it was found that the longitudinal bending stiffness of the composite walls offered substantial benefits in maintaining stability and preventing the early global buckling. However, premature local buckling of profiled sheeting plate occurred during the test. Since then, Bradford et al. [15] and Prabha et al. [16] investigated the structural performance of double skin profiled steel sheeting infilled with normal and lightweight foamed concrete under axial compression. The profiled steel facings were connected by through-through

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bolts. Mydin and Wang [17] conducted compression tests on short profiled thin-walled steel sheeting infilled with lightweight foamed concrete to evaluate the plate local buckling coefficients proposed by Uy and Bradford [18]. Calibration showed that with the Liang and Uy [19] effective width formulation, the proposed strength calculation model gave a best agreement with the experimental tests. However, the compressive strength of the foamed concrete was very low (around 6~8 MPa) which may not be applicable for offshore structures that requires high compressive resistance and impact resistance. It is found that steel profile sheets could provide some confinement to the concrete with minor enhancement of compressive strength but the failure behaviour is brittle due to inadequate bond. In order to enhance the structural performance of SCS sandwich structure, a novel interlocking J-hook shear connector [20,21] was proposed as shown in Fig. 1(e). The J-hook connector not only provides composite action between steel face plates and concrete core but also prevents tensile separation between them when the sandwich composite panels are subjected to extreme loads [22]. One critical factor affecting the structural performance of sandwich composite wall is the imperfect bond between the steel plates and the concrete core. The imperfect bonding will decrease the confining pressure provided to the concrete core, and reduce the elastic resistance and stiffness. The imperfect bonding may also increase the chance of plate buckling occurrence. This problem has drawn the attention of many researchers in recent years. Two popular approaches have been proposed to enhance the steel-concrete interface bonding in composite structures. The first is to adopt expansive concrete [23], which compensates the shrinkage and temperature variation effect. However, the steel plate still dilates more than the concrete core in compression, and thus it reduces the confining pressure on the expansive concrete. The second method is to change the cross-sectional shape or adding additional lateral restraint, which includes internal tension stiffeners, binding bars, external rings and tie bars [1–5]. However, these internal stiffeners have limitations in term of flexibility of fabrication especially when the gap between the two face plates is too small. The installation of binding bars and external rings requires welding or bolting on the plates, the quality of such welds which is difficult to control. It requires drilling of holes on the steel plates which may initiate stress concentration on the plates [24–26]. The application of ultra-lightweight concrete on nowadays practical construction is rare. One of the major reasons is owing to its brittleness and low strength as compared to normal lightweight concrete [15–17]. One recent breakthroughs in steel-concrete-steel sandwich technology in offshore structural application is the development of ultralightweight cement composite (ULCC) [7,27] to be used as infilled materials. This ULCC material can achieve a high compressive strength up to 60 MPa with density less than 1400 kg/m3. As ULCC is a relatively new material, the constitutive laws for normal strength concrete may not be directly applicable. Therefore, the validity of the current analysis and design methods for steel-concrete composite structures should be checked before applying them to the new ULCC materials in SCS sandwich composite construction. This paper develops a new Steel-Concrete-Steel (SCS) sandwich wall which consists of two external steel plates infilled with ultralightweight cementitious composite material. The two external steel plates are connected together by inter-locking J-hook connectors to overcome the abovementioned disadvantages. The J-hook bar is made of bending steel bar and welded on the external plates which are then matched together by the interlocking hooks as shown in Fig. 3. The lightweight cement composite is then cast in between the two steel plates. The hook length can be adjusted depending on the thickness of the sandwich core. The internal J-hook connectors offer an interlocking effect on the external plates preventing them from moving outward. The J-hook connectors provide the necessary shear resistance and also prevent tensile separation between the steel face plates and the concrete core when the SCS sandwich wall is subjected to compression. The earlier research focused on sandwich beams and slabs with J-hook

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Fig. 1. Patterns of stiffeners used in double skin composite structures (a) with Bi-steel [1,2]; (b) with bi-directional corrugated-strip-core [3]; (c) with overlapping headed stud; (d) with binding bar [4]; (e) with J-hook [6]; (f) with tie bar [5].

connectors subject to transverse loads [9,20–22]. This paper extends the previous work to investigate the structural performance of such SCS sandwich walls subjected to compression. 2. Experimental investigation 2.1. Material properties 2.1.1. Ultra lightweight cementitious composite (ULCC) Three grades of ultra-lightweight cementitious composite material (ULCC): C35, C45 and C60, are used in the preparation of the test specimens. The ULCC is made of Ordinary Portland Cement, lightweight cenosphere and silica fume as the supplementary cementitious material. The water-to-blinder remains low with application of superplasticiser. To reduce early shrinkage, air content and drying effect, shrinkage reducing admixture and 0.5% polyvinyl alcohol (PVA) fibres are added. Table 1 presents the material properties of PVA fibres while Table 2 shows the basic mixture proportions for each grade of ULCC. To improve the workability, superplasticiser and shrinkage reducing admixture are controlled based on trial mixes to achieve a desired flow of 210~240 mm, tested in accordance with BS EN 1015-3:1999 [28]. Such high flow is required so that the cement composite can fill the voids within the thin gap of the sandwich specimens and flow around the closely spaced shear connectors. The ULCC has a 28-day compressive strength of about 60 MPa with an average density of 1360 kg/m3, which is about 60% unit weight of the normal weight concrete of 2400 kg/m3. Besides the weight saving, the absence of the coarse aggregates leads to a highly workable material

suitable for pumping in construction. Table 3 shows the mechanical properties for each grade of ULCC. 2.1.2. Steel plate and J-hook connectors S275 mild steel plate is used to fabricate the SCS sandwich wall specimens. The new form SCS sandwich composite walls are strengthened by J-hook connectors to overcome the abovementioned disadvantages of conventional internal stiffeners for composite columns. The J-hook bar is made by cold bending of HPB235 steel bar and welded on the steel plates. The Young's modulus Es, Poisson ratio νs, yield strength fy and ultimate strength fuof the steel plates and hook bars are obtained from tensile testing of steel samples. The standard tensile test for the steel materials follows the procedure outlined in ASTM E-8M [29]. The detailed measured properties of steel plates and hook bars are listed in Table 4. 2.2. Test specimens A total of fifteen sandwich wall specimens with different parameters are prepared and tested as shown in Table 5. The test parameters are the steel contribution ratio (Asfy/Acfc), connector spacing (s), concrete compressive strength (fck), concrete core thickness (hc) and overall slenderness (λ). In Table 5, SCSW-R1 and SCSW-R2 are designed as the referential specimens. SCSW-H1 and SCSW-H2 are designed as slender walls with slenderness ratioλ = 0.54 and 0.98 by varying the height of the specimen (H). For SCSW-S1 and SCSW-S2, sparser connector arrangement is used to consider the effect of connector spacing (partial composite design). SCSW-C1 and SCSW-C2 are designed with thinner

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Fig. 2. Offshore platform structure using SCS sandwich structure.

and thicker concrete core respectively. Headed shear stud connector is used for SCSW-HS as compared to the referential specimens with Jhook connector. Specimens SCSW-C30, SCSW-C45, SCSW-T8 and SCSW-T12 are prepared to investigate the influence of concrete compressive strength and steel contribution factor. To investigate the confinement effect, closed edge and open edge plates are added in SCSWP1 and SCSW-P2 respectively, as shown in Fig. 4. However, the side plates do not contribute to load capacity but only serve as confined plates in the specimens. 2.2.1. Dimensions and connectors design Except SCSW-H1 and SCSW-H2, the rest of the sandwich wall specimens are designed to behave as short walls to avoid global buckling

with an overall height of 400 mm, as shown in Table 5 and Fig. 5(a) and (b). Two 15 mm thickness cover plates are welded on both ends of the specimens so that compression load can be applied uniformly on the cross section. 50x20x8mm plate stiffeners are also provided to strengthen both ends of the wall to ensure that the failure would occur within the instrumented region, and to prevent the premature end failure. Fig. 5(c) shows the fabricated sandwich walls with J-hook connectors and headed studs. Previous research of SCS sandwich structures faced problems with local buckling, e.g. Bi-Steel [1]. However, J-hooks can be placed at much closer spacing than Bi-Steel connectors and Sohel et al. [6] derived a formula relating maximum J-hook connector spacing and steel plate thickness. For S275 steel plate, width-to-thickness ratio of s/ts ≤ 52is needed

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Fig. 3. Fabrication procedures of SCS sandwich wall with J-hook connectors.

to prevent local buckling. A shear connector spacing of 100 mm has been selected conservatively for full composite design, as shown in Fig. 5(a) and (b). Moreover, Eurocode 4 [30] section 6.6.5.7(5) limits the stud diameter to less than 2.5 times the thickness of the part it is welded i.e., d b 2.5ts. If a shear stud with diameter d = 13 is selected, its diameterto-thickness ratio is equal to d/ts = 2.17.

2.2.2. Casting and curing Each specimen is cast individually in small quantities of ULCC to ensure better control of workability and workmanship. ULCC is poured sideways and the specimens are naturally cured in the laboratory. The sandwich walls are tested at the 28th day after curing. Concrete cylinders are also cast on the day the sandwich walls are made. The corresponding compressive strength is obtained by testing ∅100x200 mm size cylinders according to ASTM C39/39M-09 [31].

2.3. Test set-up, loading procedure Fig. 6 shows the test set-up for compression tests. A 10 MN Instron testing actuator operated in displacement-control mode is used to test the specimens. Two modified solid hinged supports are designed and one of which is set on the rigid beam base while the other is bolted to the actuator. The supports include a cylinder that simulates a line load on the specimens and which also allows rotation. The composite wall specimens are bolted onto the support plates. Hence, the boundary conditions of the wall are pin-pin end supported. A quasi-static loading procedure is introduced in four steps for the tests: (1) preload at a rate of 0.2 mm/min for 400 mm long specimen while 0.3 mm/min for 2500 mm long specimen respectively up to 10% of calculated resistance by Eurocode 4; (2) unload at a rate of 0.5 mm/min for all the specimens; (3) reload at the same rate as in Step 1 until the peak load is reached; (4) finally in the post-peak range, increase the rate to 0.5 mm/min

Table 1 Properties of PVA fibres. Fibre types

Tensile strength (MPa)

Elastic modulus (GPa)

Length (mm)

Diameter (μm)

Aspect ratio

Density (kg/m3)

Polyvinyl Alcohol (PVA)

2610

79

12

39

308

0.97

Table 2 Mixture proportion of ULCC. Mixture proportion of matrix by mass of total binder Mix ID

ULCC-60 ULCC-45 ULCC-30

PVA Fibre (vol%)

0.5 0.5 0.5

*SRA = Shrinkage reduced admixture.

Water/binder

0.32 0.45 0.51

Binder Cement

Silica fume

0.92 0.89 0.87

0.08 0.11 0.13

Cenosphere/binder

SRA/binder

0.42 0.59 0.66

0.025 0.035 0.039

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Table 3 Measured material properties of ULCC. Item

Grade

ρw (kg/m3)

fck (MPa)

fst (MPa)

Ec (GPa)

νc

Mean

C60 C45 C30

1361 1298 1283

60.9 54.4 37.8

5.40 2.83 2.84

15.4 15.4 15.4

0.25 0.25 0.25

*ρw = 1 day density; fck = cylinder compressive strength; fst = splitting tensile strength; Ec = Young's Modulus; νc = Poisson ratio.

Table 4 Mechanical properties of steel plate and J-hook bar. Item

Thickness/diameter

Es (MPa)

fy (MPa)

fu (MPa)

νs

Mild steel plate

6 mm 8 mm 12 mm HPB 235φ13

202.0 179.6 211.5 189.9

309.4 393.9 374.6 356.8

426.8 522.9 507.0 521.7

0.3

J-hook bar

νs = Poisson ratio of steel (assumed value);

until significant visible deformation is observed. During each loading step, concrete cracks or buckling of steel plates are recorded. Fig. 4. Different types of sandwich sections and loading area.

2.4. Instrumentations Each specimen is instrumented with post-yield steel strain gauges (YFLA-5) and concrete strain gauges (PFL-30-11) at the side surface to capture the strain distribution along the steel plates and to assess the compression strain distribution across the concrete core respectively, as displayed in Fig. 7. Linear variable displacement transducers (LVDTs) are used to capture the axial shortening and rotation or to measure bulging or, if applicable, buckling of outer face plates. 3. Test results and discussions 3.1. Failure modes and ultimate resistance Fig. 8 shows the failure modes of the tested specimens. There are four types of failure modes observed from the test. The first type of failure is concrete splitting and crushing followed by subsequent local buckling of steel plates, as shown in Fig. 8(a). This type of failure initiates combined crushing and splitting cracks of concrete, with the outer steel plate yielded. After that, the steel plates tend to buckle outward since the sectional stiffness degrades due to the cracking of the

concrete core. Specimens SCSW-R1, SCSW-R2, SCSW-H1, SCSW-C1, SCSW-HS, SCSW-C30, SCSW-C50 and SCSW-C2 all fail in this type of mode. The ultimate load resistance of these specimens can achieve on average 88% the cross-sectional resistance calculated by Eurocode 4 approach for encased composite section. The second type of failure observed from the test is interfacial bond failure with severe abrupt buckling of outer steel plates, as shown in Fig. 8(b). This type of failure is due to inadequate shear bond between concrete and steel face plates because of insufficient shear connector design (partial composite). Excessive slip occurred between the steelconcrete interfaces. Specimens SCSW-S1 and SCSW-S2 fail in this type of mode. It is found that SCSW-S1 and SCSW-S2 exhibit brittle behaviour with loud popping sound heard at failure rather than ductile post-peak unloading behaviour. The ultimate load resistances of these two specimens can achieve on average only 75% cross-sectional resistance calculated by Eurocode 4 approach for encased composite section. The third type of failure mode is cross-sectional failure as shown in Fig. 8(c). Specimens SCSW-T8 and SCSW-T12 fail in this type of mode. From the figures, the concrete core are fully crushed and the steel face plates are fully yielded. After failure, the shanks of the J-hook connectors

Table 5 Dimension and material properties of sandwich wall specimens. Specimen size Specimen SCSW-R1 SCSW-R2 SCSW-H1 SCSW-H2 SCSW-S1 SCSW-S2 SCSW-C1 SCSW-HS SCSW-C30 SCSW-C45 SCSW-T8 SCSW-T12 SCSW-P1 SCSW-P2 SCSW-C2

H(mm)

b (mm)

hc (mm)

ts (mm)

400 400 1000 2500 500 700 400 400 400 400 400 400 400 400 400

590 590 590 590 590 590 590 590 590 590 590 590 590 590 590

120 120 120 120 120 120 90 120 120 120 120 120 120 120 135

6.0 6.0 5.0 6.0 6.0 6.0 6.0 6.0 6.0 6.0 8.0 12.0 6.0 6.0 6.0

s(mm)

connector

b/t

s/ts

λ

fck(MPa)

fy(MPa)

100 100 100 100 201 300 100 100 100 100 100 102 103 100 100

JH JH JH JH JH JH JH HSS JH JH JH JH JH JH JH

4.47 4.47 4.54 4.47 4.44 4.44 5.78 4.47 4.47 4.47 4.34 4.10 4.47 4.47 4.01

16.67 16.67 20.00 16.67 33.50 50.00 16.67 16.67 16.67 16.67 12.50 8.50 17.17 16.67 16.67

0.29 0.29 0.51 0.92 0.32 0.38 0.37 0.30 0.27 0.29 0.30 0.25 0.30 0.29 0.26

54.3 54.3 60.9 58.9 54.6 55.0 60.0 61.1 37.8 54.4 57.1 52.0 55.3 53.8 52.6

309.4 309.4 309.4 309.4 309.4 309.4 309.4 309.4 309.4 309.4 393.9 375.0 309.4 309.4 309.4

*H = height of specimen; hc = thickness of concrete core; ts = thickness of steel plate ; t = 2ts + hc; b = width of sandwich wall section; s = spacing of connector; s/ts = plate slenderness pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ratio; λ = member slenderness ratio, λ ¼ N pl;Rk =N cr ; fck = compressive strength of concrete cylinder ∅100×200; fy = yield strength of steel; HSS = headed shear stud; JH = J-hook.

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Fig. 5. Preparation of SCS sandwich wall specimens.

are necked and finally fail in fracture. Local buckling of steel face plates is delayed due to the decrease of s/ts ratio. The ultimate load resistances of these two specimens can achieve on average 99% cross-sectional resistance calculated by Eurocode 4 approach for encased composite section. The fourth type of failure mode is global buckling of slender wall specimen, as shown in Fig. 8(d). Specimen SCSW-H2 fails in this type of mode. Concrete crushing at concave and steel yielding at convex are

observed at the mid-length region due to global bending. The ultimate load resistance of this specimens can achieve 52% cross-sectional resistance only because of the slenderness effect. 3.2. Strength index (SI) To eliminate the differences of sectional dimensions and plate thickness among specimens, strength index (SI) is introduced, which can be

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3.3. Ductility index (DI) The structural ductility index represents the ability to undergo large plastic deformation without significant strength degradation. The displacement ductility factor μ is introduced which is defined as the ratio of axial shortening corresponding to 85% ultimate resistance during descending stage δ0.85to the axial shortening δu at peak load of the axial load-shortening curves, as shown in Eq.(2): μ¼

δ0:85 δu

ð2Þ

Table 6 also summarises the ductility index for the SCS sandwich wall. The average value of ductility indexμ is 1.204 with variance of 0.174. It is found that the specimens fail in interfacial bond have lowest ductility index, e.g., SCSW-S1 and SCSW-S2. The specimens with small plate slenderness ratio s/ts (SCSW-T8 and SCSW-T12) have higher ductility index because these two specimens fail in cross-sectional failure. The decrease of plate slenderness ratio will increase the ductility index. Moreover, the ductility index increases slightly in the specimens SCSW-P1 and SCSW-P2 since the local buckling is delayed by using side plates.

Fig. 6. Specimen, supports and test rig.

3.4. Effect on load-axial shortening behaviour defined as:

SI ¼

Ntest 0:85Ac f ck þ As f y

ð1Þ

where Ntestis the peak value of axial load-shortening curve. Based on the Eq. (1), Table 6 summarises strength index SI of all the specimens. The average value of SI is 0.849 with variance of 0.121, which indicates the cross-sectional resistance of the specimens are not fully utilised. This is attributed to the different failure modes of lightweight cement composite used in the sandwich walls. When the cross-sectional aspect ratio (width to thickness ratio) becomes larger, the average compressive strength of the concrete core may not achieve its cylinder strength fck due to: (1) the boundary constraint effect of the sandwich wall is different from that of cylinder sample under compression; (2) initial imperfections (e.g., initial air voids, concrete compaction and aggregate distribution, etc.) are more pronounced which may lead to lower concrete strength; and (3) stress concentration at the interface between J-hook and its surrounding concrete which could initiate local failure of concrete core under compression. These reasons may reduce the average compressive resistance of the sandwich wall section.

3.4.1. Effect of overall slenderness ratio Fig. 9(a) shows the axial load (N) versus shortening (δ) relation curves for specimens SCSW-R2, SCSW-H1 and SCSW-H2. It is found pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi that slenderness ratio (λ ¼ Npl;Rk =Ncr ) have great influence on elastic stiffness and ultimate load resistance of sandwich wall. The elastic stiffness is reduced by 5% and 57% as the λ increases from 0.29 to 0.512 and from 0.29 to 0.925, respectively. Specimens SCSW-H1 and SCSW-H2 with larger slenderness ratio tend to fail in global buckling mode. However, only SCSW-H2 is observed to fail in global buckling with a lowest resistance. This is because the increase of slenderness ratio leads to smaller cross-sectional resistance reduction factor (χ). The compressive buckling resistance is governed by the slenderness ratio. Fig. 9(b) shows the influence of non-dimensional slenderness on compressive resistance of sandwich wall. As can be seen, the resistance is reduced by 40% when the slenderness ratio increases from 0.29 to 0.925. 3.4.2. Effect of spacing of connectors The connector spacing for SCSW-R2, SCSW-S1 and SCSW-S2 are 100, 200 and 300 mm which leads to different plate slenderness ratio (s/ts) of 16.7, 33.3 and 50. SCSW-S1 and SCSW-S2 are specimens with partial composite design in which local buckling of outer steel plates occurs showing a brittle failure. No obvious cracks or crushing of concrete are

Fig. 7. Instrumentations (S = strain gauge of steel, T = Transducer).

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observed in the specimens before reaching the maximum loads. Debonding of steel-concrete interface is found to initiate at the top region of wall at a load of 4656 kN and 3119 kN, respectively. This de-bonding develops promptly and then the outer plates buckle suddenly with a popping sound. After that, the steel face plates lose its load carrying capacity and the load transfers to the concrete core instantaneously, causing local crushing of concrete. On the other side of the wall, significant separation of steel-concrete interface is observed. Some concrete fragments are pushed out from the specimen surface. Fig. 8(b) illustrates the failure modes of SCSW-S1 and SCSW-S2, buckling phenomenon occurs between two adjacent rows of connectors. This indicates that the outer plates are not effective in providing lateral restraint to the concrete core if insufficient shear connectors with wider spacing are provided. Insufficient connector design results in partial composite action and lower compressive resistance. Fig. 10 shows the axial load-shortening relation curves for SCSW-S1 and SCSW-S2 compared to that of referential specimen SCSW-R2. From the figure, it is shown that the maximum load resistance and post-peak ductility of SCSW-S1 and SCSW-S2 decrease as the spacing of connector increases. The maximum loads reduce by 5% and 25% when the spacing increases from 100 to 200 and from 100 to 300 mm, respectively.

3.4.3. Effect of steel plate thickness By changing the thickness of steel face plates, SCSW-T8 and SCSW-T12 have smaller plate slenderness ratio (s/ts) of 12.5 and 8.3. It is found that the thickness of steel face plate significantly influences the maximum load resistance of the SCS sandwich wall, as shown in Fig. 11. These two specimens exhibit higher compressive resistance compared to referential specimen SCSW-R2. It is attributed to the higher steel contribution ratio and lateral restraint provided by the J-hook connectors. The final failure is due to cross-sectional failure. Fig. 8(c) shows the concrete core is fully crushed and the steel face plates of SCSW-T12 are fully yielded. At the back face of the sandwich specimen, shank fracture of a J-hook pair is also observed and below which a J-hook bar is yielded in tension arising from the outward buckling of steel plate at failure. Under compression, the concrete core is laterally restrained by the outer steel plates through interlocking J-hook connectors. As the load increases further, the crushing of the concrete core leads to a sudden uploading as observed from the axial load-shortening curves in Fig. 11. The J-hook connectors are under tensile force as the concrete core expands laterally pushing the steel plate outwards. The steel face plates tend to deform outward but their movements are constrained by J-hook connectors. The lateral restraining effect in SCSW-T8 and SCSW-T12 is found to be more effective than that

Fig. 8. Failure modes of SCS sandwich wall.

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151

Fig. 8 (continued).

of other specimens. The plate buckling of SCSW-T8 and SCSW-T12 is delayed compared to that of SCSW-R2 and the post-peak load displacement curve showing a more ductile behaviour. Fig. 12 shows the relation between normalised compression resistance and plate slenderness ratio including the test data of specimens SCSW-S1 and SCSW-S2. It is observed that the increase of plate slenderness ratio will reduce the normalised compressive resistance. However, the normalised compressive resistance difference is within 10% with s/t s ≤ 16.7(s ≤ 100 mm). The normalised maximum compressive load resistances are reduced by about 15% and 35% when the plate slenderness ratio s/ts increases from 8.5 to 33.3 (s = 200 mm) and from 8.5 to 50 (s = 300 mm), respectively. 3.4.4. Effect of different connector Fig. 13 shows the failure mode of SCSW-HS with headed shear studs. It is shown that de-bonding of steel-concrete interface initiates at the mid-height region of the wall. As the load increases, the de-bonding

propagates from the mid-height to both ends of the sandwich wall. The final failure is local buckling of the outer plate due to tensile separation at the steel-concrete interface. It is also observed that outer plate buckles in a multiple-wave mode while other specimens buckle in a single-wave mode. This implies that the tension separation resistance of using overlapped headed shear stud may be lower than that using the interlocking J-hook connectors in sandwich walls. The buckling phenomenon may initiate from several locations of the wall which produces multiple-wave buckling mode. However, the specimen SCSWHS with overlapped headed shear stud has 7% higher resistance compared to SCSW-R2. The small variance is due to the slightly higher concrete compressive strength for SCSW-HS. Fig. 14 compares the loadshortening curves of these two specimens. It can be observed that difference on the maximum load resistance is 7% while the elastic stiffness is less than 5%. This implies that the load displacement behaviour of the SCS sandwich wall with J-hook connector is comparable to the one with overlapped headed stud connectors.

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Table 6 Failure loads, failure modes and ductility index of SCS sandwich walls. Specimen SCSW-R1 SCSW-R2 SCSW-H1 SCSW-H2 SCSW-S1 SCSW-S2 SCSW-C1 SCSW-HS SCSW-C30 SCSW-C45 SCSW-T8 SCSW-T12 SCSW-P1 SCSW-P2 SCSW-C2 Mean Std.dev

Ntest (kN)

0.85Acfck + Asfy (kN)

4191 4906 4788 2975 4656 3670 4248 5317 3916 4689 6889 8418 5120 4933 5467

5403 5457 5466 5735 5516 5530 4832 5855 4445 5400 7099 8411 5954 5873 5767

SI 0.776 0.899 0.876 0.519 0.844 0.664 0.879 0.908 0.881 0.868 0.970 1.001 0.860 0.840 0.948 0.849 0.121

δtest (mm)

δ0.85 (mm)

3.93 3.56 4.85 8.91 4.11 3.73 3.97 4.65 3.44 4.08 4.47 4.99 3.81 4.27 4.73

5.85 4.03 6.02 9.46 4.11 4.09 4.2 6.05 3.84 4.16 5.57 8.13 4.77 5.42 5.41

μ 1.489 1.132 1.241 1.062 1.000 1.097 1.058 1.301 1.116 1.020 1.246 1.629 1.252 1.269 1.144 1.204 0.174

Failure mode CS+LB CS+CC+LB CC+LB GB LB LB CC+LB CS+LB CS+LB CS+CC+LB CSF+LB CSF+LB CC+LB CC+LB CS+CC+LB

* CS = concrete splitting; CC = concrete crushing; LB = local buckling of steel; CSF =

Fig. 10. Effect of connector spacing on load-shortening curves.

cross-sectional failure; GB = Global buckling; SI = strength index; µ = ductility index.

3.4.5. Effect of concrete compressive strength Fig. 15 shows the influence of concrete compressive strength on the maximum load resistance of the SCS sandwich wall. It is found that the maximum load resistance of the sandwich wall increases by about 20% and 25%, as the concrete strength increases from 30 to 45 MPa and from 30 to 60 MPa, respectively. The increase of concrete strength is approximately linearly proportional to the increase in the maximum load resistance, as shown in Fig. 15(b). The increase in maximum load resistance can be attributed to: (1) increase in concrete compression strength; and (2) the higher concrete core strength has less cracks and this provides better anchoring for the J-hook connectors leading to higher shear and tensile resistances.

3.4.6. Effect of concrete core thickness The concrete core thickness of SCSW-R2, SCSW-C1 and SCSW-C2 are 120, 80 and 136 mm. Fig. 16 plots the concrete core thickness versus the maximum load resistance of the three test specimens. The increase of maximum load resistance can be attributed to: (1) the increase of cross sectional area of the concrete; and (2) the larger embedded depth of J-hook connectors which provides high resistance for the connectors to resist tension force arising from the outward buckling of the face plates.

Fig. 9. Effect of slenderness ratioλ ¼

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi N pl;Rk =N cr .

Fig. 11. Effect of steel plate thickness on load-shortening curves.

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Fig. 12. Relation between normalised strength and plate slenderness (s/ts).

153

Fig. 14. Effect of different connector on load-shortening curves.

3.5. Behaviour of J-hook connectors 3.4.7. Effect of side plate confinement Open edge side plates and closed side plates are added to wall specimen SCSW-P2 and SCSW-P1, respectively. The side plates are used to seal the concrete core at the two sides but these side plates are not loaded in the tests, as shown in Fig. 4. Fig. 17 shows the failure modes of these two specimens. Local buckling and welding fracture at the corner are observed at failure. Fig. 18 compares the axial load-shortening curves of SCSW-P1 and SCSW-P2 with that of SCSW-R2, which does not have any side plates. It can be seen that there is no significant difference in the maximum load resistance of these three specimens. This indicates that the concrete confinement effect provided by the side plates is not pronounced in sandwich walls due to the larger cross-sectional aspect ratio compared to that of the composite column section. Local buckling of the face plates occurs only after the maximum load is attained as indicated in the load displacement curves in Fig. 18. The post-peak unloading curves of SCSW-P1 and SCSW-P2 are slightly more ductile as compared to SCSW-R2 (without side plates) which implies that the side plates are capable of delaying the post peak buckling of the face plates due to the additional side restraining effects.

The proposed J-hook connectors improve the structural performance for SCS sandwich wall by preventing the outward buckling of the face plates through an appropriate spacing of J-hook. This is because the local buckling of face plates only taking place between the two adjacent rows of level J-hook connectors, as shown in Fig. 19. Under compression, the J-hook pairs will subject to tension force, preventing the steel face plates from dilating outward. The concrete is constrained by steel plate so that the concrete is in a bi-axial compressive state. As expected, if no stiffener or sparser connectors are used, the J-hook connector cannot provide sufficient shear and tensile bond strength at steelconcrete interface thus lead to low level of composite action. The steel plate and concrete could share the loads independently and the steel plates could buckle outward in very early stage [4,5]. As observed from Table 5, specimens SCSW-S1 and SCSW-S2 have larger spacing of the connectors and thus the effective width to thickness ratio of the steel plate is higher. These two specimens failed by local buckling of the face plates. Once the face plate buckles and unloads, the concrete core has to resist the remaining load and this leads to the failure of

Fig. 13. Failure mode of SCSW-HS.

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Fig. 15. Effect of concrete compressive strength.

concrete core which is accompanied by an abrupt drop in load displacement curve. Large buckling wave heights of 200 mm and 300 mm are observed for SCSW-S1 and SCSW-S2, respectively, which correspond to the spacing of shear connectors. The ductility index of these two specimens is quite low. The load-shortening

curves and test observation show that SCSW-S1 and SCSW-S2 exhibit a brittle post-peak behaviour. On the contrast, by using smaller connector spacing or thicker plates which leads to smaller plate slenderness, the J-hook pairs can effectively restrain the transverse deformation of steel face plate. For example, for SCSW-T8 and SCSW-T12 with s/ts = 12.5 and 8.3, the failure mode of these two specimens is cross-sectional failure rather than local plate buckling mode. This indicates the material strength can be fully developed. After failure, a multiple-wave bulging pattern was observed on the face plates due to the progressive tensile fracture of hook shanks. 4. Compressive resistance of SCS sandwich wall 4.1. Eurocode 4 method Methods to predict the compressive resistance of normal composite columns under axial load are presented in Eurocode 4 [30] and AISC 360-10 [32]. In Eurocode 4, the characteristic cross-sectional resistance of concrete encased and partially encased steel section for composite column can be calculated by

Fig. 16. Effect of concrete thickness on maximum load resistance.

Nu ¼ Npl;Rk ¼ 0:85Ac f ck þ As f y

ð3Þ

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155

Fig. 17. Failure modes of SCSW-P1 and SCSW-P2.

For composite column, the buckling resistance may be predicted by multiplying the cross sectional resistance by a buckling reduction factor χ Nu ¼ χNpl;Rk where,χ ¼

ϕþ

ð4Þ p1ffiffiffiffiffiffiffiffiffiffiffi2 withλ ¼ ϕ2 −λ

qffiffiffiffiffiffiffiffi N pl;Rk Ncr

4.2. AISC method In AISC 360, the compressive resistance of encased composite members may be calculated by the following equations,

2

and ϕ ¼ 0:5½1 þ αðλ−0:2Þ þ λ ;α

is an imperfection factor corresponding to the appropriate buckling curve which can be obtained from Table 6.5 in Eurocode 4.

Fig. 18. Effect of side plate on load-shortening curves.

h P n0 i P n0 ≤2:25 Nu ¼ P n0 0:658 Pe ; when Pe

Fig. 19. Buckling mode of outer plates restrained by J-hook connector.

ð5Þ

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Table 7 Comparison of compressive resistance between calculated and test results. Specimen

Ntest (kN)

Nu_EC4 ( kN)

Nu_AISC ( kN)

Nu_Eq.(8) ( kN)

Nu_Eq.(9) ( kN)

Ntest/Nu_EC4

Ntest/Nu_AISC

Ntest/Nu_Eq.(8)

Ntest/Nu_Eq.(9)

SCSW-R1 SCSW-R2 SCSW-H1 SCSW-H2 SCSW-S1 SCSW-S2 SCSW-C1 SCSW-HS SCSW-C30 SCSW-C45 SCSW-T8 SCSW-T12 SCSW-P1 SCSW-P2 SCSW-C2 Mean Std.dev

4191 4906 4788 2975 4656 3670 4248 5317 3916 4689 6889 8418 5120 4933 5467

5403 5457 5457 5457 5421 5457 5801 5466 5735 5516 5531 4832 5855 4445 5400

5188 5245 4833 3886 5260 5176 4533 5611 4301 5186 6826 8169 5716 5642 5581

7458 7565 7808 7904 5818 5505 6158 8048 6312 7454 8703 10890 7724 7766 8263

8178 8248 7879 8696 7157 6565 7597 8890 6627 8174 12223 17179 9123 9047 8619

0.776 0.899 0.876 0.519 0.844 0.664 0.879 0.908 0.881 0.868 0.970 1.001 0.860 0.840 0.948 0.849 0.117

0.808 0.935 0.991 0.766 0.885 0.709 0.937 0.948 0.910 0.904 1.009 1.030 0.896 0.874 0.980 0.905 0.089

0.562 0.649 0.613 0.376 0.800 0.667 0.690 0.661 0.620 0.629 0.792 0.773 0.663 0.635 0.662 0.672 0.070

0.512 0.595 0.608 0.342 0.651 0.559 0.559 0.598 0.591 0.574 0.564 0.490 0.561 0.545 0.634 0.574 0.043

*Ntest = failure load by test; Nu_EC4 = prediction by Eurocode 4 (encased section); Nu_AISC = prediction load by AISC 360-10; Nu_Eq.(8)= prediction by Eq. (8); Nu_Eq.(9)= prediction by Eq. (9); The data of SCSW-H2 is not included in calculating the mean value and standard deviation.

Nu ¼ 0:877P e ; when

P n0 N 2:25 Pe

ð6Þ

where,   0 P n0 ¼ F y As þ 0:85f C Ac and P e ¼ π2 EIeff =ðKLÞ2

ð7Þ

Eqs. (3) and (5), which are developed for composite columns, are found to over-predict the compressive resistance of SCS sandwich wall. The predictions are compared with the test results in Table 7. The average ratios ofNtest/Nu_EC4and Ntest/Nu_AISCare 0.849 and 0.905, with standard deviations of 0.117 and 0.089, respectively. The codes’ criteria were developed to suit a variety of structural elements and therefore some variability is expected when compared to the test results. Eqs. (3) and (5) leave out the effect of the J-hook connectors on the compression resistance of the sandwich composite walls. The codes’ equations predict the same compression resistance with or without the interlocking effect of the J-hook connectors. Although the J-hook connectors cannot directly enhance the compressive resistance of the composite wall, they provide effective lateral restraints to the face plates to confine the lateral dilation of the concrete core under compression. Several researchers have investigated the compressive resistance of concrete filled tubular (CFT) columns with binding bars, tie rods and

external rings [4,5 and 24–26]. It was concluded that these reinforcement enhancements indeed improved the compression resistance and ductility of CFT composite columns because of the confinement effect from the face plates due to the present of the binding bars and tie rods. The corresponding prediction models were also derived based on calibration with the experimental results, as given below, Nu ¼ ϕ1 f y As þ ϕ2 f ck Ac ( where ϕ1 ¼

0:89 0:897R−0:7407

ð8Þ R b 0:85 R ≥ 0:85

) and ϕ2 = 1.039R−0.0861

(7.3836ς + 1.0588) [26].   2 0 Nu ¼ A f scy ¼ Aa0 b0 f scy ¼ Aa0 b0 1:1 þ Bξ þ Cξ f ck

ð9Þ

where a0 = 1.54(a/t)−0.12,b0 = −0.10 ln (s/t) + 1.34,B = 0.1381(fy/ 235) + 0.7646, C = 0.0727(fck/20) + 0.0216 and ξ = fyAs/(fckAc) [4,5]. Table 7 compares the compressive resistance obtained by the other researchers' models and the test results. Fig. 20 plots the test results and prediction by Eqs. (8) and (9). It is found that Eqs. (8) and (9) over-predict the compressive resistance of SCS sandwich wall with Jhook connectors. This is because in CFT composite columns the external steel tube provides effective confinement to concrete due to smaller aspect ratio (b 2.0) compared to that of the composite walls with larger aspect ratio (N4.5). For rectangular cross section with smaller aspect ratio (width-to-depth ratio), the concrete can be confined by steel plates and the strength of the concrete increases due to the concrete being in a triaxial compressive state. However, in the case of sandwich composite wall, the lightweight concrete core is only confined in one direction and the concrete core is un-confined in the other direction. As a result of this the concrete core tends to crush with longitudinal cracks. This type of failure mode is different from the failure mode observed for short CFT columns in which the concrete failed by crushing without longitudinal cracks. Therefore, the confinement effect for concrete in SCS sandwich wall is not pronounced and should be ignored in the design consideration. Based on the above discussions, a new formula shall be developed to predict the compressive resistance of SCS sandwich wall infilled with ultra-lightweight cement composite. 4.3. Compressive resistance of composite wall

Fig. 20. Comparison of test results with predicted results.

For simplification, the following assumptions may be applied: (1) the compressive resistance of sandwich wall should be calculated

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157

Table 8 Published test data on lightweight double skin sandwich wall. Specimen

Pc (kN)

σf (MPa)

σf/fck

n

Ah (mm2)

Asc (mm2)

σc (MPa)

σc/fck

1 2 3 4 5

389 314 378 325 270 149 191 127 1868

5.66 4.57 5.50 4.73 3.93 2.17 2.78 3.37 9.29

0.719 0.581 0.699 0.601 0.500 0.276 0.353 0.571 0.300

30 20 32 6 0 0 0 6 0

0 0 0 1566 1566 0 0 0 5400

1508 1005 1609 1085 783 0 0 170 2700

66.1 44.1 70.5 47.6 34.3 0 0 29.7 46.0

8.40 5.60 8.97 6.05 4.36 0 0 5.03 1.48

PP1 PP2 NSE1* 7

References

[16]

[17] [12]

*Pc = load on concrete only; fy = 280 MPa and Ea = 200 GPa for all tests. Prabha et al. [16]: fck = 7.87 MPa, W = 685 mm, tc = 120 mm, dstu = 8 mm, H = 870 mm; Ah = 2tH;Asc = (0.5Ah + nAstu); Mydin and Wang [17]: fck = 5.9 MPa, W = 400 mm, tc = 100 mm, dstu = 6 mm, H = 400 mm; Wright [12]: fck = 31 MPa, W = 914 mm, tc = 220 mm, dstu = -nil- ; H = 1800 mm. n = number of connecting bolts.

by adding the resistance of its components;

where, the projected area of cone surface to the free concrete surface 2

NuI ¼ ϕa As f y þ ϕc Ac f ck

ð10Þ

where ϕa,ϕc are reduction factors for steel and lightweight concrete respectively; As,Ac are cross sectional area of steel plates and concrete core; fck,fy are concrete compressive strength and yield strength of steel. (2) All the steel across the concrete failure plane can be assumed to provide confinement. The confinement stress σc may be expressed by σc ¼

nJP J WH

ð11Þ

where, nJ is the number of J-hook shear connectors, which can be determined bynJ = HW/sHsW; H,W,sH,sW are the height of the wall, the width of the wall, the connector spacing in height direction and the connector spacing in width direction. PJis the tensile resistance of J-hook connector in sandwich structures (see Fig. 19). It should be noted that there are four major failure modes of J-hook connector under tension which are breakout failure of concrete material (Pcb), pull-out failure of J-hook connector (Ppl), tensile fracture failure of J-hook connector (Pu) and punching shear failure of steel plate (Pps) as used in Ref. [8], 8 qffiffiffiffiffiffiffi > > > P cb ¼ 0:33 f ck AN > < 2 P J ¼ min P pl ¼ 0:9φf ck eh d þ 0:116d f y > > P ¼ ϕA f u se > utpffiffiffi > : P ps ¼ Av f u = 3

AN ¼ πhef ð1 þ hdh Þ; φ = 1.4 for an anchor located in a region of a conef

crete member where analysis indicates no cracking, otherwise φ = 1.0; eh is the distance from the inner shaft of a J- or L-bolt to the outer tip of the J- or L-bolt, and3d ≤ eh ≤ 4.5d; d is the diameter of stud; ϕ is the reduction factor of the steel; Ase and fut are the cross-sectional area and the ultimate strength of headed shear stud; Av = πdt and fu are the punching area and ultimate strength of steel plate. Introducing the non-dimensional form to Eq. (11) results in σ c 100n J P J ¼ f ck WH f ck

(3) Assume a linear relationship for the concrete failure stress σf against the confinement stress σc under splitting failure mode of concrete core, by regression method, an equation can be derived to describe the failure stress of any composite wall panel. In the regression model, some test data on double skin lightweight composite walls under compressive loading from other researchers [16,17 and 12] are also included as shown in Table 8. In these tests, the composite walls were confined using through-through bolts. Concrete failure stress σf can be determined by σf ¼

ð12Þ

Fig. 21. Regression of failure stress ratio (σf/fck) with confinement stress ratio (σc/fck).

ð13Þ

Nc Ac

ð14Þ

where, Nc is the failure load carried by concrete and Acis the cross sectional area of concrete. For the case where the applied load was shared by the steel and concrete, Nc = Ntest −Ns while for the case where the load was applied only on the concrete, Nc = Ntest.

Fig. 22. ϕa relationship based on regression method.

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Table 9 Comparison of compressive resistance between predictions and test results. Predictions by proposed model Ntest (kN)

Specimen SCSW-R1 SCSW-R2 SCSW-S1 SCSW-S2 SCSW-H1 SCSW-H2 SCSW-C1 SCSW-HS SCSW-C30 SCSW-C45 SCSW-T8 SCSW-T12 SCSW-P1 SCSW-P2 SCSW-C2 Mean. COV.

4190.7 4906.2 4656.4 3670.4 4787.6 2975.2 4248.4 5317.0 3915.6 4688.5 6889.7 8418.0 5119.8 4933.0 5466.7

ϕcAcfck (kN)

ϕaAafy (kN)

NuI (kN)

3523.2 3589.0 3697.0 1907.8 1702.9 2724.4 3730.4 2651.8 3529.6 3588.6 3510.3 3608.4 3578.6 4024.6

1159.4 1159.4 839.8 1159.9 1159.4 1159.0 1159.0 1159.4 1159.0 2238.3 4453.3 1395.1 1395.2 1159.0

4682.6 4748.4 4536.8 3067.6 2862.3 3883.4 4889.4 3811.3 4688.5 5826.9 7963.6 5003.5 4973.9 5183.5

Ntest/NuI

NuII (kN)

Ntest/NuII

0.89 1.03 1.06 1.52 1.28 1.09 1.09 1.03 1.00 1.18 1.06 1.02 0.99 1.05 1.09 0.14

4793.8 4859.6 4661.0 3179.4 2973.4(4804.9) (3700.7) 3993.9 4999.8 3922.4 4799.0 5923.0 7442.4 4877.9 4849.2 5294.0

0.87 1.01 1.03 1.46 1.23(1.00) (0.80) 1.06 1.06 1.00 0.98 1.16 1.13 1.05 1.02 1.03 1.08 0.13

*NuI = ϕaAafy + ϕcAcfck(Eq. (10)), NuII = befftfy + ϕcAcfck(Eq. (24)). The value in the brackets are based on Eurocode 4 method (Eq. (4)), which are not included in calculation of Mean and COV.

Using linear regression approach, an equation for the concrete failure stress against confinement stress in SCS sandwich composite walls with connectors is derived as illustrated in Fig. 21, ϕc ¼

  σf 100n J P J ¼ 0:2995 þ 0:0489 f ck WH f ck

ð15Þ

It can be found that even considering confinement effect, the lightweight wall core may not able to achieve its cross-sectional resistance (ϕc ≤ 1.0). Such finding was also confirmed by Prabha et al. [16] and Mydin and Wang [17]. This may be attributed to the larger width to thickness ratio of the SCS sandwich wall compared to composite column with smaller aspect ratio. Therefore, a strength reduction factor ϕc should be applied to lightweight concrete in SCS sandwich walls to account for this effect. (4) The failure of the steel face plates are assumed to obey the vonMises rule given as 2

2

f a −f a σ h þ σ 2h ¼ f y

Fig. 19 shows the typical buckling model of outer plates of SCS sandwich wall. To predict the local buckling strength of steel plate fb in concrete filled tube columns, Ge and Usami [33] proposed an equation as given by f b 1:2 0:3 ¼ − 2 ≤1:0 fy R R

ð17Þ

where R is the width-thickness ratio, defined as b R¼ t

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   12 1−νs 2 f y ≥0:85 4π2 Ea

ð18Þ

in which, b is the breadth of section; t is the thickness of steel plate; Ea and νs are Young’s modulus and Poisson ratio of steel plate. Combined Eq. (16) to Eq. (18), the stress states of steel plates can be determined by applying fa = fb.

ð16Þ compression : f a ¼

where, fa is compressive stress of steel plate ; σh is transverse tensile stress of steel plate; fy is yield strength of steel plate. tension : σ h ¼

  1:2 0:3 − 2 fy R R

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2 f a − 4f y −3f a 2

ð19Þ

ð20Þ

Based on the regression (Fig. 22) of test data from Ref. [26] and from the current research, ϕa is given as ϕa ¼ 0:898R−0:771 ðR ≥0:85Þ

ð21Þ

For thin-walled structures, the ϕc value in Eq. (15) and ϕa value in Eq. (21) can be used in the proposed formula Eq. (10) to obtain the cross-sectional resistance of composite wall considering local buckling of thin steel plates and concrete failure. Liang and Uy [19] also proposed a method based on effective width concept to account for post-local buckling behaviour of steel plates in steel box columns infilled with concrete.

Fig. 23. Comparison between test results and predictions by proposed methods.

beff σ cr ¼ 0:675 b fy

!1=3 for σ cr ≤ f y

ð22Þ

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159

Fig. 24. Typical finite element model for SCS sandwich wall.

beff σ cr ¼ 0:915 b σ cr þ f y

!1=3 for σ cr N f y

where, the critical elastic buckling stressσ cr ¼

ð23Þ kπ 2 Es , k is the elas12ð1−νs 2 Þðb=tÞ2

tic buckling coefficient accounts for the effect of the plate aspect ratio and boundary condition on the critical buckling stress. Uy and Bradford [18] proposed k = 5.6for steel plates in contact with elastic medium. Therefore, the cross-sectional resistance of SCS sandwich composite wall can be calculated by the following equation that replaces the first term on the right hand side of Eq. (24), as follow: NuII ¼ beff t f y þ ϕc Ac f ck

ð24Þ

4.4. Validation In general, the proposed formulae give reasonable estimations on ultimate resistance of sandwich wall under compressive loading. Table 9 summarises the calculation results and comparison to the experimental results. The average ratios of test result to prediction Ntest/NuIand Ntest/ NuII are 1.09 and 1.08, with coefficients of variance 0.14 and 0.13 respectively. Fig. 23 depicts the scatter plot between the test results and predictions by proposed methods. The comparisons show that the two proposed formulae give fairly reasonable predictions in the compressive resistance of SCS sandwich wall. For specimen SCSW-H2 which fails in global buckling, Eq. (4) in Eurocode 4 is used to predict the ultimate load resistance considering the second order effect. It is found that Eq. (4) over-estimates by 20% the buckling resistance of slender SCS sandwich wall with J-hook connectors, as seen the bracket value in Table 9. This may be due to the interaction of local and global buckling occurs in this specimen. Since only one slender wall is tested in this study, more test data should be collected to evaluate the design formulae in the future study. 4.5. Design recommendations for SCS sandwich wall 4.5.1. Full composite design The steel plate slenderness ratio (s/ts) has a great influence on the compressive resistance and ductility behaviour. With appropriated s/ts ratio, the sandwich wall would have favourable ductility and the material strength could be fully utilised. Also, it is found that providing sufficient number of J-hook connectors can ensure full bond (shear bond and

tensile bond) at the steel-concrete interface. For structural efficiency and to avoid local buckling of the face plate, full composite interaction is required between the core and the steel face plates. If insufficient connectors are provided, this may cause bond failure which results in premature buckling of steel face plates. Then the elastic stiffness, maximum resistance and the post-peak ductility of the sandwich wall will degrade. Therefore, full composite design is recommended. Based on the test investigations, it is recommended that the connector spacing qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi should not be spaced more than 20t s 235=f y , which is similar to Eurocode 4 class 3 requirement to avoid local buckling of face plates under compression. Otherwise, local buckling of steel plates must be accounted for by using the effective width method as proposed in Section 4.3. 4.5.2. Design method to predict compressive resistance of SCS sandwich wall Step I: Determine the failure mode of J-hook connector under tension force so as to choose tensile resistancePJ by Eq. (12) to determine the concrete strength reduction factor ϕc by Eq. (15); Step II: Determine the steel buckling reduction factor ϕa by Eq. (21); Step III: Determine the compressive resistance of SCS sandwich wall with J-hook connectors by Eq. (10) or Eq. (24). 5. Finite element analysis 5.1. Numerical modelling Finite element analysis is performed using ABAQUS 6.13 to study the compressive behaviour of SCS sandwich walls with J-hook connectors. Eight-node solid element with reduced integration point (C3D8R) is used to model the steel plate, J-hook bars and ULCC core. The Concrete Damage Plasticity (CDP) model is employed to model the nonlinear behaviour of ULCC core while elastic-plastic model is used to model the steel face plate and steel bar. The mechanical properties of ULCC and steel plates obtained from the tests are used in the FE model. Each pair of J-hook connectors are modelled using two solid bars connected by a nonlinear connector element. Shear and flexural deformations of the connectors can be considered by the solid bars whereas the connector element is used to represent the tension-elongation relationship between the top and bottom connectors. The connector element is coupling at the centre point of both bar section surfaces and a forcedisplacement relationship based on pull-out test of J-hook connector embedded in concrete is assigned to the connector element [34]. A

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Fig. 25. Comparison of failure mode of SCS sandwich wall (left: FE; right: Test).

mesh convergence study has been conducted to find a suitable mesh size to each component. Surface-to-surface contact algorithm is adopted to simulate the interfacial bond between concrete and steel. "Hard contact" in the normal direction and "slide contact" in the tangential direction are selected for analysis. A friction coefficient of 0.4 is assumed in the analyses. Sensitivity analyses have been carried out using friction coefficient 0.2 to 0.6 and the results showed that the friction coefficient led to little difference in both overall load-shortening response and failure mode. The difference in the predicted maximum load obtained from various friction coefficients is less than 5%. In the FE model, initial geometrical imperfection of the slender wall specimen is considered based on the measured value by using inclinometers. Fig. 24 shows the typical finite element model of SCS sandwich wall under compression. More detailed information on the FE model and parametric studies is reported by Huang and Liew [34–36].

5.2. Numerical results and comparisons Fig. 25(a) and (b) compare the primary failure modes between FE and test results. As shown in Fig. 25(a), the concrete crushing and local buckling of steel face plate are captured accurately for specimen SCSW-R2. For the slender sandwich wall specimen SCSW-H2, the FE model also captures the occurrence of global buckling, as shown in Fig. 25(b). The corresponding load-shortening relations are plotted in Fig. 26(a)-(b). A close agreement between the test and numerical results can be observed. Table 10 compares the maximum loads obtained from FE, analytical and test results. The average ratio of Ntest/ NFE, NFE/NuI and NFE/NuII are 1.05, 1.00 and 1.00 with coefficients of variance of 0.05, 0.06 and 0.07. The results obtained from FE analysis are verified against the test results to establish its accuracy in predicting load-displacement curves, maximum resistance and failure modes of the sandwich walls. This confirms that the analytical

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Fig. 26. Comparison of load-shortening curves of SCS sandwich wall.

formulae are conservative for the design of axially compressed SCS sandwich composite walls with J-hook connectors. 6. Conclusions This study investigates the compressive behaviour and maximum resistance of sandwich composite wall system with J-hook connectors infilled with the ultra-lightweight cement composite material. A series of compression tests on short and slender SCS sandwich walls have been carried out. The experimental programme consists of a wide range of parameters such as overall slenderness of sandwich wall, steel contribution ratio, concrete compressive strength, side plates, connector type and spacing of shear connectors. Analytical method has been proposed to predict the compressive resistance of lightweight cement composite infilled SCS sandwich composite wall. Nonlinear finite element analysis has been performed to further investigate the structural behaviour of SCS sandwich wall. The studies reported herein support the following observations and conclusions: (1) Test observation shows that the lightweight SCS sandwich wall with J-hooks exhibits comparable behaviour to the same wall specimen with overlapped headed stud connectors in terms of compressive resistance and ductility. Four types of failure modes are observed from the tests, namely crushing of cementitious core, steel and concrete interfacial bond failure, crosssectional failure and overall buckling of slender wall depending on the slenderness of the steel plates and height of the wall; (2) From the parametric study, it is found that: (a) an increase of

Table 10 Comparison of compressive resistance between predictions and test results. Maximum load comparison Specimen

Ntest (kN)

NFE (kN)

NuI (kN)

NuII (kN)

Ntest/NFE

NFE/NuI

NFE/NuII

SCSW-R1 SCSW-R2 SCSW-S1 SCSW-S2 SCSW-H2 SCSW-C1 SCSW-C30 SCSW-C45 SCSW-T8 SCSW-T12 SCSW-P1 SCSW-P2 SCSW-C2 Mean. COV.

4190.7 4906.2 4656.4 3670.4 2975.2 4248.4 3915.6 4688.5 6889.7 8418.0 5119.8 4933.0 5466.7

4295.8 4727.8 4394.3 3376.7 2826.0 4013.8 3626.9 4614.1 5939.2 8460.7 5117.1 4636.1 5554.0

4682.6 4748.4 4536.8 3067.6 3883.4 3811.3 4688.5 5826.9 7963.6 5003.5 4973.9 5183.5

4793.8 4859.6 4661 3179.4 3700.7 3993.9 3922.4 4799 5923 7442.4 4877.9 4849.2 5294

0.98 1.04 1.06 1.09 1.05 1.06 1.08 1.02 1.16 0.99 1.00 1.06 0.98 1.05 0.05

0.92 1.00 0.97 1.10 1.03 0.95 0.98 1.02 1.06 1.02 0.93 1.07 1.00 0.06

0.90 0.97 0.94 1.06 0.76 1.00 0.92 0.96 1.00 1.14 1.05 0.96 1.05 1.00 0.07

(3)

(4)

(5)

(6)

concrete thickness and concrete strength offer higher compressive resistance of composite wall; (b) increase of the wall slenderness ratio will reduce the compressive resistance; (c) smaller spacing of connector provides higher confinement to the concrete core and thus increase the plate buckling resistance and the compressive resistance of sandwich wall. The limqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi iting connector spacing s≤20t s 235=f y is recommended for practical design; (d) confinement effect provided by the side plates is not pronounced for rectangular SCS sandwich wall although it slightly improves the post-peak unloading behaviour. The composite column design method in Eurocode 4 and AISC 360 could over-predict the compressive resistance of sandwich composite wall. This is due to the different aspect ratio between the SCS sandwich wall and column, leading to different failure mechanisms such as concrete splitting failure, crushing failure or the combined of both in sandwich composite wall. A modified design method based on Eurocode 4 approach is proposed to predict the compressive resistance of SCS sandwich wall. A regression method is used to obtain the strength reduction of lightweight concrete core and local buckling of steel face plate. The proposed design method has been validated against the test results and is found to be adequate and conservative. The validation confirms the accuracy of proposed FE model over the test results on predicting the compressive behaviour of the sandwich wall with novel J-hook connectors. The numerical study presented herein path the way to develop design guideline to predict the compressive resistance of SCS sandwich walls. The use of inter-locking J-hook connectors in SCS sandwich wall prevents hydrostatic pressure deformation when concreting at the construction stage. The studies show that the inter-locking J-hook connectors are subjected to tension force due to the lateral expansion of cement composite core under compression. This signifies the important role of the interlocking effect of J-hook connectors in reducing the initial out-of-plane deformation of the face plates and improving the buckling resistance of the composite wall.

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