Seismic behavior of connections composed of CFSSTCs and

Journal of Constructional Steel Research 64 (2008) 680–688www.elsevier.com/locate/jcsr

Seismic behavior of connections composed of CFSSTCs and steel–concretecomposite beams — finite element analysis

Jianguo Niea, Kai Qinb, C.S. Caic,∗

a Structural Engineering Research Laboratory, Department of Civil Engineering, Tsinghua University, Beijing, 100084, Chinab China Development Bank Jiangsu Branch, Nanjing, Jiangsu Province, 210024, China

c Department of Civil and Environmental Engineering, Louisiana State University, Baton Rouge, LA 70803, United States

Received 3 November 2006; accepted 11 December 2007

Abstract

Based on the experimental results of connections composed of concrete-filled square steel tubular columns (CFSSTCs) and steel–concretecomposite beams with interior diaphragms, exterior diaphragms, or anchored studs, 3-D nonlinear finite element models were established toanalyze the mechanical properties of these three types of connection using ANSYS. Finite element analyses were conducted under both monotonicloading and cyclic loading. The load–displacement and shear force–deformation curves of the finite element analyses are in agreement withthose of the tests in terms of strength and unloading stiffness. Parametric analyses were conducted on the connections with exterior diaphragmsunder monotonic loading to investigate the influences of axial load ratio, width to thickness ratio, and dimensions of exterior diaphragms on theconnection behavior. It was found that the strength and stiffness are less influenced by the axial load ratio and the dimensions of the exteriordiaphragms, but more influenced by the width to thickness ratio of the steel tube under shear failure mode.c© 2008 Elsevier Ltd. All rights reserved.

Keywords: Connections; Interior diaphragms; Exterior diaphragms; Anchored studs; Finite element analysis; Parametric analysis; Monotonic loading; Cyclic loading

1. Introduction

In the past two decades, the incorporation of connectionbehavior into the design of moment resisting frames hasattracted much attention since modeling the real behaviorof the connections leads to more reliable and/or economicaldesigns in construction. Much of the knowledge neededto apply composite connection behavior to design hasbeen derived from detailed finite element analyses of theconnections. These finite element models have been usedfrequently to develop moment–rotation curves, to verify designmethodologies and concepts, and to assess the local behaviorof the connection components, such as interior diaphragms orexterior diaphragms.

Because of the need for practical applications, somefinite element analyses were carried out on the mechanicalproperties of connections composed of concrete-filled steeltubular columns and steel beams. Using ABAQUS, Alostaz

∗ Corresponding author. Tel.: +1 225 578 8898; fax: +1 225 578 8652.E-mail address: [email protected] (C.S. Cai).

0143-974X/$ - see front matter c© 2008 Elsevier Ltd. All rights reserved.doi:10.1016/j.jcsr.2007.12.003

et al. [1] conducted a nonlinear 3-D finite element studyon a variety of connections where beams were connected toconcrete-filled circular steel tubes. Numerical results suggestedthat connections that transfer loads from the girder to theconcrete core potentially offer better seismic performance thanconnections where beams are connected to the steel tubealone. Based on the experimental results, the welded split-teeconnections composed of concrete-filled square steel tubularcolumns (CFSSTCs) and wide flange beams were modeledusing ABAQUS by Peng [2]. The finite element modelswere verified by comparing the behavior of the models withexperimental results. The comparison showed good agreementin terms of stiffness, load–displacement relationship, straindistribution, and variation of the bolt forces. Chiew et al. [3]established a finite element model of connections composedof concrete-filled circular steel tubular columns and I-shapedsteel beams using MARC. Subsequently, the key parameterswere studied numerically through the finite element methodand an empirical formula was derived based on more than 100numerical results of parametric analysis.

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http://dx.doi.org/10.1016/j.jcsr.2007.12.003

J. Nie et al. / Journal of Constructional Steel Research 64 (2008) 680–688 681

Cyclic loading tests on eight cruciform connections andfinite element analyses using ABAQUS were carried out byKang [4], and both the numerical and experimental results werefound to be in good agreement with one another. According tothe results of the finite element analysis, by using reinforcingT-stiffeners, the failure mechanism changed from the momentyield line of the round corner of the column wall to atensile yield failure mode. Six full-size specimens composed ofconcrete-filled square tubular columns and I-shaped steel beamswere tested under cyclic loading by Zhou [5]. Based on theexperimental results, a finite element analysis model involvinggeometric large deformation, contact-friction of high-strengthbolts, and nonlinear material properties was established usingANSYS. Some suggestions for improving connection designwere proposed according to the comparison between thetheoretical and experimental results.

Four types of steel beam to rectangular concrete-filledsteel tubular column connection, including normal weldedflange plate (WFP) connections, bolted flange plate (BFP)connections, stiffened end plate (SEP) connections, and doublesplit-tee plate (DST) connections, were tested by Lin [6].Subsequently, finite element analyses were conducted on thecyclic behavior of the WFP and BFP connections usingANSYS. The finite element analysis results were not consistentwith the experimental ones for reasons of inaccurate materialproperties and not simulating the welds. Although connectionscomposed of concrete-filled steel tubular columns and steelbeams are widely studied through the finite element method,research on the mechanical properties of connections composedof concrete-filled steel tubular columns and steel–concretecomposite beams is rare. Therefore, it is necessary toinvestigate the seismic behavior of this type of connection.

As a part of a series of investigations on the seismic behav-ior of composite connections, fourteen cruciform connectionspecimens composed of CFSSTCs and steel–concrete compos-ite beams with interior diaphragms, exterior diaphragms, or an-chored studs were tested under cyclic loading and reported inthe companion paper [7]. Based on the test data, in the presentstudy finite element analyses were carried out to analyze themechanical behavior of these three types of connection usingANSYS. Subsequently, parametric analyses were conducted onthe connections with exterior diaphragms to investigate the in-fluences of axial load ratio, width to thickness ratio, and dimen-sions of exterior diaphragms.

2. Finite element model

2.1. Element types

Five kinds of element were adopted in the finite elementmodels:

• Solid65 element. This element is defined by eight nodeshaving three degrees of freedom at each node, namelytranslations in the nodal x , y, and z directions. It is used for3-D modeling of solids with or without rebars. The solid iscapable of cracking in tension and crushing in compression.This type of element was used to model the in-filled concreteand the slab concrete.

Fig. 1. Trilinear stress–strain model used for steel. (σs = stress, εs = strain,fy = yield strength, fu = ultimate strength, εA = yield strain, εB =

ultimate strain, and εC = maximum strain of steel, respectively.)

Fig. 2. Bilinear stress–strain relation model used for rebars. (σr = stress, εr= strain, fr y = yield strength, fru = ultimate strength, εr y = yield strain,Es = elastic modulus, and Et = hardening modulus of rebars, respectively.)

• Solid45 element. This element is defined by eight nodeshaving three degrees of freedom at each node, namelytranslations in the nodal x , y, and z directions. It was used tomodel the rollers and supports.

• Shell181 element. This is a four-node element with sixdegrees of freedom at each node, namely translations in thex , y, and z directions, and rotations about the x , y, and zaxes. It was used to model the steel tubes and beams.

• Link8 element. This 3-D spar element is a uniaxialtension–compression element with three degrees of freedomat each node, namely translations in the nodal x, y, and zdirections. It was used to model the longitudinal rebars inthe concrete slabs.

• Beam4 element. This is a uniaxial element with tension,compression, torsion, and bending capabilities. The elementhas six degrees of freedom at each node, namely translationin the nodal x, y, and z directions and rotations about thenodal x, y, and z axes. It was used to model the studsanchored in the columns and the studs between the steelbeams and the concrete slab.

2.2. Material properties

The Von Mises yield criterion with kinematic hardening rulewas adopted to model the steel material in the finite elementanalysis. The trilinear stress–strain relation was used to modelthe steel plates of the steel tubes and steel beams as shownin Fig. 1, where εA = fy/Es , εB = 10( fu − fy)/Es , and εC =

0.03. The bilinear stress–strain relationship, as shown in Fig. 2,was used to model the rebars, where Et = 0.1Es . For the steelmaterial in the finite element model, the elastic modulus Es andPoisson’s ratio υs were assumed as 2.06 × 105 MPa and 0.3,respectively.

682 J. Nie et al. / Journal of Constructional Steel Research 64 (2008) 680–688

Fig. 3. Stress–strain relation model used for in-filled concrete. (σc = stress ofconcrete, and εc = strain of concrete.)

Fig. 4. Stress–strain relation model used for slab concrete.

Solid65 Element in the ANSYS program was used tomodel the concrete. The William–Warnker concrete model [8]was adopted as the failure criterion for the in-filled concreteand slab concrete where nine parameters are required asinputs to describe the model. The in-filled concrete was underconfining pressure, and as a result, the material property wasdifferent from that of the slab concrete. Han’s stress–strainrelationship [9] was simplified as a multi-linear model, whichwas used for the in-filled concrete in the ANSYS finite elementmodel (see Fig. 3). Guo’s stress–strain relationship [10] wasalso simplified to model the slab concrete, as shown in Fig. 4.The Poisson’s ratio of concrete υc was assumed as 0.2, and theelastic modulus of concrete was calculated as

Ec =105

2.2 +34.7fcu,k

(MPa) (1)

where fcu,k is the cube strength of concrete.

2.3. Finite element mesh and boundary conditions

Three-dimensional numerical models were established torepresent the test specimens. The finite element meshes of theconnections with interior diaphragms, exterior diaphragms, oranchored studs are shown in Fig. 5(a)–(b), (c)–(d), and (e)–(f),respectively.

The interface between the steel tube and concrete core wasassumed to have a perfect bond condition. As a result thelongitudinal slip between the concrete and the steel tube isneglected in the finite element analysis. On this aspect, the studsare not effective. However, the studs are somewhat effectivein transferring transverse tension forces. Also, the existence ofthe anchored studs influences the strain and stress distributionof the in-filled concrete, which was hard to measure in theexperiments. Therefore, the anchored studs are modeled toinvestigate these effects.

To simulate the experiments, the same loading proceduresand constraints as the experiments were used in the finiteelement analysis. The numerical models were loaded in twosteps. The axial load was applied to the column at first, and thenthe monotonic or cyclic loads were applied at the ends of thebeams. The rollers of the test setup were modeled in ANSYSand the axial load was applied to them. The displacements in thex and y directions and the rotations about the x and z axes of themiddle line at the top of the upper rollers were constrained. Alsothe displacements in the x, y, and z directions and the rotationsabout the x and z axes of the middle line at the bottom of thelower rollers were constrained. Therefore, rotation about the yaxis of the model and displacement in the z direction of the topsupport are free. Thus the rollers of the test setup were formedin the ANSYS model and the axial forces can also be applied onthis model. The Newton–Raphson equilibrium iteration methodwas used to solve these nonlinear problems.

3. Numerical results

3.1. Results under monotonic loading

Numerical load–deformation curves (called P–∆ curves forconvenience) under monotonic loading are compared with theexperimental P–∆ skeleton curves, as shown in Fig. 6. Thecomparisons of the yield strength and ultimate strength of theconnections are shown in Table 1. As the curves predictedby ANSYS did not decline at the end of the analysis, thecorresponding load when the story rotation reached 0.06 radwas taken as the ultimate strength. The ultimate strengths of thetests are defined as the maximum strengths of the specimens.The yield strength was obtained by a graphical method [11].The finite element analysis results show fairly good agreementwith the experimental ones in terms of strength and stiffness forconnections with interior or exterior diaphragms, while for theconnections with anchored studs there are certain differencesfor stiffness and positive strength, perhaps due to the samereason as discussed below.

The Q j –γ j curves obtained from the numerical analysesand experimental tests are compared in Fig. 7. It can beseen that the two curves agree well for the connections withinterior or exterior diaphragms, while a larger difference isfound for the connections with anchored studs. The reason forthis difference is that while the punching shear failure modeor fracture of the welds between the steel tube flange andweb occurred during the test for this type of connection, it isdifficult to simulate these failures in the finite element model.Therefore, the numerical and experimental results for the sheardeformation and stiffness are different.

As has been illustrated in the companion paper [7], everycomponent of the connections deforms when the subassembliesare loaded under cyclic loading. For example, when an exteriorconnection is loaded, the displacement at the end of the beam∆ is composed of the displacement caused by the beam ∆b, thedisplacement caused by the column ∆c, and the displacementcaused by the joint ∆ j , i.e. ∆ = ∆b + ∆c + ∆ j . Furthermore,∆ j is composed of the displacement caused by the shear


(a) Finite element mesh of the connection withinterior diaphragms.

(b) Finite element mesh of the interiordiaphragms.

(c) Finite element mesh of the connection withexterior diaphragms.

(d) Close-up of finite element mesh of theexterior diaphragms.

(e) Finite element mesh of the connection withanchored studs.

(f) Close-up of finite element mesh of theanchored studs.

Fig. 5. Typical finite element mesh.

(a) CFRTJ-3(ID). (b) CFRTJ-4 (ED). (c) CFRTJ-11 (AS).

Fig. 6. Comparison of experimental and numerical P–∆ curves under monotonic loading. (P = load applied to the beam end; ∆ = displacement at the beam end;ID represents the connections with interior diaphragms; ED represents the connections with exterior diaphragms, and AS represents the connections with anchoredstuds.)

Table 1Comparison of experimental and numerical results under monotonic loading

Specimen Loading direction Yield strength (kN) Ultimate strength (kN)Pym,FEA Py,e Pym,FEA/Py,e Pum,FEA Pu,e Pum,FEA/Py,e

CFRTJ-3 (ID)+ 148.2 145.7 1.02 185.5 167.7 1.11− −166.7 −153.0 1.09 −198.0 −169.1 1.17

CFRTJ-4 (ED)+ 160.3 160.3 1.00 194.3 187.6 1.04− −152.5 −179.6 0.85 −181.1 −194.3 0.93

CFRTJ-11 (AS)+ 114.5 75.2 1.52 122.5 90.7 1.35− −100.9 −108.5 0.93 −125.1 −115.7 1.08

Average 1.07 1.11Standard deviation 0.217 0.129

Note: Pym,F E A and Pum,F E A are the yield strength and ultimate strength under monotonic loading from finite element prediction, respectively; and Py,e and Pu,eare the yield strength and ultimate strength from tests, respectively.



Fig. 7. Comparison of experimental and numerical Q j –γ j skeleton curves under monotonic loading.(Qj = shear force of the panel zone, and γ j = sheardeformation of the panel zone.)

Fig. 8. Typical numerical distribution of stress and deformation of the connections.

deformation ∆s and the displacement caused by the non-sheardeformation ∆ns , i.e., ∆ j = ∆s + ∆ns .

The stress distribution of the interior diaphragm in CFRTJ-3at the ultimate state is shown in Fig. 8(a). It can be seen thatthe tension yield line is in the diagonal 45◦ direction, whichcorresponds to the conclusions of Yu [12]. As a result, thetension capacity of the interior diaphragms can be analyzedbased on the diagonal 45◦ yield line mechanism.

The ultimate deformation of CFRTJ-4 is shown in Fig. 8(b).The displacement is mainly caused by beam deformation andjoint shear deformation, while the portions of the columndeformation and the joint’s non-shear deformation are small.This is consistent with the test results [7]. The stress distributionof the steel structure in CFRTJ-4 at the ultimate state is shownin Fig. 8(c). It can be seen that the steel tube webs andpart of the bottom exterior diaphragms reached their ultimatestrength, which corresponds to the experimental observation.The principal stress orientation of the steel tube in the panelzone at the ultimate stage is in the two diagonal directions,

as shown in Fig. 8(d). The stress distribution of the concretecore in the panel zone of CFRTJ-4 at the elastic stage is shownin Fig. 8(e). The concrete formed a strut to resist the shear forceof the panel zone. Therefore, the strut model [5] can be used inthe shear capacity analysis of the connections.

The ultimate deformation of CFRTJ-11 is shown in Fig. 8(f).The displacement is mainly caused by the joint’s non-sheardeformation, while the other deformations are relatively small.At the end of the test the punching shear failure of the steel tubeflanges or fracture of the welds between the steel tube flangesand webs occurred, which is difficult to simulate in the finiteelement analysis.

3.2. Results under cyclic loading

Numerical P–∆ curves under cyclic loading are comparedwith the experimental curves, as shown in Fig. 9. The finiteelement analysis results showed fairly good agreement withthe experimental ones in terms of strength, deformation, andunloading stiffness, while the reloading stiffness did not show



Fig. 9. Comparison of experimental and numerical P–∆ curves under cyclic loading.

Table 2Comparison of experimental and numerical results under cyclic loading

Specimen Loading direction Yield strength (kN) Ultimate strength (kN)Pyc,FEA Py,e Pyc,FEA/Py,e Puc,FEA Pu,e Puc,FEA/Py,e

CFRTJ-3 (ID)+ 152.4 145.7 1.05 179.4 167.7 1.07− −146.2 −153.0 0.96 −170.1 −169.1 1.01

CFRTJ-4 (ED)+ 140.7 160.3 0.88 188.0 187.6 1.00− −143.8 −179.6 0.80 −172.6 −194.3 0.89

CFRTJ-11 (AS)+ 78.5 75.2 1.04 106.8 90.7 1.18− −142.7 −108.5 1.31 −142.7 −115.7 1.23

Average 1.01 1.06Standard deviation 0.163 0.115

Note: Pyc,F E A and Puc,F E A are the yield strength and ultimate strength under cyclic loading from finite element prediction, respectively.


Fig. 10. Comparison of experimental and numerical Q j –γ j curves under cyclic loading.

as good agreement. The degradation of reloading stiffness inthe finite element analysis is less obvious than that of the testcurves.

There is no pinching phenomenon in the numericalhysteresis curves for the connection with anchored studs for thereason that the punching shear failure mode of the steel tubeflanges cannot be simulated in the finite element model.

The comparisons of the yield strength and ultimate strengthof the connections are given in Table 2. Based on thecomparison of Tables 1 and 2, it is found that the strengthscalculated with numerical models under cyclic loading arecloser to the test values than those under monotonic loading.

The Q j –γ j curves obtained from numerical analyses andexperimental tests are compared in Fig. 10. It can be seen thatthe two curves agree well in terms of strength and unloadingstiffness, while the shear deformation and reloading stiffnessare significantly different. The fracture of welds and steel platesis difficult to simulate in the finite element analysis, and as aresult, the shear deformation calculated by ANSYS may not bereliable.

Distinct unsymmetrical behaviors have been shown inFigs. 9(c) and 10(c). As has been explained in the companionpaper [7], the negative flexural capacities are larger than thepositive ones for specimen CFRTJ-11 due to the existence ofthe steel angle. Therefore, the negative hysteresis loops showednearly linear characteristics when cyclic loads were applied.After the positive loads decreased to 85% of the positiveultimate strength of the connections, the cyclic loads wereterminated. Then, the static loads were applied to the beams inthe negative direction, and the monotonic curves were obtained.The unsymmetrical behaviors are caused by this factor.

Comparison of experimental results and numerical ones forEn–nh curves and Ec–nh curves are shown in Figs. 11 and12, respectively. The results of average energy dissipated perhemicycle calculated by the ANSYS model and obtained fromtests are summarized in Table 3. The dissipated energy fromthe finite element analysis is larger than that of the tests. Thereare two reasons for this difference: (1) the degradation ofreloading stiffness calculated in the numerical models is lessthan that of the tests, limited to the material properties/models



Fig. 11. Comparison of experimental and numerical results for En–nh curves. (En = energy dissipated per hemicycle of the hysteresis loops, and nh = number ofhemicycles)


Fig. 12. Comparison of experimental results and numerical ones for Ec–nh curves (Ec = cumulative dissipated energy of the specimens).

Table 3Comparison of experimental and numerical results for dissipated energy undercyclic loading

Specimen EFEA/kN mm Ee/kN mmEFEA

Ee

CFRTJ-3(ID) 1584.5 1371.7 1.16CFRTJ-4(ED) 5371.0 3674.9 1.46CFRTJ-11(AS) 980.2 794.9 1.23

Average 1.28Standard deviation 0.130

Note: EFEA = average energy dissipated per hemicycle from finite elementprediction, and Ee = average energy dissipated per hemicycle from tests.

available in ANSYS, and as a result, the numerical hysteresisloops are fuller than the experimental ones; and (2) since thepunching shear failure mode of the steel tube flanges is difficultto simulate in ANSYS, there is no pinching phenomenon inthe numerical hysteresis curves of CFRTJ-11. Although thereare some differences in terms of dissipated energy, the trendsare almost the same for the numerical results and experimentalones. Therefore, the finite element models can still be usedto provide some guidance in the design of the compositeconnections.

4. Parametric analyses

In order to investigate the effects of parameters onthe behavior of the connections, parametric analyses wereconducted based on the numerical model of the connection withexterior diaphragms under monotonic loading. The axial loadratio, width to thickness ratio, and dimensions of the exteriordiaphragms were varied in the finite element model, and theireffects were observed and are summarized below.

4.1. Axial load ratio

A comparison of numerical results for different axial loadratios n is shown in Fig. 13. It can be seen that the strength andstiffness are not significantly influenced by this parameter forthis type of connection under shear failure mode, which agreeswith the experimental results. With the increase of the axialload ratio, the strength of the connection decreases slightly, andthe displacement proportion of the joint shear deformation alsodrops slightly.

4.2. Width to thickness ratio

A comparison of numerical results for different tubethicknesses tc f is shown in Fig. 14. It can be seen thatthe strength and stiffness are significantly influenced by thisparameter. The strength of the connection increases with theincrease in the thickness of the steel tube; the stiffness ofthe connection also increases up to a certain extent, while thedisplacement due to the joint shear deformation drops.

4.3. Dimensions of exterior diaphragms

A comparison of numerical results for different dimensionsof the exterior diaphragms is shown in Fig. 15. It can be seenthat the strength and stiffness are not significantly influencedby this parameter for this type of connection under shear failuremode, which is consistent with the experimental results. Thisdemonstrates that the contribution of the exterior diaphragmsfor the shear capacity of the connections is negligible.

5. Conclusions and remarks

Based on the experimental results of connections composedof CFSSTCs and steel–concrete composite beams with interior


(a) P–∆ curves. (b) Q j –γ j curves.

Fig. 13. Comparison of numerical results for different axial load ratio (n = axial load ratio).


Fig. 14. Comparison of numerical results for different width to thickness ratio (tcf = thickness of the steel tube in mm).


Fig. 15. Comparison of numerical results for different dimensions of exterior diaphragms (bexu = width of the top exterior diaphragms in mm, and bexd = widthof the bottom exterior diaphragms in mm).

diaphragms, exterior diaphragms, or anchored studs, 3-Dnonlinear finite element models were established to analyzethe mechanical behavior of these three types of connectionusing ANSYS. Finite element analyses were conducted underboth monotonic loading and cyclic loading using appropriatematerial stress–strain relations and failure criteria. Parametricanalyses were conducted on the connections with exteriordiaphragms under monotonic loading to investigate theinfluences of axial load ratio, width to thickness ratio, anddimensions of exterior diaphragms. The main numerical resultscan be summarized as follows:

1. The load–displacement curves and shear force–deformationcurves of the finite element analyses are in good agreementwith those of the tests in terms of strength and unloadingstiffness, while the reloading stiffness and dissipated energyshow some difference.

2. There are three reasons for the differences between thefinite element analysis results and experimental ones: (1) thefracture of welds is difficult to simulate in the numerical

models, while during the tests this phenomenon wasobserved frequently; consequently, the shear deformationdeveloped in the tests is less than that predicted in thefinite element analyses; (2) the punching shear failuremode of the steel tube flanges is difficult to simulate inANSYS, and as a result, there is no pinching phenomenonin the numerical hysteresis loops of the connection withanchored studs; (3) the degradation of the reloading stiffnesspredicted in the numerical models is less than that of thetests due to the limitation of the material properties/modelsin ANSYS. Therefore, the dissipated energy of the finiteelement analysis is larger than that of the experiments.

3. Parametric analyses show that the strength and stiffness ofconnections are not significantly influenced by the axial loadratio and the dimensions of exterior diaphragms under shearfailure mode for the connections with exterior diaphragms,while they are significantly influenced by the tube thickness.

A few concluding remarks are offered here. Firstly, finiteelement analysis on the behaviors of connections composed


of CFSSTCs and steel–concrete composite beams are rarelyreported in the literature. This paper demonstrated that finiteelement analysis is useful not only for monotonic load analysisbut also for cyclic load analysis. Secondly, experimental andfinite element analysis results are the basis of theoretical studyon the connections. Based on the stress distribution of theconnections, the deformation of the connections, and the keyparameters of the connections, a theoretical study can beconducted. For example, the results that the concrete formeda strut to resist the shear force of the panel zone are found infinite element analysis, and based on this the strut model can beused in the shear capacity analysis of the connections. Thirdly,the brittle failure modes are difficult to simulate in the ANSYSprogram, especially under cyclic loading. Therefore, this paperdoes not discuss such failure modes. Fourthly, the failure modesof the connections are not the beam hinge failure modes, but thejoint shear failure modes or the local large deformation failuremodes. As a result the impacts of the concrete slabs are notobvious, and this is not discussed in this paper.

Acknowledgments

This research is a part of the research program on“Connections composed of CFSSTCs and steel–concretecomposite beams” at Tsinghua University, China. The writersgratefully acknowledge the financial support provided bythe National Natural Science Foundation of China (GrantNo. 50438020). Assistance for experimental studies from thestaff at Tsinghua University is also appreciated. The third writerappreciates the support from Louisiana State University.

References

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[2] Peng SW. Seismic resistant connections for concrete filled tube column-to-WF beam moment resisting frames. Ph.D. dissertation. Bethlehem(Pa): Lehigh Univ.; 2001.

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[4] Kang CH, Shin KJ, Oh YS, Moon TS. Hysteresis behavior of CFT columnto H-beam connections with external T-stiffeners and penetrated elements.Eng Struct 2001;23(9):1194–201.

[5] Zhou TH. Study on seismic behavior and load-carrying capacity ofconcrete-filled square tubular column to steel beam connection. Ph.D.dissertation. Xi’an (Shaanxi): Xi’an Univ. of Architecture & Technology;2004.

[6] Lin J. Experimental study on seismic performance of rectangularconcrete-filled steel tubular column to steel beam connections. Masterthesis. Fuzhou (Fujian): Fuzhou Univ.; 2004.

[7] Nie JG, Qin K, Cai CS. Seismic behavior of connections composedof CFSSTCs and steel–concrete composite beams-experimental study. JConstruct Steel Res 2008, in press [doi:10.1016/j.jcsr.2007.12.004].

[8] Willam KJ, Warnke EP. Constitutive model for the triaxial behaviorof concrete. In: Proceedings of international association for bridge andstructural engineering. Bergamo (Italy): ISMES; 1975.

[9] Han LH. Concrete-filled steel tubular structures. Beijing: Science Press;2000.

[10] Guo ZH. Theory of reinforced concrete. Beijing: Tsinghua Univ. Press;1999.

[11] Tang JR. Seismic resistance of joints in reinforced concrete frames.Nanjing (China): Southeast Univ. Press; 1989.

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http://dx.doi.org/doi:10.1016/j.jcsr.2007.12.004

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Seismic behavior of connections composed of CFSSTCs and