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Journal of Constructional Steel Research 64 (2008) 1283–1293

Contents lists available at ScienceDirect

Journal of Constructional Steel Research

journal homepage: www.elsevier.com/locate/jcsr

Tests of concrete-filled stainless steel tubular T-joints

Ran Feng, Ben Young ∗Department of Civil Engineering, The University of Hong Kong, Pokfulam Road, Hong Kong

a r t i c l e i n f o

Article history:

Received 12 November 2007Accepted 30 April 2008

Keywords:

Cold-formed steelStainless steelConcrete-filledHigh strengthRectangular hollow sectionSquare hollow sectionTubular structuresT-joints

a b s t r a c t

This paper describes a test program on a wide range of concrete-filledcold-formed stainless steel tubularT-joints fabricated from square hollow section (SHS) and rectangular hollow section (RHS) brace andchord members. A total of 27 tests wasperformed. Thechordmember of thetest specimen wasfilled withconcrete along its full length. Both high strength stainless steel (duplex and high strength austenitic) andnormal strength stainless steel (AISI 304) specimens filled with nominal concrete cylinder strength of 30 MPa were tested. The axial compression force was applied to the top end of the brace member, whichwas welded to the center of the chord member. Local buckling failure of brace member was the mainfailure mode observed during the tests. Hence, the axial compression force was then applied by means of a steel bearing plate to avoid failure of brace member. The failure modes of chord face failure and chordside wall failure as well as crushing of the concrete infill were observed. All the tests were performedby supporting the chord member of the specimen along its entire length to apply the pure concentratedforce without any bending moment. The test results were also compared with design rules for carbonsteel tubular structures, which is the only existing design guideline for concrete-filled tubular joints. Itis shown that the design strengths predicted by the current design rules are quite conservative for thetest specimens. It is also recommended that thecontribution of stainless steel tubes should be included inthe designrules since it has significant effects on theultimate bearing capacity of concrete-filled stainlesssteel tubular T-joints.

© 2008 Elsevier Ltd. All rights reserved.

1. Introduction

Cold-formed stainless steel tubular connections are being usedincreasingly for architectural and structural purposes in recentyears. Typical application of stainless steel in tubular constructionincludes pedestrian bridges. Connection reinforcement is widelyused when a truss connection has an inadequate resistance andits primary hollow section members cannot be changed. Thecommonly used method of strengthening tubular connections isto weld a stiffening plate to the exterior of the chord member. Oneof the disadvantages of this form of connection reinforcement isthe resulting structure may lose its aesthetic appearance due to thewelded stiffening plate. Another less visible and novel alternativefor certain connection types is to fill the hollow section withconcrete or grout. For short span trusses, the chord members of the tubular joints can be filled with concrete along the full lengthof the chord to improve the member capacity as well as the fireresistance. For long span trusses, only some parts of the chordmembers, especially in the vicinity of critical connections, needto be filled with concrete to increase the connection strengths.Concrete filling, instead of adding stiffening plates to the exterior

∗ Corresponding author. Tel.: +852 2859 2674; fax: +852 2559 5337.E-mail address: [email protected] (B. Young).

of tubular connections is particularly appealing for architecturallyexposed steelwork.

Numerous research studies have been conducted on concrete-filled hollow section carbon steel members as beams, columns andbeam–columns. However, little research has been carried out onconcrete-filled tubular connections. An experimental investigationof both grouted and ungrouted tubular T-joints fabricated fromcircular hollow section was carried out by Tebbett et al. [1].The ultimate strength of tubular connections was shown to besignificantly improved for axial tension, axial compression and in-plane bending load cases when the connections were filled withconcrete. A punching shear failure of the chord face was observedin the joints without concrete filled. However, the punching shearfailure was prevented by filling with concrete in the joints. Anextensive test program on a range of concrete-filled rectangularhollow section (RHS) in the chord members of X-connections wasconducted by Packer and Fear [2].A total of14 RHS specimens wastested subjected to transverse compression load applied throughbearing plates to avoid failure of the RHS brace members. Thetests were performed on a RHS of size 177.8 × 127.0 × 4.78 mmmember that varied in the amount of concrete filling, loaded withdifferent bearing areas and had the tube oriented in differentdirections. Packer and Fear [2] proposed a conservative lower-bound method for estimating the strength of concrete-filled RHSunder transverse compression that ignore the contribution of the

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1284 R. Feng, B. Young / Journal of Constructional Steel Research 64 (2008) 1283–1293

Notation

A1 Bearing area over which the transverse load isapplied

A2 Dispersed bearing areab0 Overall width of chord memberb1 Overallwidthofbracememberorsteelbearingplate

COV Coefficient of variation f c Crushing strength of concrete by cylinder testh0 Overall depth of chord memberh1 OveralldepthofbracememberorsteelbearingplateLc Length of concrete in chord memberL0 Overall length of chord memberL1 Overall length of brace member or overall height of

steel bearing plateN Axial compression loadN f Failure loadN max Maximum test strength (peak load)N s Test serviceability strengthN 3%b0 Test strength at the deformation of 3%bo

N ∗1 Factored resistance of concrete-filled tubular joint

N ∗1n Nominal strength of concrete-filled tubular jointq Maximum bearing stress of concrete infillr 0 Inner corner radius of chord memberr 1 Inner corner radius of brace memberR1 Corner radius of steel bearing platet 0 Overall thickness of chord membert 1 Overall thickness of brace memberu Chord flange indentationv Chord web deflectionw Weld sizew Weld size for full width jointβ Brace or steel bearing plate to chord width ratio

(b1/b0)2γ Chord width to thickness ratio (b0/t 0)

θ 1 Inclined angle between brace and chord memberτ Brace to chord thickness ratio (t 1/t 0)φc Resistance factor for limit states design of concrete

in bearing

steel and determine the capacity by the bearing strength of theconcrete. In 1995, a large number of tests on a range of concrete-filled T-, X- and K-joints fabricated from square and rectangularhollow sections (SHS and RHS) were performed by Packer [3].Special attention wasgiven to the comparison between the full andpartial concrete filling of RHSin the chord membersof the X-joints.The tests were performed by using bearing plates to transmittransverse compression load to the hollow section chord members

to prevent the failure mode of squashing the compression bracemembers. It should be noted that the aforementioned testsconducted by Tebbett et al. [1], Packer and Fear [2] and Packer [3]were focused on concrete-filled carbon steel tubular joints.

Design rules for concrete-filled tubular joints are available inthe Comité International pour le Développement et l’Étude de laConstructionTubulaire (CIDECT) Monograph No. 6 [4]. Currently,itis the only existing design guideline that can be used for concrete-filled tubular joints. The design procedure given in the CIDECT isbased on the studies of Packer and Fear [2] and Packer [3], which isa conservative lower-bound method by ignoring the contributionof the steel.

This paper focuses on the strength of concrete-filled stainlesssteel tubular T-joints fabricated from square and rectangular

hollow sections. Both high strength and normal strength stainlesssteel specimens were tested. In this study, the design guidelines

given by Packer [3] for concrete-filled carbon steel tubular T-jointswere used for concrete-filled stainless steel tubular T-joints. Theflange indentation andweb deflection of chord membersof the testspecimens as well as the observed failure modes were reported inthis paper. The joint deformations under service loads were alsoexamined in this study.

2. Experimental investigation

2.1. General

The test strengths of concrete-filled tubular T-joints are mainlybased on the failure modes. When local buckling failure of bracemembers governs, the concrete infill in the chord members couldnot contribute to the improvement of the strengths of T-joints.When the test specimens mainly failed by plastification of chordmembers as well as crushing of the concrete infill, the strengths of concrete-filled tubular T-joints are dependent mainly on: (1) theratio (β) of brace width to chord width (b1/b0); (2) the bearingarea ( A1) over which the transverse load was applied; (3) thedispersed bearing area ( A2); and (4) the crushing strength ( f c ) of

concrete infill. The concrete infill in the chord member plays animportant role for the significant improvement of failure load of tubular connections.

Laboratory tests were conducted to investigate the structuralbehaviour of concrete-filled stainless steel SHS and RHS T-joints.The concrete was filled in the chord member along its full lengthfor most of the test specimens. Axial compression force was firstlyapplied to the top end of the brace member, which was weldedto the center of the chord member and perpendicular to thechord. The concrete-filled tubular T-joints have non-dimensionalparameters of β ranging from 0.5 to 1.0, τ from 0.5 to 2.0, and 2γ from 10 to 50. Local buckling failure of brace member was foundto be the main failure mode for most of the test specimens. Hence,the axial compression force was then applied through steel bearingplates to avoid failure of brace members. The bearing area ( A

1) of

steel plate was ranging from 1545 to 22 448 mm2. Most of the testspecimens were filled with a nominal concrete strength of 30 MPain the chord members. The chord members were supported by arigid base along its entire length to apply the pure concentratedforce without any bending moment.

2.2. Test specimens

2.2.1. Stainless steel tubular T-joints with brace member

The compression tests were performed on concrete-filled cold-formed stainless steel tubular T-joints of square and rectangularhollow sections. A total of 11 specimens was fabricated withbrace member fully welded at right angle to the center of the

continuous chord member. The square and rectangular hollowsections consisted of a large range of section sizes. For the chordmembers, the tubular hollow sections had nominal overall flangewidth (b0) ranged from 40 to 200 mm, nominal overall depth of the web (h0) from 40 to 160 mm, and nominal thickness ( t 0) from1.5 to 4.0 mm. For the brace members, the nominal overall flangewidth (b1) ranged from 40 to 150 mm, nominal overall depth of the web (h1) from 40 to 150 mm, and nominal thickness ( t 1) from1.5 to 3.0 mm. The measured values of the specimen dimensionsare shown in Table 1 using the nomenclature defined in Figs. 1 and2. The values of plate width, depth, thickness, and weld size arebased on the average measurements of all four sides and welds of the intersection for each specimen. The end of the brace memberof tubular joint loaded in axial compression was milled flat to an

accuracy of 0.02 mm to ensure full contact between the specimenand end plate of the testing machine. The length of brace member

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Table 1

Measured dimensions and test results of test specimens with brace members

Specimen Chord (mm) Brace (mm) Weld (mm) Failure mode Test strength (kN)

h0 b0 t 0 r 0 L0 h1 b1 t 1 r 1 L1 w w β N f

TD-C40 × 2F1-B40 × 2 40.0 40.4 1.98 2.0 241 39.9 40.3 1.94 2.0 98 4.9 8.2 1.00 C 236.1TD-C50 × 1.5F1-B40 × 2 50.2 50.0 1.54 1.5 291 40.0 40.3 1.95 2.0 98 5.3 – 0.81 B+D 184.2TD-C160 × 3F1-B40 × 2 160.3 80.6 2.89 6.0 843 40.0 40.2 1.97 2.0 99 6.4 – 0.50 C 241.1TD-C50

× 1.5F1-B50

× 1.5 50.2 50.1 1.54 1.5 300 49.9 50.2 1.56 1.5 122 4.1 5.9 1.00 C 162.0

TD-C140 × 3F1-B50 × 1.5 140.1 80.0 3.09 6.5 750 50.0 50.3 1.52 1.5 122 6.9 – 0.63 C 168.9TD-C160 × 3F1-B50 × 1.5 160.4 80.4 2.89 6.0 853 50.1 50.1 1.54 1.5 121 6.2 – 0.62 C 173.8TD-C140 × 3F1-B140 × 3 140.1 80.0 3.15 6.5 841 140.3 80.4 3.10 6.5 345 7.2 10.2 1.00 C 571.8TH-C110 × 4F1-B150 × 3 110.0 196.0 4.00 8.5 703 150.9 150.4 2.79 4.8 369 8.1 – 0.77 C 426.4TN-C40 × 4-B40 × 2F3 40.1 40.0 3.91 4.0 240 40.1 40.1 2.00 2.0 99 8.5 11.1 1.00 C 235.8TN-C100 × 4-B40 × 2F3 99.8 49.7 3.85 4.0 540 40.0 40.1 2.02 2.0 99 8.3 – 0.81 B+D 193.1TN-C100 × 4-B100 × 2F3 99.7 49.6 3.82 4.0 602 99.9 50.1 1.97 2.0 246 8.6 11.2 1.00 B+D 324.6

Note: A = Chord face failure; B = Chord side wall failure; C = Local buckling failure of brace; D = Crushing of concrete.

Fig. 1. Definition of symbols for test specimen with brace member.

Fig. 2. 3D view of test specimen with brace member.

(L1) was chosen as 2.5h1 to avoid the overall buckling of bracemember. The length of chord member (L0) was chosen as 5h0 + h1

to ensure that the stresses at the brace and chord intersection arenot affected by the ends of the chord.

According to the tests conducted by Feng and Young [ 5] onSHS and RHS stainless steel tubular T-joints without concrete infill,several specimens failed by local buckling in the brace member.Hence, the same specimens would also failed by local bucklingof the brace member if concrete is filled in the chord memberand having similar failure loads. Therefore, concrete was filled inthe brace member of these specimens (TN-C40 × 4-B40 × 2F3,

TN-C100 × 4-B40 × 2F3 and TN-C100 × 4-B100 × 2F3) and thechord member was not filled with concrete. It should be noted that

the rest of the specimens were filled with concrete in the chord

member only, but not the brace member.

2.2.2. Stainless steel tubular T-joints with steel bearing plate

Based on the study of concrete-filled compression-loaded T-connections carried out by Tebbett et al. [1], the ultimate strengthsof tubular joints were actually governed by the failure of thecompression brace without concrete infill, as expected with sucha rigid base for the brace member. This was also demonstrated inthis study for concrete-filled stainless steel tubular T-joints.

Local buckling failure of brace member does not really reflectthe true ultimate capacity of the tubular joints. Hence, localbuckling in the brace member was purposely prevented by usingsteel bearing plate to transmit transverse compression force to theconcrete-filled chord member. A total of 16 specimens was testedwith the axial compression loads applied through steel bearingplate rather than brace member. The ratio of bearing plate widthto chord width of the specimens (β) varied from 0.5 to 1.0 forobtaining different failure modes as chord face failure and chordside wall failure as well as crushing of the concrete infill. For thechord members, the tubular hollow sections had nominal overallflangewidth(b0) ranged from40 to200 mm,nominaloverall depthof the web (h0) from 40 to 200 mm, and nominal thickness (t 0)from 1.5 to 6.0 mm. For the steel bearing plates that simulated thebrace members, they were fabricated from high strength steel withheight (L1) of 40 mm, the nominal overall width (b1) ranged from40 to 150 mm, nominal overall depth (h1) from 40 to 200 mm, andnominal corner radius (R1) from 3.0 to 12.5 mm. The measureddimensions of specimens and steel bearing plates are shown inTable 2 using the nomenclature defined in Figs. 3 and 4.

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Fig. 3. Definition of symbols for test specimen with steel bearing plate.

Fig. 4. 3D view of test specimen with steel bearing plate.

Table 2

Measured dimensions and test results of test specimens with steel bearing plates

Specimen Chord (mm) Steel bearing plate (mm) Weld (mm) Failure mode Test strength (kN)

h0 b0 t 0 r 0 L0 h1 b1 R1 L1 w w β N f

TD-C40 × 2F1-P40 × 40 40.0 40.4 1.98 2.0 241 40 40 4.0 40 4.9 8.2 0.99 B + D 259.8TD-C140 × 3F2-P40 × 40 140.2 80.0 3.14 6.5 740 40 40 4.0 40 6.3 – 0.50 A + D 232.9TD-C160 × 3F1-P40 × 40 160.3 80.6 2.89 6.0 843 40 40 4.0 40 6.4 – 0.50 A + D 272.8TD-C50 × 1.5F1-P50 × 50 50.2 50.1 1.54 1.5 300 50 50 3.0 40 4.1 5.9 1.00 B + D 248.1TD-C140 × 3F1-P50 × 50 140.1 80.0 3.09 6.5 750 50 50 3.0 40 6.9 – 0.63 A + D 323.1TD-C160 × 3F1-P50 × 50 160.4 80.4 2.89 6.0 853 50 50 3.0 40 6.2 – 0.62 A + D 300.4TD-C140 × 3F1-P140 × 80 140.1 80.0 3.15 6.5 841 140 80 9.5 40 7.2 10.2 1.00 B + D 734.1TD-C160 × 3F3-P140 × 80 160.4 80.6 2.89 6.0 939 140 80 9.5 40 6.9 10.6 0.99 B + D 635.4TH-C110 × 4F1-P150 × 150 110.0 196.0 4.00 8.5 703 150 150 7.8 40 8.1 – 0.77 A + B + D 3322.2TH-C150 × 6F2-P150 × 150 150.0 150.1 5.88 6.0 900 150 150 12.0 40 10.8 14.8 1.00 B + D 2212.3TH-C200 × 4F2-P200 × 110 197.0 109.2 4.02 8.5 1204 200 110 12.5 40 8.5 15.8 1.00 B + D 1003.9TN-C40 × 2F2-P40 × 40 40.1 40.1 2.00 2.0 241 40 40 4.0 40 5.0 7.8 1.00 B + D 200.4TN-C80 × 2F3-P40 × 40 80.1 80.1 1.87 4.0 443 40 40 4.0 40 4.7 – 0.50 A + D 283.2TN-C100 × 2F3-P40 × 40 99.9 50.1 1.94 2.0 540 40 40 4.0 40 5.4 – 0.80 A + D 149.6TN-C40 × 4F2-P40 × 40 40.0 40.0 3.92 4.0 241 40 40 8.0 40 8.0 11.7 1.00 B + D 308.5TN-C40 × 4F2-P40 × 40-R 40.1 40.1 3.82 4.0 241 40 40 8.0 40 8.3 10.8 1.00 B + D 319.8

Note: A = Chord face failure; B = Chord side wall failure; C = Local buckling failure of brace; D = Crushing of concrete.

2.3. Specimen labeling


The specimens are labeled according to their joint con-figuration, steel types, cross-section dimensions of chord andbrace members and concrete infill. For example, the labels ‘TD-C160 × 3F1-B40 × 2’ and ‘TN-C100 × 4-B100 × 2F3’ define thefollowing concrete-filled tubular T-joints:

• The first letter ‘T’ indicates the T-joint specimens.• The second letters ‘D’ and ‘N’ indicate that the steel type of the

specimens are duplex and normal strength austenitic stainless

steel type AISI 304. If the letter is ‘H’, it refers to high strengthaustenitic stainless steel.

• The third letter ‘C’ refers to chord member and the followingexpressions ‘160 × 3’ and ‘100 × 4’ indicate the cross-section dimensions of the chord members, which havingnominal overall depth of the webs (h0) of 160 and 100 mm,respectively as well as the wall thicknesses (t 0) of 3 and 4 mm,respectively. The overall flange width is purposely not shownfor simplification, and the dimension of the flange can be foundin Table 1. The notation ‘F1’ means concrete of the first batchwas filled in the chord member only. If the notation is notshown, then the chord member was not filled with concrete.

• The letter ‘B’ refers to brace member and the followingexpressions ‘40 × 2’ and ‘100 × 2’ indicate the cross-section

dimensions of the brace members, which having nominaloverall depth of the webs (h1) of 40 and 100 mm, respectively

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Table 3

Concrete mix design

Nominal concretestrength (MPa)

Water/cementratio

Mix proportions (to the weight of cement)

Cement Water Fineaggregate

Coarseaggregate

30 0.66 1.0 0.66 2.0 4.0

as well as the wall thickness (t 1) of 2 mm for both specimens.Once again, the notation ‘F3’ means concrete of the third batchwas filled in the brace member only.


The specimens are labeled according to their joint configu-ration, steel types, cross-section dimensions of chord members,steel bearing plates and concrete infill. The label ‘TN-C40 × 4F2-P40 × 40-R’ defines the specimen with nearly the same meaningas the label system for concrete-filled tubular T-joints with bracemember. The only difference is the part ‘P40 × 40-R’, in whichthe letter ‘P’ refers to steel bearing plate and the following expres-sion ‘40

× 40’ indicates the cross-section dimensions of the steel

bearing plate, which having nominal overall depth (h1) and overallwidth (b1) of 40 mm. If a test is repeated, then the letter ‘R’ indi-cates the repeated test.

2.4. Material properties of cold-formed stainless steel tubes

The specimens were cold-rolled from austenitic stainless steeltype AISI 304 (EN 1.4301), high strength austenitic (HSA) andduplex (EN 1.4462) stainless steel sheets. The stainless steel typeAISI 304 is considered as normal strength material, whereas theHSA and duplex are considered as high strength material. In thisstudy, the stainless steel tube specimens were obtained fromthe same batch of specimens conducted by Zhou and Young [6]for flexural members. The material properties of the stainless

steel tubes were determined by tensile coupon tests. The duplexstainless steel tubes are 40 × 40 × 2, 50 × 50 × 1.5, 140 × 80 × 3and 160 × 80 × 3 having the measured 0.2% tensile proof stress of 707, 622, 486 and 536 MPa, respectively; the highstrength austenitic (HSA) stainless steel tubes are 150 × 150 × 3,150 × 150 × 6 and 200 × 110 × 4 having the measured 0.2%tensile proof stress of 448, 497 and 503 MPa, respectively; thenormal strength stainless steel (AISI 304) tubes are 40 × 40 × 2,40 × 40 × 4, 80 × 80 × 2, 100 × 50 × 2 and 100 × 50 × 4 havingthemeasured0.2%tensileproofstressof447,565,398,320and378MPa, respectively.The tensilecoupon tests aredetailedin Zhou andYoung [6].

2.5. Material properties of concrete

Thematerialproperties of concrete were determinedfrom com-pressive concrete cylinder tests. The standard concrete cylinderhas nominal diameter of 150 mm and height of 300 mm. Theconcrete cylinders were produced using commercially availablematerials with normal mixing and curing techniques. In this study,a relatively low strength concrete was deliberately used. Theconcrete mix design for the expected nominal concrete cylinderstrength of 30 MPa is shown in Table 3. A total of 18 concrete cylin-der tests was conducted according to the test procedures in theAmerican Specification [7] for compressive concrete cylinder test-ing. The tests include three batches of concrete cylinders havingthe same concrete mix. The material properties of the concrete aresummarized in Table 4 that includes the measured concrete cylin-

der strength ( f c ) and the corresponding coefficients of variation(COV) for the three batches of concrete cylinders, respectively.

Table 4

Measured concrete cylinder strengths

Nominalconcretestrength (MPa)

Batch Measuredconcrete cylinderstrength f c (MPa)

Coefficientof variation(COV)

Number of concretecylinder tests

30 1 28.5 0.020 630 2 29.9 0.043 630 3 27.5 0.073 6

It is known that the concrete strength will be enhanced by theconfinement of the tube side walls. Hence, the observed maximumbearing stress (q) of the concrete is always greater than its crushingstrength ( f c ) determined by the concrete cylinder test. Studiesof confined concrete in bearing have been carried out by manyresearchers. The ratio of the bearing stress to the crushing strengthof concrete (q/ f c ) was given in concrete design codes as

√ A2/ A1

with different upper limit, where A1 = bearing area over whichthe transverse load is applied, and A2 = dispersed bearing area.It was recommended by Packer and Fear [2] and Packer [3] thatthe dispersed bearing area ( A2) should be calculated based onlongitudinal load dispersion along the chord member at a slopeof 2:1 until the edge of the concrete is reached, which has been

adoptedintheCIDECT[ 4].Theupperlimitof √ A2/ A1 was proposedby the American concrete code [8] of 2.0, the European concretecode [9] of 3.0, whereas the CIDECT [4] has the upper limit of 3.3. The design guidelines given in the CIDECT were used for thecalculation of concrete bearing stress in this study, as shown inTable 5.

2.6. Procedure of welding

The welds connecting brace and chord members were designedaccording to the American Welding Society (AWS) D1.1/D1.1Mspecification [10] and laid using shielded metal arc welding. Theweld sizes (w and w ) in the test specimens are all greater thanthe larger value of 1.5t or 3 mm, where t is the thickness of

thinner part between brace and chord members. The 2.5, 3.25and 4.0 mm electrodes of type E2209-17 with nominal 0.2% proof stress, tensile strength, and elongation of 635 MPa, 830 MPa, and25%, respectively, were used for welding high strength stainlesssteel (duplex and high strength austenitic) specimens. The 2.0and 2.5 mm electrodes of type E308L-17 with nominal 0.2% proof stress, tensile strength, and elongation of 440 MPa, 570 MPa, and37%, respectively, were used for welding normal strength stainlesssteel (AISI 304) specimens. The electrodes are described in detailsin the AWS A5.11 specification [11]. All welds consisted of 2 to 3runs of welding to guarantee that failure of specimens occurred inthe brace or chord members rather than the welds. The measuredweld sizes w and w (full width joint) are shown in Tables 1 and 2.

2.7. Test rig and procedure


The schematic sketches of the test arrangement are shownin Fig. 5. Axial compression force was applied to the specimenby using a servo-controlled hydraulic machine. The upper endsupport was movable to allow tests to be conducted at variousspecimen dimensions. A special fixed-ended bearing was designedto simulate the pure axial compression test without any bendingmoment applied to the specimen. The special bearing wasconnected to the upper end support. The chord member of thetest specimen was rest on the bottom end plate, which connectedto the bottom support of the testing machine. This providedsupport to the specimen along its entire length. The special bearing

was restrained from rotations by using four vertical bolts. Hence,the special bearing became a fixed-ended bearing which was

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Table 5

Limit-state loads and failure modes of test specimens with steel bearing plates

Specimen Failuremode

Test strength Nominalstrength

Ultimatelimit state

Serviceabilitylimit state

Flangeindentation

Webdeflection

β N f (kN) q

f c N s (kN) N ∗1n (kN)

N f

N ∗1n

N sφc N ∗1n/1.5

u (mm) v (mm)

TD-C40 × 2F1-P40 × 40 0.99 B + D 259.8 2.24 110.3 102.0 2.55 2.70 1.1 1.2TD-C140

× 3F2-P40

× 40 0.50 A

+ D 232.9 3.88 171.5 185.7 1.25 2.31 2.4 1.4

TD-C160 × 3F1-P40 × 40 0.50 A + D 272.8 4.13 196.4 188.3 1.45 2.61 1.9 1.0TD-C50 × 1.5F1-P50 × 50 1.00 B + D 248.1 2.24 117.4 159.7 1.55 1.84 1.1 1.5TD-C140 × 3F1-P50 × 50 0.63 A + D 323.1 3.49 231.1 249.1 1.30 2.32 2.4 1.5TD-C160 × 3F1-P50 × 50 0.62 A + D 300.4 3.72 208.1 265.2 1.13 1.96 1.6 0.9TD-C140 × 3F1-P140 × 80 1.00 B + D 734.1 2.24 320.4 714.5 1.03 1.12 1.4 2.4TD-C160 × 3F3-P140 × 80 0.99 B + D 635.4 2.36 298.7 726.4 0.87 1.03 1.4 2.4TH-C110 × 4F1-P150 × 150 0.77 A + B + D 3322.2 1.98 1044.9 1272.7 2.61 2.05 5.9 4.8TH-C150 × 6F2-P150 × 150 1.00 B + D 2212.3 2.24 1330.7 1506.3 1.47 2.21 2.9 4.5TH-C200 × 4F2-P200 × 110 1.00 B + D 1003.9 2.22 524.8 1464.0 0.69 0.90 2.2 3.3TN-C40 × 2F2-P40 × 40 1.00 B + D 200.4 2.24 73.1 107.2 1.87 1.70 1.2 1.1TN-C80 × 2F3-P40 × 40 0.50 A + D 283.2 3.00 150.2 131.8 2.15 2.85 2.4 1.1TN-C100 × 2F3-P40 × 40 0.80 B + D 149.6 3.32 105.1 145.6 1.03 1.80 1.2 1.0TN-C40 × 4F2-P40 × 40 1.00 B + D 308.5 2.24 112.0 107.1 2.88 2.61 1.2 0.5TN-C40 × 4F2-P40 × 40-R 1.00 B + D 319.8 2.24 112.3 107.2 2.98 2.62 1.2 0.6

Mean 1.68 2.04COV 0.442 0.301

Note: A =

Chord face failure; B =

Chord side wall failure; C =

Local buckling failure of brace; D =

Crushing of concrete. φc =

0.6.

(a) End view.

(b) Elevation.

Fig. 5. Schematic sketch of compression tests applied by brace member.

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Fig. 6. Deformations of test specimen with brace member.

considered to be restrained against both minor and major axesrotations as well as warping. The pure axial compression force wasthen applied to the specimen.

Two displacement transducers were positioned on either sideof the brace member measuring the vertical deflections at thecenter of the connecting face of the chord. The transducers werepositioned 20 mm away from the faces of the brace member, as

shown in Fig. 6. The flange indentation (u) in the chord memberwas obtained from the average reading of these two transducers.It is clearly demonstrated in Fig. 6 that the maximum outwarddeflection (v) of the chord web does not occur at the mid-heightof the chord side wall. It may approximately appear near the two-thirds of the overall depth of thechord web (h0). The exact locationof the maximum deformation of the chord side wall cannot beeasily predicted, as it depends on the initial plate imperfection of the chord side wall. Hence, two displacement transducers werepositioned at the mid-height of the chord side wall to recordthe side wall deflection. The average of these readings was alsotaken as the chord web deflection (v), as shown in Fig. 6. Twoother displacement transducers were positioned diagonally on thebottom end plate to measure the axial shortening of the specimen.A 1000 kN capacity servo-controlled hydraulic testing machinewas used to apply axial compression force to the specimens.Displacement control was used to drive the hydraulic actuator at aconstant speed of 0.1 mm/min for all the concrete-filled tubular T-

joints. Most of test specimens were failed by local buckling of bracemembers, as shown in Fig. 7 for a photograph of the test setup of specimen TH-C110 × 4F1-B150 × 3.


The schematic sketches of the test arrangement are shown inFig. 8. The test setup for specimens with steel bearing plate issimilar to those specimens with brace member. The deformationof test specimens was measured. Due to the space limitation, fourdisplacement transducers were positioned on the four aluminiumangles, as shown in Fig. 8. Two other displacement transducers

were positioned diagonally on the bottom end plate to measurethe axial shortening of the specimen. Therefore, the flangeindentation (u) in the chord member was obtained from thesetransducers, as shown in Fig. 9. Similar to the test procedure forthe specimens with brace member, two displacement transducerswere positioned at the mid-height of the chord side wall to recordthe side wall deflection. The average of these readings was alsotaken as thechord web deflection (v),as shown in Fig. 9. A5000kNcapacity servo-controlled hydraulic testing machine was used toapplyaxialcompressionforcethroughthesteelbearingplatetothespecimen. Displacement control was used to drive the hydraulicactuator at a constant speed of 0.2 mm/min. The applied loads andreadings of displacement transducerswere recorded automaticallyat regular interval by using a data acquisition system.A photographof thetest setup of specimen TH-C150

×6F2-P150

×150is shown

in Fig. 10.

Fig. 7. Test setup of concrete-filled tubular T-joint of specimen TH-C110 × 4F1-B150 × 3.

2.8. Test strengths and failure modes

2.8.1. Definition of failure load

In this study, there are four types of failure modes for thecompression tests of concrete-filled stainless steel tubular T-joints,namely chord face failure (plastic failure of the chord face), chordside wall failure (chord wall bearing or local buckling under the

compression bracing member), localbuckling failure of brace(localbuckling of the compression bracing member) and crushing of concrete. According to the design rules in the CIDECT [ 4], thefailure modes of hollow section tubular joints without concreteinfill are defined based on the value of brace width to chordwidth ratio (β). This design guidance, however is only suitablefor the hollow section tubular joints without concrete infill, andit is not applicable to the concrete-filled tubular joints. Designrules in the CIDECT [4], which is the only existing design guidelinefor concrete-filled tubular joints, do not include any definition of failure modes for concrete-filled tubular joints.

The designof concrete-filledtubular T-joints is more likely to begoverned by the ultimate limit state rather than the serviceabilitylimit state, the determination of failure load proposed by Zhao [12]

forhollow section tubularT-jointswithout concrete infillcannot befully used for concrete-filled tubular joints. For the concrete-filled

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(a) End view.

(b) Elevation.

Fig. 8. Schematic sketch of compression tests applied by steel bearing plate.

Fig. 9. Deformations of test specimen with steel bearing plate.

tubular T-joints, if the maximum test strength (N max) ata deforma-tion smaller than 3%b0, the maximum test strength is consideredto be the failure load; if the maximum test strength (peak load) ata deformation larger than 3%b0, the test strength N 3%b0 at the de-formation of 3%b0 is considered to be the failure load.


The failure loads (N f ) of the test specimens and the observedfailure modes are shown in Table 1. All the specimens failed bylocal buckling of brace members, except for the specimens TD-C50 × 1.5F1-B40 × 2, TN-C100 × 4-B40 × 2F3 and TN-C100 × 4-B100 × 2F3 failed by chord side wall and crushing of concrete. It

is shown that the joint deformations (u, v) for most test specimensarequitesmallwhentheysuddenlyfailedbylocalbucklingofbrace

members. In contrast, the joint deformations (u, v) of the threespecimens TD-C50 × 1.5F1-B40 × 2, TN-C100 × 4-B40 × 2F3 andTN-C100 × 4-B100 × 2F3 are relatively large when they failed byplastification of chord members as well as crushing of the concreteinfill.


The failure loads (N f ) of the test specimens and the observedfailure modes are shown in Table 2. It was found that thedeformations corresponding to the maximum test strengths for allthe specimens are greater than the respective 3%b0, except for thespecimens TD-C160×3F1-P40×40, TD-C160× 3F1-P50× 50and

TN-C100 × 2F3-P40 × 40, indicating that the 3%b0 deformationlimit for the ultimate strength of hollow section carbon steel

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follows:

N ∗1 = (φc f c A1/ sin θ 1)

( A2/ A1) (1)

where φc is the resistance factor for concrete in bearing (φc =0.6 in the CIDECT [4]), f c is the crushing strength of concrete bycylinder tests, A1 is the bearing area over which the transverse loadis applied, A2 is the dispersed bearing area, and θ 1 is the inclined

angle between brace and chord members. For which, A1 = h1b1,where h1 and b1 aretheoveralldepthandwidthofthesteelbearingplate, respectively. It should be noted that the design rules givenby Packer and Fear [2], Packer [3] and CIDECT [4] for concrete-filled carbon steel tubular joints under compression are onlyapplicable to test specimens with steel bearing plate rather thantest specimens with brace member. Therefore, the comparison of test strengths with design strengths of test specimens was notcarried out in this paper forthe testspecimenswith brace member.

Based on the recommendation given by Packer [3] for theconcrete-filled tubular T-joints, the dispersed bearing area ( A2)

was calculated by dispersion of the bearing load at a slope of 2:1longitudinally along the chord member, and through the entiredepth of the chord rather than limited by the

√ ( A2/ A1). The

equation for calculating the dispersed bearing area ( A2) proposedby Packer [3] is as follows:

A2 = (h1/ sin θ 1 + 4h0)b1 (2)

where h1 and b1 aretheoveralldepthandwidthofthesteelbearingplate,respectively, h0 istheoveralldepthofthechordmember,andθ 1 is the inclined angle between brace and chord members. Therange of validity for the length of concrete in chord member (Lc )was also recommended by Packer [3] as Lc ≥ (h1/ sin θ 1) + 4h0.

3.3. Ultimate limit state

In this study, the design rules given by Packer [ 3] for concrete-filled carbon steel tubular T-joints are used for the design of

concrete-filled stainless steeltubular T-joints. The designstrengths(N ∗1 ) given by Packer [3] have already incorporated the resistancefactor (φc ) for limit states design of concrete in bearing, whichare the products of nominal strengths and resistance factor, whereφc = 0.6 is recommended in the CIDECT [4]. Hence, the nominalstrengths (N ∗1n) can be obtained by dividing the design strengths(N ∗1 ) by resistance factor (φc ) as follow:

N ∗1n = N ∗1 /φc = N ∗1 /0.6. (3)

The calculated nominal strengths (N ∗1n) and the ratio of teststrengths to nominal strengths (N f /N ∗1n) are shown in Table 5. If the value of the ratio is greater than unity, it indicates conservativedesign strength. It follows from the values as shown in Table 5that the nominal strengths are conservative for both high strengthand normal strength stainless steel concrete-filled tubular T-

joints, except for specimens TD-C160 × 3F3-P140 × 80 and TH-C200× 4F2-P200× 110.It is clearly shown that the contributionof stainless steel tubes should be included in the design rules since ithas significant effects on the ultimate bearing capacity of concrete-filled stainless steel tubular T-joints. Furthermore, the effect of theparameter (h0/t 0) on the confinement of concrete should also betaken into account in the design rules, where h0 and t 0 are theoverall depth and thickness of the chord member.

3.4. Serviceability limit state

The design of stainless steel structures is more likely to be

governed by the serviceability limit state compared to carbonsteel structures, because the loss of stiffness associated with the

low proportionality stress precipitates growth of deformationsat loads well below ultimate (Rasmussen and Young [15]). Itwas proposed in the CIDECT [4] recommendations for hollowsection tubular joints that carbon steel joint deformations underservice loads should be limited to 1% of the chord width (b0). Thisrecommendation was also adopted in this study for concrete-filledstainless steel tubular T-joints.

For test specimens failed by plastification of chord membersand crushing of the concrete infill, the test serviceability strengths(N s) at which the measured deflection (max{u, v}) equaled 1% of the chord width (b0) are shown in Table 5. The test serviceabilitystrengths were compared with the design serviceability strengthsdetermined by dividing the design strengths (φc N ∗1n) by 1.5,where the value of 1.5 is consistent with the recommendations of CIDECT [4], as well as the American Institute of Steel Construction(AISC) [16] of using a load factor of 1.5 on the design strengthin allowable stress design. The ratio of the test serviceabilitystrengths to the design serviceability strengths (N s/(φc N ∗1n/1.5))is shown in Table 5. It follows from the table that the valuesof the ratio for all specimens are much greater than unity,except for specimen TH-C200 × 4F2-P200 × 110, indicating thatthe serviceability limit state generally will not be reached. For

concrete-filled stainless steel tubular T-joints, the larger webslenderness ratio (h0/t 0) may cause faster growth of thechord webdeflection (v) under relatively small service loads.

3.5. Comparison of test strengths with design strengths

The failure loads (N f ) were compared with the nominalstrengths (N ∗1n)predictedusingthedesignrulesgivenbyPacker[ 3].In addition,the testserviceability strengths (N s) at the deformationof 1%b0 were also compared with the design serviceabilitystrengths predicted using the recommendations of CIDECT [4] aswell as the American Institute of Steel Construction (AISC) [16].Table 5 shows the comparison of the test strengths with thedesign strengths for the ultimate limit state and serviceability

limit state of concrete-filled stainless steel tubular T-joints. Thedesign strengths were calculated using the measured cross-sectiondimensions as shown in Table 2 and the measured materialproperties as summarized in Table 4.

It is shown that the design strengths predicted by the designrules given by Packer [3] are generally conservative for bothhigh strength and normal strength stainless steel concrete-filledtubular T-joints. For the ultimate limit state, the mean value of the tested-to-predicted strength ratio (N f /N ∗1n) is 1.68, with thecorresponding coefficient of variation (COV) of 0.442. For theserviceability limit state, the mean value of the tested-to-predictedstrength ratio (N s/(φc N ∗1n/1.5)) is 2.04, with the correspondingCOV of 0.301, as shown in Table 5.

4. Conclusions

An experimental investigation of concrete-filled cold-formedstainless steel tubular T-joints of square and rectangular hollowsections has been presented in this paper. The test specimenswere cold-rolled from high strength and normal strength stainlesssteel materials having different bearing area over which thetransverse load was applied and dispersed bearing area. The teststrengths, failure modes and load-deformation curves for all thetest specimens have been reported. The failure modes involvedchord face failure, chord side wall failure, local buckling failure of brace as well as crushing of the concrete infill.

The test strengths were compared with the design strengthscalculated using the proposed design rules given by Packer [3] for

concrete-filled carbon steel tubular T-joints. It is shown that thedesign strengths are generally conservative for both high strength

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and normal strength stainless steel concrete-filled tubular T-joints.It is demonstrated that the concrete-filled cold-formed stainlesssteel tubular T-joints of square and rectangular hollow sectionscan be designed using the design rules for concrete-filled carbonsteel tubular T-joints. Furthermore, it would not be necessary tocheck the deformations of concrete-filled tubular T-joints underserviceability limit state. It should be noted that the contribution of stainless steel tubes should be included in the design rules since ithassignificant effects on the ultimate bearing capacity of concrete-filled stainless steel tubular T-joints. Further research is requiredto propose accurate design equations for concrete-filled stainlesssteel tubular T-joints.

Acknowledgements

The authors are grateful to STALA Tube Finland for supplyingthe test specimens. The authors are also thankful to Mr. Wai-ManWong for his assistance in the experimental program as part of hisfinal year undergraduate research projectat TheUniversity of HongKong.

References

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[2] Packer JA, Fear CE. Concrete-filled rectangular hollow section X and Tconnections.In: Proceedings of the fourth internationalsymposiumon tubularstructures. 1991. p. 382–91.

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[12] Zhao XL. Deformation limit and ultimate strength of welded T-joints in cold-formed RHS sections. Journal of Constructional Steel Research 2000;53(2):149–65.

[13] Lu LH, de Winkel GD, Yu Y, Wardenier J. Deformation limit for the ultimatestrength of hollow section joints. In: Proceedings of the sixth international

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