WANG, W. et al. - Bidirectional seismic performance of steel beam to circular tubular column connections with outer diaphragm (2011).pdf

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EARTHQUAKE ENGINEERING AND STRUCTURAL DYNAMICSEarthquake Engng Struct. Dyn. 2011; 40:10631081Published online 22 November 2010 in Wiley Online Library (wileyonlinelibrary.com). DOI: 10.1002/eqe.1070

Bidirectional seismic performance of steel beam to circular tubularcolumn connections with outer diaphragm

Wei Wang1,2,3,,, Yiyi Chen1,2, Wanqi Li2 and Roberto T. Leon3

1State Key Laboratory of Disaster Reduction in Civil Engineering, Tongji University, Shanghai 200092,

Peoples Republic of China2Department of Structural Engineering, Tongji University, Shanghai 200092, Peoples Republic of China

3School of Civil and Environmental Engineering, Georgia Tech, Atlanta, GA 30332-0355, U.S.A.

SUMMARY

This paper presents an experimental investigation on the seismic behavior of H-beam to circular tubular

column connections stiffened by an outer ring diaphragm. An innovative three-dimensional (3D) connectionsubassembly testing system was first described. Specimens representative of two-dimensional (2D) interiorcolumns, 3D interior and exterior columns in a steel building frame were then tested to failure underunidirectional or bidirectional cyclic loads. Various specimen parameters are used to evaluate their effectson connection behavior. Test results indicate significantly different failure modes for 2D and 3D weakpanel connections, with panel shear buckling and local distortion of outer diaphragm occurring only for 3Dconnections. The weak beam connections unexceptionally exhibited final fracture at the junction betweendiaphragm and beam flange. In contrast with weak beam connections, weak panel connections demonstratedbetter seismic performance and ductility. As a result, a seismic design philosophy considering panel zoneyielding before beam flexural yielding is proposed. Based on experiment observations, small diaphragmwidth and simplified fillet welding are found to be feasible especially for weak beam connections, improvingarchitectural appearance and facilitating construction. Strength evaluations also suggest that current AIJdesign provisions may be appropriate when applied to panel zones in 3D connections. Copyright 2010John Wiley & Sons, Ltd.

Received 20 April 2009; Revised 23 August 2010; Accepted 24 August 2010

KEY WORDS: connections; cyclic tests; circular hollow sections; three-dimensional tests; seismic design

1. INTRODUCTION

Circular hollow sections have excellent properties in resisting compression, bending and torsion

in terms of loading in all directions, and their shape is aesthetically pleasing. Their use in modern

steel-framed structures is becoming more and more popular. These tube sections, when used as

columns, have to be connected to beams of H-sections. When an H-beam frames into a circular

tube column, the width of the H-beam flange is normally smaller than the diameter of the column.Such joints, when the beam is directly welded to the column without any stiffeners, have been

found to be very weak in terms of their stiffness and load-carrying capacity. This can be avoided

by welding continuity plates inside the column such that the continuity plates and beam flanges

are at the same levels. This method, however, is expensive and complicated for fabrication. An

alternative method is though to use diaphragm type of connection, which is fabricated by first

Correspondence to: Wei Wang, State Key Laboratory of Disaster Reduction in Civil Engineering, Tongji University,Shanghai 200092, Peoples Republic of China.

E-mail: [email protected]

Copyright 2010 John Wiley & Sons, Ltd.


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1064 W. WANG ET AL.

cutting the steel tube into three pieces and then welding them together with two diaphragms.

Accordingly, this type of connection requires a large amount of welding. Moreover, if the depths

of beams coming into a connection are different, the tube is separated into more layers than that

in an ordinary beam-to-column connection. This requires a greater amount of welding resulting

in increased possibility of defects. Therefore, researchers have proposed outer stiffeners for such

joints involving tubular columns and H-beams. The use of outer stiffeners is perhaps the most

efficient form of force transfer from both the structural and constructional point of view.Most of previous research work has been devoted to studies of the strength and behavior of

externally stiffened steel beam to concrete-filled tubular (CFT) column connections, and significant

efforts have been made to understand their structural performance. These representative investiga-

tions include that of Kang et al. [1], Nishiyama et al. [2], Azizinamini and Schneider [3], Fukumoto

and Morita [4], Wu et al. [5], Wang et al. [6] and Shin et al. [7], all of which involved experimental

studies to assess elastoplastic behavior from subassemblage tests. On the other hand, limited work

exists in the literature on the behavior of externally stiffened connection between tube column and

steel beam. Ting et al. [8] presented the results of finite element analysis of externally stiffened

box-column to I-beam connections with different types of stiffener. T-stiffeners were reported to

be the most efficient form in redistributing stresses and improving stiffness. Shanmugam and Ting

[9] carried out experimental investigations on the ultimate load behavior of interior I-beam to

boxcolumn connections stiffened by T-sections under static and fluctuating loads. Experiment

results showed that these connections satisfy the basic criteria of strength, rotation capacity and

stiffness. Kumar and Rao [10] proposed a new and efficient connection between rectangular hollow

section beams and columns, which employed channel connectors welded to the column flange

and bolted to the beam to transfer beam flange forces into the column webs thereby avoiding

internal diaphragms in the column. The behavior of the connection was evaluated by cyclic tests

and non-linear finite element analysis. Failure was observed to occur at the beam net section away

from the column face in the case of channel connectors of high strength. Design guidelines were

then given for evaluating its ductility and energy dissipation capacity.

Although considerable amount of investigations have been carried out on outer-stiffened connec-

tion system of steel beams to vacant or CFT columns as stated above, no research is currently

available on the seismic behavior of three-dimensional (3D) circular tubular column to H-beam

connections with outer ring diaphragm under the severe earthquake. Because of the lack of test

evidences, the seismic design criteria and correlative detailing requirements of this type of connec-tion remain unclear and require further investigation. Therefore, this paper presents an experimental

investigation of circular tube column to steel beam subassemblies with outer ring diaphragms,

including three two-dimensional (2D) interior subassemblies, five 3D interior subassemblies and

one 3D exterior subassembly. An innovative spatial testing system for beamcolumn connections

has been developed. Test results are presented and discussed on the hysteretic behavior of the

connections subjected to unidirectional and bidirectional cyclic lateral loadings. The seismic perfor-

mance of the connections is evaluated in terms of strength, ductility and energy dissipation. The

design implications are proposed based on the comparison of the effects of various test parameters.

The work in this paper provides a basis for further development of an analytical model, which

will be described in another paper, and will help to establish a more reasonable seismic design

approach of this type of connection.

2. EXPERIMENTAL PROGRAM

2.1. Design of test specimens

The experimental program consisted of nine specimens to investigate the seismic behavior of the

steel beam-to-tubular column moment connections with outer diaphragm. Figure 1 shows the details

of the connection, where dc and tc are the diameter and thickness of the circular tube column,

respectively; bf, tf, hb and tw are the width, flange thickness, overall depth and web thickness

of the H-beam, respectively; and hs and ts are the width and thickness of the outer diaphragm,

Copyright 2010 John Wiley & Sons, Ltd. Earthquake Engng Struct. Dyn. 2011; 40:10631081

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BIDIRECTIONAL SEISMIC PERFORMANCE 1065

H beam

h

R

t t

Outer diaphragm

RR

R

h

R

RR

R Outer diaphragm

h

R

Outer diaphragm

d

t

h

b b b

d

t

d

t

t t

h

t t

h

Circular tube column Circular tube column Circular tube column

H beamH beam

(a) (b) (c)

Figure 1. Connection details: (a) interior (2D); (b) interior (3D); and (c) exterior (3D).

respectively. The outer diaphragms were first welded around the tube column and then jointed to

steel beams by welded flange-bolted web connections. In order to mitigate the abrupt geometric

changes in load transferring from a beam to the diaphragm, the rounded edges of the diaphragm

plate were formed by making it tangent simultaneously to the borders of two adjacent orthogonal

beam flanges and an auxiliary circle, which was marked with a dashed line in Figure 1. The radius

of the auxiliary circle was the sum of column tube radius and outer diaphragm width, hs. Thus,

the radius of diaphragm plate edge, R1, can be determined by the following equation:

R1=ctg(22.5o) dc2 +hs

2

2 bf

(1)

The moment of the beam may be carried by the beam flanges in the form of a couple axial force,

Tf (see Figure 2). The axial forces from the beam are first resisted by the outer diaphragm. In this

case, the outer diaphragm plays an essential role in transferring the forces from the beam flanges

to the column tube. The key parameter for the diaphragm design is the width (hs) of the critical

section. By assuming that the force from the beam flange is entirely transferred through the outer

diaphragm, as shown in Figure 2, the following equilibrium expression can be obtained:

Tf=

2Td (2)

where Td is the internal axial force based on critical section of the diaphragm. If the thickness

of the diaphragm is set equal to that of beam flange and the diaphragm is required to yield afterbeam flange yields, a simple design criterion for outer diaphragm can be derived as follows:

hs0.7bf (3)

Table I summarizes the dimensions of test specimens for the experimental study presented

in this paper. The column height, H, was 3025 mm. The beam length, L , was 3600mm for

interior subassemblies and 1800 mm for exterior subassemblies. Three of these specimens were 2D

subassemblies (C1C3), to which a reversed cyclic lateral load and column axial force were applied,

and the remaining specimens were 3D subassemblies (C4C9) and were subjected to reversed

cyclic lateral forces in two directions under constant column axial force. The subassemblies were


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Tf Tf

TdTd

Td Td

sh

Figure 2. Design assumption for outer diaphragm.

approximately one-half size scale models of the preliminary design of joints for a prototype 20-

story building. The parameters varied in the study including the loading direction of lateral force

(biaxial, uniaxial), the configuration of subassembly (interior, exterior), diameterthickness ratio

for the tubular column (dc/tc=39,25), the ratio of outer diaphragm width to beam flange width(hs/bf=0.71, 0.43, 0.41, 0.26, 0.23), and the weld type between column and outer diaphragm(completely penetrated welds, fillet welds from both sides). It should be noted that the specimens,

except for C1, were designed to have the hs/bf ratio less than 0.7 predicted by Equation (3) in order

to investigate the potential feasibility of smaller width of outer ring diaphragm because this is moredesirable for architectural appearance in practice. Moreover, the specimens with the weak panel

configuration (i.e. Specimens C1, C2, C3, C4 and C5) were designed to develop yielding primarily

in the panel zone to examine their failure modes and inelastic behavior. The shear strength of

the panel zone of the circular columns was computed according to the AIJ recommendations for

2D connections [11]. By increasing the thickness of the steel tube column, the specimens with

the weak beam configuration (i.e. Specimens C6, C7, C8 and C9) were also designed to develop

yielding primarily in the beams to investigate the performance of the panel zone in the weak-

beamstrong-column system. The outer diaphragms were welded to the column using fillet welds

from both sides for Specimen C9, with each leg size 7 mm, and using complete joint penetration

(CJP) single-bevel-groove welds for all other specimens. The beam flanges were welded to the

outer diaphragms using CJP single-bevel-groove welds. For CJP welds, backing bars were used

and removed after welding. Gas metal arc welding with CO2 shielding was adopted to fabricate

the welded connections of test specimens. Welding electrodes designated as E50 with a specifiedminimum CVN toughness of 80 J at 20C were used. The material properties and thickness ofthe steel plates or tubes for the beam, column, and outer diaphragm are given in Table II.

2.2. Three-dimensional testing system for beamcolumn connections

Figure 3 shows the configuration and loading condition for the interior and exterior subassemblies.

A 3D testing system was designed in order to simulate the bidirectional seismic lateral loading.

In this test setup, as can be shown in Figure 4, the test specimen was idealized as pinned at

both the top and the bottom of the column. The pinned connections were achieved using 3D

spherical plain bearings. The column bottom bearing was fixed on the foundation. The horizontal

movement of the specimen at the top and the bottom was prevented by two orthogonal braces

attached to the L-shaped strong reaction wall, respectively. Two sets of servo hydraulic actuator

pairs were available for cyclic loading in this test, identified as JB and JS, respectively. JB actuator

pair, capable of applying maximum 1000kN compression or 500 kN tension, delivering an anti-

symmetrical vertical loading at west and east beam ends. The maximum stroke of JB is 250mm.JS, with a capacity of 500 kN compression or 300 kN tension, served as anti-symmetrical vertical

loading at north and south beam ends. The maximum stroke of JS is 300mm. As is known, fora 3D beamcolumn subassembly under cyclic loading, the flexural deformation in one plane will

result in the beam torsion in the other orthogonal plane. If it is constrained, additional twisting

moment will be induced on the beam. This can be avoided by designing the details of cyclic

loading apparatus at the beam ends as shown in Figure 5(a). The beam section was clamped by

two stiffened steel plates through four threaded rods. Spherical bearings were then set between


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TableI.Summaryofouter

ringdiaphragmbeam-to-columnsubassemblies(inmm).

Weldtype

between

Joint

Column

Beam

Column

Beam

columnand

Constant

Spec

type

(circulartube)

(Hshape)

length,

H

length,

L

hs

ts

hs/

bf

dc/tc

diaphrag

m

axialforce

C1

Interior(2D)

350

9

300

175

8

12

3025

3600

125

14

0.71

39

CJPW

0.27Npc

C2

Interior(2D)

350

9

300

175

8

12

3025

3600

75

14

0.43

39

CJPW

0.27Npc

C3

Interior(2D)

350

9

300

175

8

12

3025

3600

45

14

0.26

39

CJPW

0.27Npc

C4

Interior(3D)

350

9

300

175

8

12

3025

3600

75

14

0.43

39

CJPW

0.27Npc

C5

Interior(3D)

350

9

300

175

8

12

3025

3600

45

14

0.26

39

CJPW

0.27Npc

C6

Interior(3D)

350

14

220

110

8

12

3025

3600

45

12

0.41

25

CJPW

0.27Npc

C7

Exterior(3D)

350

14

220

110

8

12

3025

1800

45

12

0.41

25

CJPW

0.27Npc

C8

Interior(3D)

350

14

220

110

8

12

3025

3600

25

12

0.23

25

CJPW

0.27Npc

C9

Interior(3D)

350

14

220

110

8

12

3025

3600

25

12

0.23

25

BFW

0.27Npc

Note:Npcistheaxialyield

strengthofCHScolumns;CJPWmean

scompletejointpenetratedwelds;BFW

meansfilletweldsfrombothsides.


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1068 W. WANG ET AL.

Table II. Material properties.

Specimen Steel type t (mm) fy (N/mm2) fu (N/mm

2) Es (N/mm2) (%)

C1C5 CHS tube 9 464 632 2.05105 32Beam flange 12 389 537 2.05105 27Beam web 8 509 601 2.07

105 24

Outer diaphragm 14 420 550 2.06105 28C6C9 CHS tube 14 348 558 2.09105 30

Beam flange 12 422 550 2.07105 27Beam web 8 328 454 1.99105 26

Outer diaphragm 12 422 550 2.07105 27

West

WEP

N

EastWEP

L

H West East

South

North North

East

NN

WEP

WEP

NSP

NSP

WEP

NSP

L

H

L

H

(a) (b) (c)

Figure 3. Subassembly configurations and loading conditions: (a) interior (2D);(b) interior (3D); and (c) exterior (3D).

Figure 4. Overview of test setup.

steel plates and top or bottom flange of the beam. A PTFE plate with friction coefficient of 0.03

was used for contacting surface of the spherical bearing. The basic idea behind this detailing is toalleviate the friction due to the compression between loading apparatus and the beam. As a result,

the beam twisted along its axis more freely in the test (Figure 5(b)). This approach was taken so

that all the bearings could be reused. In addition, a load cell was mounted between the actuator

and the loading apparatus to monitor the actual loading value. During the test, the beam tips were

braced laterally to prevent excessive out-of-plane displacements.

2.3. Loading procedure

At the beginning, a constant axial compression force equal to 0.27Npc was applied on the top

of the specimen by a hydraulic jack and maintained throughout the test. This axial load level of


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Beam

Threaded rod

Spherical bearing

Spherical bearing

PTFE plateBeam

(a) (b)

Figure 5. Details of cyclic loading apparatus at the beam ends: (a) before test and (b) during test.

the column is determined according to the maximum capacity of the vertical reaction frame. The

alternately repeated vertical loads, P , were then synchronously applied at beam ends by servo

actuators. For 2D interior subassemblies (C1C3), each load step started with the west beam going

up whereas the east beam went down (see Figure 3(a)). This uniaxial cyclic loading program

assumes that the seismic load is input within the westeast plane. For 3D interior subassemblies

(C4C6 and C8C9), each load step started with the west beam and north beam going up whereas

the east beam and south beam went down (see Figure 3(b)). For 3D exterior subassemblies (C7),

each load step started with the north beam going up whereas the east beam went down (see Figure

3(c)). Considering the maximum numbers and loading capacities of servo actuators that can be

offered by the laboratory, this biaxial cyclic loading program assumes that the seismic load is input

simultaneously in two directions, i.e. major loading direction along the westeast plane and minor

loading direction along the northsouth plane. The loading ratio of PWE in the westeast plane to

PNS in the northsouth plane is 1:0.75, which was realized by setting parallel connection of twosets of actuator pairs with different maximum loading capacities to the oil pump.

The loading protocol was based on a load history that consists of stepwise increasing deformation

cycles similar to the SAC loading protocol [12]. The deformation parameter used to determine the

loading history was the interstory drift angle, R, defined as the beam tip deflection divided by the

beam span. The interstory drift was applied in each principal direction, with the resultant at 37 having a magnitude equal to the drift in either principal direction multiplied by 1.25.

2.4. Measurements

The beam end displacement of the subassemblies, and the diagonal displacement of the shear

panels were measured by displacement transducers. Readings from the diagonal displacement

transducers at the panel zone were used to determine the shear deformation . Strain gauges were

installed to track the yielding process of the outer diaphragms, shear panels, beams and columns.

The vertical cyclic force at the beam end, P , was acquired through the load cell. As shown in

Figure 6, for each loading plane, the interstory drift, and the story shear force, Q, can be related

to and P by the following equations:

= 2H/L (for interior subassembly) (4)= H/L (for exterior subassembly) (5)


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L

H

L

P

P

Q

RR

R

Figure 6. Equivalent transformation from forcedisplacement relationship at beamend to story sheardrift relationship at column top.

Q = PL/H (6)

Thus, the vertical forcedisplacement relationships at two beam ends can be equivalently

expressed as the story shear force versus story drift angle relationship of one column, which makes

the test results for 2D and 3D subassemblies more intercomparable.

3. TEST RESULTS AND DISCUSSION

3.1. Yield mechanisms and failure modes

Ductile behavior are observed in all test subassemblies and the tests proceeded in a smooth and

controlled fashion. Readings from strain gauges showed that they differed in yielding sequence

for key components of the connection. For 2D interior specimens, C1 was observed to yield in

the shear panel first. Local yielding in the shear panel occurred at the center of the tube wall, and

then spread to the entire panel zone and the adjacent column section. Next, local yielding occurred

in the outer diaphragm. C2 was observed to first yield locally in the shear panel and the outer

diaphragm almost at the same time and then spread to the adjacent column section. C3 yielded inthe outer diaphragm first. After that the shear panel and the adjacent column were found to yield

simultaneously. No beam yielding was observed during the tests because the moment carrying

capacity of the beam was designed to be much stronger than the shear capacity of the panel zone,

as shown in Table III. It can be concluded that 2D specimens exhibited a strong-beamweak-panel

yield mechanism. For 3D specimens, they were divided into two different groups in terms of

yield mechanism. C4 and C5 showed similarities to C2 and C3 of 2D specimens, where yielding

is observed mainly in the panel zone, in the outer diaphragm and in the column near the shear

panel. However, with increased lateral displacement, C6, C7, C8 and C9 developed significant

yielding in the beam and the outer diaphragm and finally formed a plastic hinge. No obvious

shear deformation of the panel zone was observed during these tests. These specimens failed in a

weak-beamstrong-panel mode.

The failure modes varied for different test parameters. Generally, the failure mode mainly

depends on the tube wall thickness of the column, the width of the diaphragm and the loading

direction. Table III and Figure 7 show all the failure modes of connection specimens with 2D

configuration. The four observed failure modes are: excessive shear deformation of panel zone

(Figure 7(a)), occurred in C1 and C3; local buckling of column wall (Figure 7(b)), occurred in

C1, C2 and C3; weld crack between the column and the diaphragm (Figure 7(c)), occurred in C1;

and the doglegged deflection (Figure 7(d)), occurred in C2. Table III and Figure 8 present all the

failure modes of the connection specimens with 3D configuration. The five kinds of failure modes

identified are: shear buckling of panel zone (Figure 8(a)), occurring in C4 and C5; local buckling

of column wall (Figure 8(a)), also occurring in C4 and C5; local distortion of the outer diaphragm

(Figure 8(b)), occurring in C5; weld cracking between the column and the diaphragm caused by


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TableIII.Summaryoftestresults(inkN).

Calculatedvalue

Beam

Panelzone

Specimen

YieldstrengthQ

by

PlasticstrengthQbp

YieldstrengthQpy

UltimatestrengthQpp

Maxim

umloadQmax

Rmax

Qmax

/

Qbp

Failuremodes

C1

243

279

146

185

235

0.091

0.84

Figure7(a),(b),(c)

C2

222

255

146

185

218

0.067

0.85

Figure7(b),(d)

C3

211

243

146

185

195

0.069

0.80

Figure7(a),(b)

C4(WE)

222

255

146

185

185

0.072

C4(NS)

222

255

146

185

125

0.068

C4(37)

278

319

223

0.099

0.70

Figure8(a),(c)

C5(WE)

211

243

146

185

167

0.062

C5(NS)

211

243

146

185

117

0.043

C5(37)

264

304

204

0.075

0.67

Figure8(a),(b),(c)

C6(WE)

110

126

119

152

138

0.048

C6(NS)

110

126

119

152

105

0.024

C6(37)

138

158

173

0.054

1.09

Figure8(d)

C7(WE)

55

63

119

152

63

0.041

C7(NS)

55

63

119

152

47

0.019

C7(37)

69

79

79

0.045

1.00

Figure8(d)

C8(WE)

106

122

119

152

129

0.076

C8(NS)

106

122

119

152

101

0.033

C8(37)

133

153

164

0.083

1.07

Figure8(d)

C9(WE)

106

122

119

152

132

0.051

C9(NS)

106

122

119

152

95

0.027

C9(37)

133

153

163

0.058

1.07

Figure8(d)


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Figure 7. Typical failure modes of specimens with two-dimensional configuration: (a) excessive plasticshear deformation of panel zone; (b) local buckling of column wall; (c) weld crack between column and

diaphragm; and (d) doglegged deflection.

Figure 8. Typical failure modes of specimens with three-dimensional configuration: (a) local buckling ofcolumn wall and shear buckling of panel zone; (b) local distortion of outer diaphragm; (c) weld crackbetween column and diaphragm; and (d) fracture at the junction between diaphragm and beam flange.

local kinking (Figure 8(c)), occurring in C4 and C5; fracture at the junction between the diaphragm

and the beam flange (Figure 8(d)), occurring in C6, C7, C8 and C9. It can be clearly seen that the

specimens with thicker column walls led to yielding primarily in the beams so that the required

width of outer diaphragm can be greatly reduced. Differences in failure modes of C3 and C5 can

be attributed to loading directions. In Specimen C3, shear buckling of panel zone was prevented

by stiffening effect of two webs perpendicular to the loading direction. But for Specimen C5, shear

buckling occurred in the panel zone because no such effect existed in the resultant loading plane.

3.2. Seismic loading resistance

Resistances from all yield mechanisms and failure modes need to be compared and evaluated to

control connection behavior. However, resistances associated with these modes and mechanisms

may not be directly comparable because they occur at different locations. Therefore, the predicted

resistances must be adjusted for their location by equilibrium methods. In this study, the comparison

is conducted for forces at the top of the column. The test results are summarized in Table III, in

which the strengths are all expressed in terms of the story shear.

A method of calculating the strength of circular tube connection panels is available in the

AIJ literature [11], but the scope of application of this design formula does not include the 3D

subassemblies investigated in this study. The shear yield strength of a steel tube panel, Vpy, which

is used in allowable stress design against moderate earthquakes, is given by an equation that takes

into consideration the axial stress of the steel tube panel, using the von Mises yield criterion:

Vpy=(dc tc)tc

2

1n2 fy3

(7)

For the limit state design against a severe earthquake, the ultimate strength is considered to be

1.27 times the design yield shear strength.

Vpp=2(dctc)tc

1n2 fy3

(8)

where n is axial compression ratio.


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R (rad)

West-east plane

North-south plane

R (rad)

West-east plane

R (rad)

Q

(kN)

Q

(kN)

Q

(kN)

West-east planeNorth-south plane

R (rad)

Q

(kN)

West-east plane

North-south plane

R (rad)

Q

(kN)

West-east plane

R (rad)

West-east plane

R (rad)

West-east plane

North-south plane

R (rad)

Q

(kN)

West-east plane

North-south plane

R (rad)

Q

(kN)

Q

(kN)

Q

(kN)

West-east plane

North-south plane

(a) (b) (c)

(d) (e) (f)

(g) (h) (i)

Figure 9. Story shear versus story drift angle response for the connections: (a) C1; (b) C2; (c) C3;(d) C4; (e) C5; (f) C6; (g) C7; (h) C8; and (i) C9.

3.3. Hysteretic behavior

The story shear versus story drift relationships of the subassemblies are shown in Figure 9. For 2Dspecimens, Figure 9(a)(c) show the hysteretic curves for the uniaxial loading direction (westeast

plane). For 3D specimens, Figure 9(d)(i) show the hysteretic curves for the major loading direction

along the westeast plane in solid lines and minor direction along the north-south plane in dotted

lines, respectively. It is obvious that these specimens developed different extents of plasticity in two

directions, consistent to the loading ratio as planned. The curves of the subassemblies with weak

panel connection (C1C5) are of a fatter shape with full and stable hysteretic loops, indicating

large energy absorption. The cyclic responses were first represented by a steady increase of strength

up to the peak force. After maximum strength was reached, rapid loss of lateral strength was

not observed. Instead, the reduction in strength was slight and gradual with the increase of drift

amplitude. In comparison, the hysteretic curves of the subassemblies with weak beam connection

(C6C9) exhibited a noticeable shuttle-like shape with stable but not very full cyclic behavior.

After the specimens were monotonically subjected to final large deformation, a sudden drop in

the strength occurred, corresponding to the fracture at the junction between the diaphragm and

the beam flange (see Figure 8(d)). It can be concluded that the weak panel connections had better

energy dissipating capacity than the weak beam connections.

The shear forcedeformation (Vpp) responses in the major loading plane of the panel zonefor the specimens of both the weak panel and weak beam configurations are given in Figures 10

and 11, respectively. The panel zone of the specimens with the weak panel connection showed no

deterioration in shear resistance up to a deformation p over 0.04rad, and provided appreciable

deformation capacity. However, the panel zone of the specimens with the weak beam connection

only achieved a deformation p of 0.02 rad. Also plotted on the same figures are the predicted

shear yielding forces, calculated using Equations (4) and (5). For 2D specimens C1C3, it is clear


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1074 W. WANG ET AL.

(a) (b) (c)

(e)(d)

2500West-east plane

2000

1500

1000

500

0

500Vp(KN)

p(102 rad)

1000

1500

2000

2500Eq. 7

10 8 6 4 2 0 2 4 6 8 10

Eq. 8

2500

West-east planeCoupled shear

2000

1500

1000

500

0

500Vp(KN)

p(102 rad)

1000

1500

2000

2500Eq. 7

10 8 6 4 2 0 2 4 6 8 10

Eq. 8

2500West-east plane

Coupled shear2000

1500

1000

500

0

500Vp(KN)

p(102 rad)

1000

1500

2000

2500Eq. 7

10 8 6 4 2 0 2 4 6 8 10

Eq. 8

2500West-east plane

2000

1500

1000

500

0

500Vp(KN)

p(102 rad)

1000

1500

2000

2500Eq. 7

10 8 6 4 2 0 2 4 6 8 10

Eq. 8

2500West-east plane

2000

1500

1000

500

0

500Vp(KN)

p(102 rad)

1000

1500

2000

2500Eq. 7

10 8 6 4 2 0 2 4 6 8 10

Eq. 8

Figure 10. Panel zone sheardeformation results for weak panel zone specimens: (a) C1; (b) C2; (c) C3;(d) C4; and (e) C5.

(a)

2500


2000

1500

1000

500

0

500Vp(KN)

p(102 rad)

1000

1500

2000

2500Eq. 7

10 8 6 4 2 0 2 4 6 8 10

Eq. 8 Eq. 7 Eq. 8 Eq. 7 Eq. 8

(b)

2500


2000

1500

1000

500

0

500Vp(KN)

p(102 rad)

1000

1500

2000

250010 8 6 4 2 0 2 4 6 8 10

(c)

2500


2000

1500

1000

500

0

500Vp(KN)

p(102 rad)

1000

1500

2000

250010 8 6 4 2 0 2 4 6 8 10

Figure 11. Panel zone sheardeformation results for weak beam specimens: (a) C6; (b) C8; and (c) C9.

from Figure 10 that Equation (5) reasonably predicts the strength for connections with different

out diaphragm widths. Equation (4) closely predicts strength at the onset of inelastic panel zone

deformation for connections with different diaphragm widths. These equations, on the other hand,

seems not to be applicable for prediction of panel shear strength in the major loading plane for

3D specimens C4 and C5. If shear forces from two decoupled orthogonal directions parallel to

the beams are coupled and plotted against shear deformation in the major plane, as shown in gray

lines of Figures 10(d)(e) and 11, then Equations (4) and (5) seem to be on the conservative side

and still acceptable for predicting the strengths of tubular panel zone.

Table IV summarizes the ratios of the story drift angle caused by panel zone shear deformation,

Rp, to the total story drift angle R for each specimen when the maximum story drift was reached.

From the table, the percentages of drift angle contributed by the panel zone are 7090% for

Specimens C1C5 and 3846% for Specimens C6C9. It shows that the beam contributed themost story drift to the structural system with thick-wall column to beam connections, whereas

the panel zone contributed the most drift to the structural system with thin-wall column to beam

connections. This undoubtedly led to different energy dissipating mechanisms.

3.4. Ductility evaluation and energy dissipation

Table III lists the maximum story drift angle of each specimen. The total rotational capacity of all

specimens in both major loading plane and coupled loading plane exceeded 0.04rad. For special

moment resisting frames, the AISC Seismic Provisions [13] requires a total story drift capacity for

connections of 0.04 rad prior to degrading to 80% of the nominal beam capacity. Hence, based on


DOI: 10.1002/eqe


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Table IV. Drift angle percentage contributed by the panel zoneat maximum story drift.

Specimen Rp/R (%)

C1 80.6C2 71.1C3 90.7

C4 90.0C5 82.3C6 38.0C8 36.1C9 45.7

O

D

F EC R

Q

B

A

Figure 12. Illustration of equivalent damping coefficient he.

Table V. Equivalent damping coefficient of all specimens.

Specimen he

C1 0.437C2 0.399C3 0.440C4 0.366C5 0.445C6 0.291C7 0.303C8 0.291C9 0.278

a comparison of specimen response with AISC Seismic Provisions, all connections are observed

to have good ductility and are suitable for seismic resistant application.

The capacity of structural connections to dissipate energy when subjected to seismic loads

is as important as their strength or ductility in the evaluation process. The equivalent damping

coefficient he, as expressed in Equation (9), is a normalized value to evaluate the energy dissipation

of one hysteresis loop, as shown in Figure 12. The calculated he for the last completed loop of

the specimens is given in Table V. It should be noted that the area of the hysteresis loop stated in

Equation (6) has been calculated by integration and, therefore, represents the energy absorbed by

the specimen.

he=1

2

area(ABC+CDA)area(OBE+ODF) (9)


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1076 W. WANG ET AL.

-1.8

-1.2

-0.6

0.0

0.6

1.2

1.8

-6

R/Ry

Q/Qpy

C1

C2

C3

-4 -2 0 2 4 6

(a) R/Ry R/Ry

Q/Qpy

Q/Qpy

-1.8

-1.2

-0.6

0.0

0.6

1.2

1.8C4

C5

-1.8

-1.2

-0.6

0.0

0.6

1.2

1.8C6

C8

-6 -4 -2 0 2 4 6 -6 -4 -2 0 2 4 6(b) (c)

Figure 13. Effect of outer diaphragm width: (a) 2D weak panel connection; (b) 3D weak panel connection;and (c) 3D weak beam connection.

-1.8

-1.2

-0.6

0.0

0.6

1.2

1.8

-6 -4 -2 0 2 4 6

R/Ry

Q

/Qpy

C5

C6

Figure 14. Effect of column tube thickness.

It can be observed that the he values for C1C5 are obviously higher than C6C9 on average. In

this case, the thinner the column tube, the higher the percentage of energy dissipation contributedby the panel zone becomes. Therefore, the weak panel connections can be deemed to have better

energy dissipating capacity than the weak beam connections, having the same tendency as in the

previous section.

3.5. Effects of different parameters on the connection behavior

Figure 1317 shows the envelopes of the normalized story shear force and normalized story drift

angle, which are presented here to help compare the effects of different test variables. For 3D

specimens, only the envelopes for the major loading direction along the westeast plane are plotted

here. The story shear force, Q, is normalized by the AIJ design yield shear strength of a circular

tube joint panel, Qpy, expressed in terms of story shear. The measured story drift angle, R, is

normalized by the story drift angle at initial yield, Ry, representing the ductility index of the

specimens.

3.5.1. Effect of outer diaphragm width. Figure 13(a) shows a comparison of the behavior of the

2D weak panel subassemblies, C1, C2 and C3, with diaphragm widths of 125, 75 and 45 mm,

respectively. It was found that a larger width of the outer diaphragm may lead to larger strength

and better ductility. Figure 13(b) shows a comparison between 3D weak panel subassemblies, C4

and C5, with diaphragm widths of 75 and 45 mm, respectively. Figure 13(c) shows a comparison

between 3D weak beam subassemblies, C6 and C8, with diaphragm widths of 45 and 25 mm,

respectively. Similar effect was observed, and the difference was quite small. This can be attributed

to the fact that either panel zone failure or beam failure controlled connection behavior.


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-1.8

-1.2

-0.6

0.0

0.6

1.2

1.8

-6

R/Ry R/Ry

Q/Qpy

-1.8

-1.2

-0.6

0.0

0.6

1.2

1.8

Q/Qpy

C2

C4

C3

C5

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

(a) (b)

Figure 15. Effect of loading direction: (a) hs=75mm and (b) hs=45mm.

-1.8

-1.2

-0.6

0.0

0.6

1.2

1.8

-6 -4 -2 0 2 4 6

R/Ry

Q/Qpy

C6

C7

Figure 16. Effect of subassembly configuration.

-1.8

-1.2

-0.6

0.0

0.6

1.2

1.8

-6 -4 -2 0 2 4 6

R/Ry

Q/Qpy

C8

C9

Figure 17. Effect of welding type between column and outer diaphragm.

3.5.2. Effect of column tube thickness. Figure 14 shows the envelopes of C5 and C6, which had

different column wall thicknesses. C5, with smaller column tube thickness, shows more ductile

behavior than C6. The additional shear deformation capacity of the panel zone can be attributed

to this improved ductility.


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1078 W. WANG ET AL.

3.5.3. Effect of loading direction. Figure 15(a) and (b) shows the comparison of envelopes between

subassemblies C2 versus C4, and C3 versus C5. C4 and C5 were the 3D subassemblies loaded

simultaneously in the major and minor directions. It is observed that bidirectional loading may

reduce the strength in the decoupled loading plane. However, better ductility can also be achieved.

If story shear forces from two loading planes are coupled, the strengths of C4 and C5 would be

increased to be higher than that of C2 and C3, respectively.

3.5.4. Effect of subassembly configuration. Figure 16 shows the comparison of envelopes between

interior subassembly, C6, and exterior subassembly, C7. It was found that the ultimate lateral load

decreased significantly with exterior subassembly. However, the ductility was almost same for two

specimens.

3.5.5. Effect of welding type between column and outer diaphragm. Figure 17 shows the envelopes

of subassemblies C8 and C9, which were studied in order to understand the effect of welding type

between the column and the outer ring diaphragm. Good agreement can be observed, indicating

that fillet welding from both sides of the diaphragm, which is preferable from the of construc-

tion standpoint, can be used instead of complete penetration welds especially for weak beam

connections.

4. DESIGN IMPLICATIONS

4.1. Design philosophy considering effect of panel zone yielding

For structural design of circular tubular column-to-beam connections with outer diaphragms, the

effect of panel zone yielding on connection performance is often an issue of concern. Tests has

shown that panel zone yielding provided considerable ductility in inelastic cyclic deformation,

and recent building codes have increasingly emphasized utilizing this ductility in seismic design.

Figure 18 shows the maximum total story drift rotation achieved in the tests as a function of the

maximum shear force in the panel zone, Qpmax, normalized by Qpp from Equation (8), expressed

in terms of the story shear. Specimens with large Qpmax/Qpp ratios are those specimens that have

large amounts of panel zone shear yielding and strain hardening, and it can be observed that thesespecimens generally develop more plastic deformation than that of the specimens with less panel

zone yielding. Moreover, it can also be observed that the specimens, which do not develop the AIJ

panel zone shear capacity, also develop large story drift (greater than 0.04 rad).

In order to provide a comparison between initiation of the yield mechanism level for beam

flexure and panel zone yielding, Figure 19 plots the maximum total story drift rotation achieved

0.00

0.02

0.04

0.06

0.08

0.10

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8

Qmax/ Qpp

Rmax

Specimens which

develop AIJ panel

zone shear

capacity

Specimens which do

not develop AIJ panel

zone shear capacity

Specimens with limited

ductility

Specimens with relatively

large ductility

Figure 18. Total rotation as a function of normalized shear force.


DOI: 10.1002/eqe


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0.00

0.02

0.04

0.06

0.08

0.10

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8

Qby/ Qpy

Rmax

Specimens with initial

yield in panel zone

Specimens with initial

yield as flexure of beam

Limited rotational

capacity

Significant

rotational capacity

Figure 19. Total rotation as a function of relative beam flexure and panel zone yielding.

in the tests, as a function of the story shear force associated with initiation of flexural yielding,

Qby, divided by that associated with the panel zone shear yield force, Qpy. Specimens with aQby/Qpy ratio less than 1.0 develop flexural yielding of the beam before panel zone yielding

occurs. Specimens with ratios greater than 1.0 experience panel zone shear yielding before flexural

yielding occurs. Again the test data show that specimens yielding in beam flexure first, generally,

have smaller ductility than do specimens yielding first in panel zone shear. The largest rotational

capacities are achieved with specimens that have maximum Qby/Qpy ratio.

This leads to a design philosophy that the excessive panel zone yield deformation occurring

before beam flexural yielding will provide the greatest potential for connection ductility. Although

excessive large yield deformation of panel zone significantly increased demands for the weld

toughness or even led to the weld fracture between the diaphragm and the column, the ring

diaphragm stiffened connections in this program, unlike the conventional welded-flange-welded-

web connection, demonstrated excellent ductility hardly with any decrease in strength. The reason

can be attributed to the tying action formed by the outer diaphragm plate as whole. The design

equation based on this philosophy can be expressed as follows:

Mby

hb

L

Ldc

Hhb

H

(dc tc)tc

2

1n2 fy3

(10)

4.2. Design considerations for the width of outer diaphragm

Determining the width of critical section using Equation (3) may be conservative for the design of

outer ring diaphragm, since all specimens except C5 did not fail in this component. In fact, axial

forces from the beam flange were resisted not only by the diaphragm, but also by the diaphragm

and the column. Experimental stress analyses also have verified this force distribution mechanism.

It suggests that smaller width of outer ring diaphragm may be adopted to satisfy the requirement

by the architectural appearance in practice. However, further research should be carried out in the

future to provide reasonable width value for the outer diaphragm. In particular, as discussed in

Section 3.1, local distortion of outer diaphragm occurred in Specimen C5 but not in Specimen

C6, although they have same diaphragm width. This distinction implies that the width of outer

diaphragm can be designed even smaller in weak beam connections than in weak panel connections,

because beams usually fail before large plastic deformation develops throughout the diaphragm for

weak beam connection. In this case, it is also recommended that the junction between diaphragm

and beam flange should be set at a certain distance away from the critical section of the beam in

order to prevent possible premature fracture caused by the welding defects and high weld toughness

demand.


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1080 W. WANG ET AL.

5. CONCLUSIONS

An innovative 3D testing system for beamcolumn connections has been designed and the seismic

behavior of steel beam to circular hollow section column connections stiffened by outer diaphragms

was studied based on the cyclic loading tests on nine specimens with 2D or 3D configuration. The

main findings can be summarized as follows:

(1) Failure modes were mainly dependant on the column wall thickness. Specimens with a weak

panel connection (thin walls) and those with a weak beam connection (thick walls) failed

in significantly different modes.

(2) All connection subassemblies behaved in a ductile manner. However, in contrast with weak

beam connections, weak panel connections demonstrated better seismic performance and

ductility. A design philosophy considering panel zone yielding before beam flexural yielding

is proposed.

(3) Compared with unidirectional loading, bidirectional loading may reduce the connection

strength in the decoupled loading plane but increase the connection strength and ductility

in the coupled loading plane. Although the application scope of current tubular panel zone

provisions is intended to only include 2D connection subassemblies, it also gave reasonably

conservative estimates for coupled shear resistance of 3D connection subassemblies.

(4) Axial forces from the beam flange were resisted together by the diaphragm and the column.

It was therefore inferred that small width of outer ring diaphragm can be adopted to satisfy

architectural requirements. This is especially true for weak beam connections, since beams

usually fail before plastic deformation fully develops throughout the outer diaphragm.

(5) Fillet welding from both sides of the diaphragm, which is preferred by construction compa-

nies, may be adopted as a replacement of complete penetration welds in weak beam connec-

tions.

ACKNOWLEDGEMENTS

The presented work was supported by the Ministry of Science and Technology of China, Grant No.SLDRCE 09-B-02, National Natural Science Foundation of China, Grant No. 51008220 and Shanghai

Pujiang Program. Any opinions, findings, conclusions and recommendations expressed in this paper arethose of the writers and do not necessarily reflect the views of the sponsors. Technical help from theChina Northwest Building Design Research Institute is greatly appreciated.

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WANG, W. et al. - Bidirectional seismic performance of steel beam to circular tubular column connections with outer diaphragm (2011).pdf