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SLOTTED END CONNECTIONS TO HOLLOW SECTIONS

by

G. Martinez-Saucedo

and

J. A. Packer

Department of Civil Engineering, University of Toronto, Canada

FINAL REPORT TO CIDECT ON PROGRAMME 8G

CIDECT Report 8G-10/06

August 2006

ii

ABSTRACT

This Report deals with the structural behaviour and design of concentrically-aligned single

gusset plate welded connections to the ends of steel hollow section members. Such

connections are commonly found in diagonal brace members of steel framed buildings and also

in roof truss web-to-chord member connections. The types of sections considered are circular

hollow sections and elliptical hollow sections, with the plate either slotted and welded into the

tube or the tube welded into a slotted plate. In addition, the presence (or lack) of an open slot at

the end of a slotted tube connection - a fabrication method particularly favoured in North

America - is evaluated within the scope of this work.

Under quasi-static loading, the behaviour of the connection has been rigorously studied

under both axial tension and axial compression loadings, by both large-scale laboratory

experiments and numerical (finite element) analysis. In addition, an exhaustive review and

analysis of all prior international work in this field has been made. Non-linear finite element

models, validated for all 13 laboratory test specimens, formed the basis of an extensive

parametric study resulting in a further 891 "numerical tests" to supplement the data base of

experiments by the author and other international researchers. In tension the tube failure modes

of circumferential fracture (with or without the presence of shear lag) and tear out (or "block

shear" failure) were clearly identified by both experimental and numerical investigations and the

parameters influencing these limit states were thus clarified. As a result, new unified design

provisions for such connections in tension are presented, which are shown to be a significant

improvement over current international design provisions. In compression, the tube failure

mode of local buckling governed throughout the connection study and the influence of the shear

lag phenomenon - hitherto completely disregarded by all design provisions under compression

loading - has been highlighted. A new static design method for slotted end connections in

compression is hence advocated, which is shown to be applicable to circular, elliptical, square

and rectangular hollow sections. Guidance on the proportioning of the longitudinal fillet welds,

so that these do not govern the connection capacity, is also presented.

The above static design recommendations, which now more truly reflect the actual

connection performance, allow connections to be designed in a more efficient manner.

iii

TABLE OF CONTENTS

ABSTRACT ..............................................................................................................................iiTABLE OF CONTENTS ..........................................................................................................iii

NOTATION................................................................................................................................x

CHAPTER 1:INTRODUCTION............................................................................................. 1-1

1.1 Project overview......................................................................................................... 1-2

CHAPTER 2:LITERATURE REVIEW................................................................................... 2-1

2.1 The shear lag phenomenon ....................................................................................... 2-1

2.2 Tear-out failure ........................................................................................................... 2-4

2.3 International specifications......................................................................................... 2-6

2.4 Summary of Chapter 2............................................................................................. 2-10

CHAPTER 3:EXPERIMENTAL PROGRAM......................................................................... 3-1

3.1 Material properties ..................................................................................................... 3-1

3.1.1 Stub column tests.......................................................................................... 3-4

3.2 Test specimens and instrumentation.......................................................................... 3-6

3.3 Experimental test results ......................................................................................... 3-10

3.3.1 Slotted CHS connection - slot end not filled (type A) .................................. 3-10

3.3.2 Slotted CHS connection - slot end filled (with a weld return) (type B) ........ 3-13

3.3.3 Slotted EHS connection - slot end not filled (gusset plate oriented

to give a large eccentricity) ......................................................................... 3-15


to give small eccentricity)............................................................................ 3-18

3.3.5 Slotted gusset plate to tube connections in tension.................................... 3-21

3.3.5.1 Slotted gusset plate to CHS connection (type C)........................................ 3-21

3.3.5.2 Slotted gusset plate to EHS connection (gusset plate oriented to

give a large eccentricity) ............................................................................ 3-24

3.3.6 Connections under compression load......................................................... 3-27

3.3.6.1 Slotted CHS to gusset plate connection - slot end not filled ....................... 3-28

3.3.6.2 Slotted gusset plate to CHS connection ..................................................... 3-30

3.4 Summary of this experimental program ................................................................... 3-32

iv

CHAPTER 4:EVALUATION OF EXPERIMENTS AGAINST DESIGN PROVISIONS ......... 4-1

4.1 Experimental program by British Steel (1992) ........................................................... 4-2

4.2 Experimental program by Korol et al. (1994) ............................................................. 4-3

4.3 Experimental program by Zhao and Hancock (1995) ................................................ 4-4

4.4 Experimental program by Cheng et al. (1996) ........................................................... 4-7

4.5 Experimental program by Zhao et al. (1999) ............................................................. 4-8

4.6 Experimental program by Wilkinson et al. (2002) .................................................... 4-10

4.7 Experimental program by the Authors ..................................................................... 4-10

4.8 Experimental program by Ling (2005)...................................................................... 4-12


CHAPTER 5:FE MODELLING OF CONNECTIONS ........................................................... 5-1

5.1 Material properties ..................................................................................................... 5-1

5.2 Connection modelling ................................................................................................ 5-4

5.2.1 Element selection.......................................................................................... 5-6

5.2.2 Analysis considerations ............................................................................... 5-6

5.3 Evaluation of FE models against experimental results .............................................. 5-8

5.3.1 Slotted CHS connection - slot end not filled (Type A) ................................... 5-9

5.3.2 Slotted CHS connection - slot end filled (weld return) (Type B).................. 5-12




to give small eccentricity)............................................................................ 5-20

5.3.5 Slotted gusset plate to tube connections in tension.................................... 5-23

5.3.5.1 Slotted gusset plate to CHS connection (Type C)....................................... 5-23

5.3.5.2 Slotted gusset plate to EHS (gusset plate oriented


5.3.6 Connections under compression load......................................................... 5-30

5.3.6.1 Slotted CHS to gusset plate connection - slot end not filled ....................... 5-30

5.3.6.2 Slotted gusset plate to CHS connection ..................................................... 5-32


CHAPTER 6:PARAMETRIC FINITE ELEMENT ANALYSIS ............................................... 6-1

6.1 Parametric analysis results of slotted CHS connection - slot end not filled ............... 6-1

v

6.2 Parametric analysis results of slotted CHS connection -

slot end filled (weld return)......................................................................................... 6-7

6.3 Parametric analysis results of slotted EHS connection -

slot end not filled (gusset plate oriented to give a large eccentricity)....................... 6-11

6.4 Parametric analysis results of slotted EHS connection -

slot end not filled (gusset plate oriented to give small eccentricity) ......................... 6-14

6.5 Slotted gusset plate to tube connection in tension................................................... 6-17

6.5.1 Parametric analysis results of slotted gusset plate to

CHS connection .......................................................................................... 6-17


EHS connection (gusset plate oriented to give a large eccentricity)........... 6-23

6.6 Connections under compression load...................................................................... 6-29

6.6.1 Parametric analysis results of slotted CHS connection -

slot end not filled ......................................................................................... 6-29


CHS connection .......................................................................................... 6-32

6.7 Weld design ............................................................................................................. 6-35


CHAPTER 7: ANALYSIS OF FE AND EXPERIMENTAL RESULTS................................... 7-1

7.1 CHS connections in tension - CF failure .................................................................... 7-1

7.1.1 Shear lag equations suggested for CSA design provision format................. 7-1

7.1.1.1 Equation suggested for slotted CHS to gusset plate connections ................ 7-1

7.1.1.2 Equation suggested for slotted gusset plate to CHS connections

based on ultimate strength............................................................................ 7-3

7.1.1.3 Equation suggested for slotted gusset plate connections based on

deformation limit (0.03D)............................................................................... 7-5

7.1.2 Shear lag equations suggested for AISC design provision format................ 7-6

7.1.2.1 Equation suggested for slotted CHS to gusset plate connections ................ 7-7

7.1.2.2 Equation suggested for slotted gusset plate to CHS connections

based on ultimate strength............................................................................ 7-8


deformation limit (0.03D)............................................................................... 7-9

7.2 EHS connections in tension - CF failure .................................................................. 7-11

vi

7.2.1 Shear lag equations suggested for CSA design provision format............... 7-11

7.2.1.1 Equation suggested for slotted EHS to gusset plate connections .............. 7-11

7.2.1.2 Equation suggested for slotted gusset plate to EHS connections

based on ultimate strength.......................................................................... 7-12


deformation limit (0.03D2) ........................................................................... 7-14

7.2.2 Shear lag equations suggested for AISC design provision format.............. 7-15

7.2.2.1 Equations suggested for slotted EHS to gusset plate connections............. 7-15

7.2.2.2 Equations suggested for slotted gusset plate to EHS connections

based on ultimate strength.......................................................................... 7-17

7.2.2.3 Equation suggested for slotted gusset plate to EHS connections

based on deformation limit (0.03D2) ........................................................... 7-18

7.3 CHS and EHS connections in tension - TO failure .................................................. 7-19

7.4 CHS connections in compression ............................................................................ 7-29

7.4.1 Equation suggested for slotted CHS to gusset plate connections

(under compression loading) ...................................................................... 7-29

7.4.2 Equation suggested for slotted gusset plate connections

(under compression loading) ...................................................................... 7-30

7.5 Evaluation of recommended equations against experimental data ......................... 7-31

7.5.1 Experimental program by British Steel (1992) ............................................ 7-33

7.5.2 Experimental program by Korol (1994) ....................................................... 7-34

7.5.3 Experimental program by Cheng et al. (1996) ............................................ 7-34

7.5.4 Experimental program by the Authors ........................................................ 7-35

7.6 Derivation of reduction (resistance) factors for the recommended equations.......... 7-37

7.6.1 Reduction factors for CHS connections in tension - CF failure................... 7-37

7.6.1.1 Reduction factors for suggested equations for slotted

CHS connections (CSA design provision format) ....................................... 7-37

7.6.1.2 Reduction factors for suggested equations for slotted gusset plate to

CHS connections based on ultimate strength

(CSA design provision format) .................................................................... 7-38


CHS connections based on deformation limit


vii

7.6.1.4 Reduction factors for suggested equations for

slotted CHS connections (AISC design provision format)........................... 7-39


CHS connections based on ultimate strength

(AISC design provision format) ................................................................... 7-40


CHS connections based on deformation limit


7.6.2 Reduction factors for EHS connections in tension - CF failure ................... 7-41

7.6.2.1 Reduction factors for suggested equations for

slotted EHS connections (CSA design provision format)............................ 7-41


EHS connections based on ultimate strength



EHS connections based on deformation limit


7.6.2.4 Reduction factors for suggested equations for slotted EHS connections (AISC

design provision format).............................................................................. 7-42


EHS connections based on ultimate strength


7.6.2.6 Reduction factors for suggested equation for slotted gusset plate to

EHS connections based on deformation limit


7.6.3 Reduction factors for CHS and EHS connection in tension - TO failure ..... 7-44

7.6.3.1 Reduction factors for slotted CHS connections - TO failure ....................... 7-44

7.6.3.2 Reduction factors for slotted gusset plate to CHS connections -

TO failure .................................................................................................... 7-44

7.6.3.3 Reduction factors for slotted EHS connections - TO failure........................ 7-44

7.6.3.4 Reduction factors for slotted gusset plate to EHS connections -

viii

TO failure .................................................................................................... 7-45

7.6.4 Reduction factors for CHS connections in compression............................. 7-45

7.6.4.1 Reduction factors for slotted CHS connections in compression ................. 7-45

7.6.4.2 Reduction factors for slotted gusset plate to CHS connections

in compression............................................................................................ 7-46

7.7 Summary of Chapter 7- recommended static design methods................................ 7-46

7.7.1 Recommended static design method for CHS connections in tension ....... 7-47

7.7.2 Recommended static design method for EHS connections in tension ....... 7-49

7.7.3 Recommended static design method for CHS connections in compression7-50

CHAPTER 8:CONCLUSIONS AND RECOMMENDATIONS FOR FURTHER RESEARCH8-1

8.1 Overview.................................................................................................................... 8-1

8.2 Recommended static design methods...................................................................... 8-2

8.2.1 Recommended static design method for CHS connections in tension ......... 8-2

8.2.2 Recommended static design method for EHS connections in tension ......... 8-4

8.2.3 Recommended static design method for CHS connections in compression 8-5

8.3 Design recommendation for seismic applications...................................................... 8-6

8.4 Recommendations for further research ..................................................................... 8-6

CHAPTER 9:REFERENCES................................................................................................ 9-1

APPENDIX A: EXPERIMENTAL PROGRAM......................................................................A-1

A.1 Slotted end connections to CHS................................................................................A-1

A.2 Slotted end connection to EHS..................................................................................A-3

APPENDIX B: EVALUATION OF EXPERIMENTS ..............................................................B-1

B.1 Experimental program by British Steel (1992) ...........................................................B-1

B.2 Experimental program by Korol el al. (1994) .............................................................B-2

B.3 Experimental program by Zhao and Hancock (1995) ................................................B-3

B.4 Experimental program by Cheng et al. (1996) ...........................................................B-5

B.5 Experimental program by Zhao et al. (1999) .............................................................B-6

B.6 Experimental program by the Authors .......................................................................B-8

B.7 Experimental program by Ling (2005)........................................................................B-9

APPENDIX C: STRAIN READINGS ....................................................................................C-1

C.1 Connections under tension ........................................................................................C-1

ix

C.1.1 Slotted CHS connections - slot end not filled (A1 and A2).........................................C-1

C.1.2 Slotted CHS connections - slot end filled (weld return) (B1 and B2) .........................C-3

C.1.3 Slotted EHS connections - slot end not filled (E1 and E2).........................................C-5

C.1.4 Slotted EHS connections - slot end not filled (E5) .....................................................C-7

C.1.5 Slotted gusset plate to CHS connection (C1 and C2)................................................C-8

C.1.6 Slotted gusset plate to EHS connection (E3 and E4) ..............................................C-10

C.2 Connections under compression .............................................................................C-12

C.2.1 Slotted CHS to gusset plate connection - slot end not filled (A3C)..........................C-12

C.2.2 Slotted gusset plate to CHS connection (C3C)........................................................C-13

NOTATION

Ag = gross cross-sectional area of hollow section

Agt = gross area in tension for block failure

Agv = gross area in shear for block failure

al = weld leg length (size)

An = net cross-sectional area of hollow section

A'ne = effective net cross-sectional area of hollow section

Ant = net area in tension for block failure

Anv = net area in shear for block failure

Aw = area of effective weld throat

B = width of overlapped gusset plate

b = overall width of RHS and SHS, measured 90 degrees to the plane of the connection

CSC = compressive strength of stub column

CHS = Circular Hollow Section

D = outside diameter of CHS

D1 = larger dimension of EHS

D2 = smaller dimension of EHS

Davg = average between larger and smaller dimension of EHS

D/t = ratio between outside diameter and wall thickness of CHS

E = modulus of elasticity

EHS = Elliptical Hollow Section

Fy = yield tensile stress

Fu = ultimate tensile stress

h = overall height of RHS and SHS, measured in the plane of the connection

HAZ = Heat Affected Zone

HSS = Hollow Structural Section

x

K = effective length factor

LVDT = Linear Variable Differential Transformer

Lw = weld length

Lw/w = ratio between weld length and distance between welds

Lw/D = ratio between weld length and outside diameter of CHS

lsl = length of slot in hollow section

Nu = calculated connection strength according to design provisions

Nux = measured connection strength

NuFE = connection strength from FE analysis

NuFE-D= connection strength from FE analysis based on distortion limit

RHS = Rectangular Hollow Section

Rt = tension area mean stress correction factor

Rv = shear area mean stress correction factor

SHS = Square Hollow Section

t = wall thickness of CHS

tp = thickness of gusset plate

Tr = factored tensile resistance

tsl = width of slot in CHS

Tσ-Tε= uniaxial true stress - true strain curve

U = reduction coefficient for shear lag in net section fracture calculation

Ubs = reduction factor for non-uniform stress in block shear

Vr = factored shear resistance

VR = coefficient of variation

w = distance between the welds, measured around the perimeter of the CHS

wp = width of gusset plate

= eccentricity x

xi

= eccentricity reduced by half of flange-plate thickness ( = - tp /2)

/Lw = ratio between the eccentricity and weld length

/Lw = ratio between the reduced eccentricity and weld length

z = longitudinal distance between strain gauges

= safety index or “reliability index”

γM0 = Eurocode 3 partial safety factor when neither buckling phenomena nor ultimate

resistance in tension is under consideration (= 1.0)

γM2 = Eurocode 3 partial safety factor when ultimate resistance in tension is under

consideration (= 1.25)

εu = ultimate strain at rupture

εef = equivalent fracture strain

= mean actual-to-predicted ratio

= resistance factor

x' x

x

x'

β

ϕm

φ

xii

1-1

CHAPTER 1: INTRODUCTION

Circular hollow sections (CHS) have gained in popularity in recent years, particularly for

architecturally exposed structural steel. Architects appreciate the clear form of CHS as well as

their excellent structural properties in compression and torsion. In order to take full advantage of

these properties, the complete tube cross-section should ideally be engaged at the connection.

However, the feasibility of doing this is determined by the shape of the elements merging at the

connection, which may result in a complicated task for detailing and fabrication. As a result, the

use of a simplified connection detail will always be desirable whenever possible.

Gusset plate connections represent one of the easiest methods to connect CHS used as

web members in roof trusses and brace members in buildings. During the fabrication of these

connections, the gusset plate or the CHS can be slotted resulting in several possible fabrication

details. The application of either detail will depend on existing tolerances during the process of

fabrication and erection of the structure. Despite these connection details providing the simplest

manner for connecting CHS, it is important to recognize that an incorrect understanding of their

behaviour may result in their failure or an expensive conservative design. As a consequence of

only part of the CHS cross-section being connected, an uneven stress distribution around the

tube circumference always occurs during the load transfer at the connection. Shear lag (see

Figure 1.1) leads to stress peaks at the beginning of the weld which may result in connection

failure by a circumferential failure (CF) mode. Moreover, a tear-out (or “block shear”) failure

(TO) may also occur under tension loading.

Beginning of theweld

Figure 1.1 Shear lag in slotted CHS connection

SLOTTED END CONNECTIONS TO HOLLOW SECTIONS, CH1: INTRODUCCION

1-2

Despite both these limit states being addressed in current North American design

provisions (AISC 2005 and CSA 2001), it has been found that the predicted connection strength

(in the parameter range when CF is governing failure mode) will always differ as these two

design provisions use dissimilar methods to account for this phenomenon. Although it is

expected that these AISC and CSA design methods will always predict conservative connection

capacities when CF governs, it has been found that the number of studies (specifically in slotted

end connections to hollow sections) is limited to verify the accuracy and validity limits for each

method. Moreover, the model currently used in design provisions (AISC 2005, CEN 2005 and

CSA 2001) to account for TO failure, which was initially developed for bolted connections, lacks

studies verifying its accuracy and validity limits for these connection types. In a similar manner

to tension loading, an uneven stress distribution can be expected at the connection under

compression loading. However, it has been found that this phenomenon is completely

disregarded by design provisions, despite the fact that it may induce tube local buckling at the

beginning of the welds.

1.1 Project overview

This Report is directed to clarify the behaviour of slotted end connections fabricated with

CHS and Elliptical Hollow Sections (EHS), their possible failure mechanisms and the relation of

these failure modes to the connection geometrical dimensions, under tension and compression

loading. In order to verify the accuracy of models currently used by design provisions, these are

compared against available experimental data from previous studies and data from an

experimental program undertaken at the University of Toronto. Results from these comparisons

revealed the deficiency of these provisions to correctly predict the connection strength and

governing failure mechanisms. A further parametric analysis based on finite element models of

CHS and EHS connections has provided information on the behaviour of these connections and

also provided further evidence of the imprecision of current design provisions. Therefore, a new

comprehensive static design method is recommended here which also illustrates the possibility

of effectively diminishing the influence of shear lag in these connections.

SLOTTED END CONNECTIONS TO HOLLOW SECTIONS, CH1: INTRODUCCION

2-1

CHAPTER 2: LITERATURE REVIEW

The use of slotted end connections to hollow sections is very popular nowadays.

However, the design methods against the most frequent failure modes such as circumferential

tensile fracture (CF) of the HSS (see Figure 2.2) and tear-out (TO) failure along the weld (see

Figure 2.3), seem to still require further attention. During the load transfer from the tube to the

gusset plate, a nonuniform strain distribution takes place in the tube cross-section as the

unconnected material is less able to participate in the load transfer. This phenomenon, known

as Shear Lag, creates a high strain concentration at the weld region which eventually can

trigger the fracture of the tube material there. Moreover, the propagation of this crack (defining a

typical failure mode) and the connection strength are strongly influenced by the weld length

(Lw).

2.1 The shear lag phenomenon

Since the first model to account for the shear lag phenomenon was proposed by Chesson

and Munse (1963), it has been included in several design specifications. Initially it was applied

to riveted and bolted connections. Afterwards, the same model was utilized for the design of

welded connections. Even though this phenomenon has been studied extensively for open

structural sections, studies from Easterling and Giroux (1993) and Kirkham and Miller (2000)

Figure 2.2 Circumferential tensile fracture Figure 2.3 Tear-out failure

SLOTTED END CONNECTIONS TO HOLLOW SECTIONS, CH2: LITERATURE REVIEW

2-2

have revealed that existing design approaches are overly conservative and that further

research may be required. In addition, this model has been applied to tubular connections.

However, the research for these connection types is relatively recent and limited in scope.

To allow for shear lag on connections fabricated with Hollow Structural Sections (HSS),

Packer and Henderson (1992) proposed that the distance between the welds (w) be measured

along the developed perimeter of the HSS (see Figure 2.4). In addition, they also suggested an

efficiency coefficient for connections with Lw/w ratios less than unit. At this time, the use of small

ratios was not considered for CAN/CSA-S16.1-M89 (CSA 1989) since it was estimated that the

weld was critical for Lw/w ratios less than one.

A specific study of shear lag-induced fracture in tubular connections started in early 1990s

when British Steel (1992) studied gusset plate connections to circular hollow sections (CHS),

square hollow sections (SHS) and rectangular hollow sections (RHS) under tension and

compression loading. An experimental program on slotted SHS and RHS to gusset plate

connections was undertaken by Korol et al. (1994). In this program, a total of 18 specimens with

Lw/w ratios ranging from 0.40 to 1.00 were tested. Their results confirmed that a net section

failure can occur in connections with ratios Lw/w < 1.00. Moreover, a ratio of Lw/w = 0.60 was

proposed as a lower limit for the net section failure mode. A FE analysis of these connections

was made considering only their elastic response, hence the FE models could not predict the

failure mode. Based on these models, a further parametric analysis determined the influence

that geometrical ratios have on the shear lag phenomenon; the Lw/w ratio was shown to have

the major influence and tube effective depth-to-width ratio a minor influence. Finally, the results

indicated the need for variable shear lag factors for slotted SHS and RHS connections.

Figure 2.4 Important dimensions in slotted end connections


2-3

Girard et al. (1995) generated a FE model of a connection between a SHS and a gusset

plate. Even though this FE model exhibited some limitations, their results displayed differences

with the equations in CAN/CSA-S16.1-M89 (CSA 1989). Cheng et al. (1996) studied the

phenomenon in CHS, undertaking an experimental program and a FE analysis of these

connections. A total of nine connections were tested, these connections were fabricated with a

slotted tube and, except for one, all had a weld return. The results showed the inaccuracy of the

shear lag factors in CAN/CSA-S16.1-94 (CSA 1994) for this type of connection. Additionally, the

results for the connection with no weld return always presented an uneven strain distribution at

the slotted end. For the same CAN/CSA-S16.1-94 (CSA 1994), Korol (1996) reached a similar

conclusion for slotted gusset plate connections fabricated with SHS and RHS. Cheng et al.

(1998) and Cheng and Kulak (2000) suggested that the reduction in the effective net area would

be eliminated for CHS connections if a minimum weld length (Lw) of 1.3 times the tube diameter

is provided.

Experimental programs in gusset plates slotted into RHS were also undertaken by Zhao

and Hancock (1995), Zhao et al. (1999) and Wilkinson et al. (2002). Although the failure mode

in the latter was not directly related to the shear lag effect, the results suggested the need to

verify the factors to account for shear lag. Recently, CHS connections with very high strength

tubes have been studied by Ling (2005), resulting in a design method which considers the heat

affected zone. However, due to the characteristics of the tube material used during this

experimental program these results may not be suitable for regular grade HSS connections.

Humphries and Birkemoe (2004) studied primarly double channel to gusset plate connections

but these were compared with RHS to gusset plate connections. The results showed that the

channels had a better behaviour than the RHS as they were able to deform reducing the

eccentricity ( ), thus increasing the connection effiency. This study also pointed out the

influence that the weld leg size (al) has on the connection strength, as an increase in this was

associated with an enhancement of the connection efficiency.

Although these research studies have contributed information related to the influence that

shear lag has in tubular connections, they have also showed the need to continue with more

definitive studies in order to provide design provisions with formulae that accurately reflects this

phenomenon.

x


2-4

2.2 Tear-out failure

In general, the research on tear-out failure or block failure has been mainly aimed at

bolted connections, using gusset plates, coped beams or angles. The first model for tear-out

failure (based on tests of coped beam connections) was proposed by Birkemoe and Gilmor

(1978) and was eventually included in the AISC specification (1978). This model calculates the

connection resistance by adding the shear resistance of the shear area and the tensile

resistance of the net tensile area. Since then, several investigations have been undertaken on

different bolted connections types.

In order to verify the accuracy of the AISC specification (1978), Yura et al. (1982) tested

twelve beam web shear connections. During these tests, several parameters such as: the edge

distance, standard and slotted holes, coped beams, uncoped beams and bolt arrangement

were studied. The results revealed a decrease in the connection capacity (approximately 20%)

when slotted holes were used, and the use of two rows of bolts clustered at the top of the web

produced a lower safety factor than that expected. Finally, for a connection with a single row of

bolts, a recommendation to calculate the connection capacity as the sum of the bolts single

capacity rather than a group capacity was made. In a further study (Ricles and Yura 1983), a

finite element analysis of these connections (considering only the connection elastic response)

showed a uneven stress distribution along the vertical plane at the cope. These results

disagreed with an ideal stress distribution calculated by simple beam theory. In general, fracture

initially started at the tension region where an uneven stress distribution was taking place and it

was combined with a substantial material yielding along the shear plane. Based on these

results, a new block shear model (with a triangular stress distribution on the tension region) was

proposed for double row bolted connections. Hardash and Bjorhovde (1985) evaluated the

application of the block-shear concept in gusset plates connections via the testing of 28

specimens. The test specimens were fabricated with two lines of bolts with various bolt rows,

pitch spacing and bolt diameters. During these tests, the dominating failure mode corresponded

to the attainment of the ultimate stress along the net area in tension (at the last row of bolts) and

yielding of the gross area in shear (outside of the line of the bolts). In addition to this, the data

from experimental programs at the University of Illinois and the University of Alberta were

combined with these results to develop a new block shear model. In general, this new model

followed the original “block failure” model. Nevertheless, it included several new factors to

calculate the ultimate resistance of the connection which made its use difficult.


2-5

Epstein (1992) undertook an experimental program to study block shear failure in angles,

with a total of 38 angle connection tests. These results showed variations with the values

recommended by AISC design provisions (1986, 1989) at that time. These differences were

mainly associated with the effect that the unconnected leg eccentricity had over the connection

behavior, modifying the failure mechanism. Gross et al. (1995) tested 13 angle connections

fabricated with a single line of bolts and steel grades A-36 and A588. In general, the results

showed good correlation with AISC design provisions (1989, 1994) based on agreement with

the failure load. However, an inconsistency was observed between the failure mechanism

predicted by design provisions and experimental test. Based on data published in previous

experimental programs, Cunningham et al. (1995) suggested a model to predict block shear

failure in connections fabricated with angles and bolts. Orbison et al. (1999) tested several

angles, WT and W sections which failed in block shear (a total of 17 specimens). The failure

mechanism observed during the tests consisted of a fracture at the tension area which was

combined with a considerable inelastic deformation along the gross shear area. Even though

the predicted connection capacity by the (then-current) design provision (AISC 1994) resulted in

conservative values, the expected failure mechanism disagreed with the tests results.

Additionally, several factors such as: low ductility, hole fabrication (punched or drilled) and large

in plane and out-of-plane eccentricities were found to have an influence on the connection

capacity. Finally, a further study of these factors was suggested since they were not considered

in design provisions. Swanson and Leon (2000) tested 48 T-stub specimens under monotonic

and cyclic loading. From all these test specimens, only one failed by block failure (this specimen

was tested under cyclic loading). For this test specimen, the predicted failure mechanism (AISC

1994) did not coincide with the failure observed during the test. Aalberg and Larsen (2000)

tested splice plates, beam web connections loaded in shear and beams connections with a

coped end using high strength steels. The results were compared with design provisions such

as: Eurocode (CEN 1992), CSA (1989) and AISC (1994). In general, an important decrease in

the connection ductility was observed as a result of the use of these steel types and the

importance of limiting the deformation of these connections was addressed. For block shear

failure, only the CSA (1989) method was found to be suitable for high strength steels. A review

of the rules for block shear design (AISC 1999) by Kulak and Grondin (2001) suggested that

these may be conservative for gusset plates, acceptable for angles and non-conservative for

coped beams. This study recommended that further research of this failure mode was required.


2-6

In addition to these experimental programs, several studies have been undertaken with

the assistance of finite element models by Epstein and Chamarajanagar (1996), Epstein and

McGinnis (2000), Barth et al. (2002) and Topkaya (2004). In this last study, new models to

calculate the tear out failure have been suggested.

As result of these research programs, the governing failure criteria defining the block

shear as well as resistance factors have experienced several modifications in existing design

provisions (Geschwindner, 2004). However, the initial model (suggested by Birkemoe and

Gilmor) which adds the resistances in tension and shear continues in use. Nowadays, the new

trend to design by block shear follows this model, but with the use of several reduction factors.

As an example of this, the AISC design provision (2005) has suggested a reduction factor (Ubs)

to consider the uneven stress distribution that can be found in coped beams. Finally, a unified

equation suitable for all types of connections has been recently proposed by Driver et al. (2006),

wherein the initial model is used but several factors are applied depending on the connection

type.

2.3 International specifications

When the capacity of a tension member is governed by the limit state of tensile fracture

affected by shear lag, several values can be calculated from current design provisions as they

do exhibit differences. In general, these provisions consider the non-uniform stress distribution

caused by shear lag by including an efficiency factor (U). This factor decreases the tube net

area (An) at the connection to an effective net area (Ae or A'ne).

Ae = An · U (as in AISC 2000, 2005) (2-1)

A'ne = An · U (as in CSA 2001) (2-2)

This effective net are is then used to calculate the connection strength. In order to

calculate this efficiency factor (U), two general methods are currently most common. The first

method can be found in American specifications (AISC 2000, 2005), where the connection

eccentricity ( ) is compared with the weld length (Lw), as proposed by Cheeson and Munse

(1963) to allow for the shear lag phenomenon in riveted and bolted connections. Specifications

using this approach are summarized in Table 2.1. By this method:

x


2-7

where for CHS; (2-3)

and for EHS (see Figure 2.5). (2-4)

Equation 2-4 considers that the gusset plate is aligned with the dimension D2 (see Figure

2.5). When the gusset plate orientation is parallel to the dimension D1, the dimension D12

should be replaced by D22. The conventional interpretation of has been the measurement

from the tube centroidal axis. However, when a thick gusset plate is utilized. It may be feasible

to consider a reduced , which is the distance from the gusset plate surface to the centre of

gravity of the half tube as shown in Figure 2.4.

The second method compares the circumferential distance between the welds (w) with

the weld length (Lw). Here the efficiency factor (U) is determined by values assigned to the ratio

Lw/w (see Table 2.1). This method can be found in the Canadian specification (CSA 1994,

2001) as well as in the design guide for hollow structural sections by Packer and Henderson

(1997). Moreover, for slotted connections to hollow sections the distance w equals half of the

HSS circumference minus the gusset plate thickness (tp) or the slot width (tsl). Eurocode3 (CEN

2005) only considers the effect of shear lag on bolted connections using angles connected by

U 1 xLw------–=

x Dπ----=

x 23π------D1

2 2D1D2+D1 D2+

------------------------------=

x

x'

Figure 2.5 Eccentricity of top half, for EHS.x


2-8

one leg and other unsymmetrically connected tension members. Eurocode3 (2005) hence is not

listed in Table 2.1.

Tr = Ae Fu (AISC Specification, = 0.75) or Tr = 0.85 A'ne Fu (CSA Specification, = 0.9).

For block shear failure, the connection resistance is calculated by adding the portion of

the load transferred as tension load, Tr, and the portion of load transferred as shear load, Vr.

The different national/regional design specifications (AISC, CSA, Eurocode) either use the

gross or net area for the calculation of Tr and Vr (see Table 2.2). In welded connections, the

gross area becomes equal to the net area for the calculation or Tr and Vr due the absence of

bolt holes. For the calculation of the shear load, the material strength is reduced to 0.60 Fy or Fy

Table 2.1 Shear lag design provisions for round (and elliptical) hollow sections

Specification or design guide

Effective net area Shear lag coefficients Range of

validity

AISC (1999):LRFD Specification for Structural Steel Buildings

Ae = An · U

with (for CHS)

(EHS, see Figure 2.5)

no restric-tionsAISC (2000):

LRFD Specification for Steel Hollow Structural Sections

AISC (2005):Specification for Structural Steel Buildings

U = 1- for

U = 1 for (only CHS)

CSA (1994):Limit States Design of Steel Structures

A'ne = An · U

U = 1.0 for

U = 0.87 for 2.0 >

U = 0.75 for 1.5 >


U = 1.0 for

U = 0.5 + 0.25 for 2.0>

U = 0.75 for < 1.0

no restric-tions

Packer and Henderson (1997):Hollow Structural Section Connections and Trusses - A Design Guide

U = 1.0 for

U = 0.87 for 2.0 >

U = 0.75 for 1.5 >

U = 0.62 for 1.0 >

shear lag not critical

for < 0.6

U 1 xLw------– 0.90≤=

x Dπ----=

x 23π------D1

2 2D1D2+D1 D2+

------------------------------=

xLw------ 1.3D Lw> D≥

Lw 1.3D≥

Lw D≥

Lw w⁄ 2.0≥

Lw w⁄ 1.5≥

Lw w⁄ 1.0≥

Lw w≥

Lw w⁄ 2.0≥

Lw w⁄ Lw w⁄ 1.0≥

Lw w⁄ Lw w⁄

Lw w⁄ 1.0≥

Lw w⁄ 1.5≥

Lw w⁄ 1.0≥

Lw w⁄ 0.6≥ Lw w

φ φ φ φ


2-9

/ . The factor Ubs used in the American specification (AISC 2005) has been introduced to

account for the stress distribution that can be found in coped beams, where Ubs=0.5 is

recommended. In gusset plate connections Ubs is taken equal to unity.

The Canadian specification (CSA 2001) uses a separate design formula for coped beams

but it also results in the same reduction factor as the American specification. It is worthwhile

noting that the latest Canadian and American specifications, while having essentially the same

model for the block shear limit state, result in considerably different safety levels due to their

different resistance factors ( ), as shown in Table 2.2 (although the Canadian value is currently

under review). This is not the case for the shear lag design provisions (Table 2.1), where

.

a) Design rule for bolted connections differs slightly.

Table 2.2 Block shear design provisions

Specification or design guide Block shear strength

AISC (1999):LRFD Specification for Struc-tural Steel Buildings

When Ant Fu 0.6 Anv Fu: Tr + Vr = [Ant Fu + 0.6 Agv Fy] [Ant Fu + 0.6 Anv Fu]

When Ant Fu < 0.6Anv Fu:Tr + Vr = [Agt Fy + 0.6 Anv Fu] φ [Ant Fu + 0.6Anv Fu]

with = 0.75

AISC (2000):LRFD Specification for Steel Hollow Structural Sections

AISC (2005):Specification for Structural Steel Buildings

Tr + Vr = Ubs Ant Fu + 0.6 Agv Fy Ubs Ant Fu + 0.6 Anv Fu

with = 0.75 and Ubs= 1


Tr + Vr = Ant Fu + 0.6 Agv Fy Ant Fu + 0.6 Anv Fu

with = 0.9

Eurocode (CEN 2005):Design of Steel Structures - General Rules - Part 1-8: Design of Joints a)

Tr + Vr = Ant Fu Anv Fy

= 1.0 and = 1.25

3

φ

0.9( ) 0.85( ) 0.75≈

≥φ ≤ φ

φ ≤φ

φ φ ≤ φ φφ

φ φ ≤ φ φφ

2

1Mγ 3

110Mγ

+

0Mγ 2Mγ


2-10

2.4 Summary of Chapter 2

As has been exposed throughout this chapter, research on TO failure has been mainly

focused on several types of bolted connections. As a result of this, the first model suggested to

predict the connection strength has experienced several modifications throughout the years.

Nevertheless, the accuracy of this model still seems to need further attention or verification

(especially for welded tubular connections).

To account for shear lag (inducing a CF) in tubular connections, two general approaches

are prevalent nowadays in current design provisions. However, the accuracy of these models

has not been totally verified for slotted end connections to CHS or EHS.

In order to asses the accuracy and suitability of the models recommended in current

design provisions (which are suggested for the TO failure limit state and to account for shear lag

phenomenon), these models are compared against the results from an experimental program

carried out at the University of Toronto (Chapter 3 of this Report) and other relevant research

programs undertaken on tubular connections (Chapter 4 of this Report).


3-1CHAPTER 3: EXPERIMENTAL PROGRAM

An experimental program has been undertaken at the University of Toronto on slotted end

connections to hollow sections (CHS and EHS). The objective of this study was to identify the

influence of parameters such as: the weld length (Lw), the eccentricity of the connection ( ), the

gusset plate orientation (for the EHS) and fabrication detail on the connection strength. In

general, these parameters have been shown to affect the shear lag phenomenon in previous

experimental programs and the calculated connection strength by current design codes is

based on these parameters. As part of this program, a total of 13 connections were fabricated

and tested under quasi-static tension and compression loading. A description of the

connections, the material properties, the testing arrangement and results from the tests are

given in this chapter.

3.1 Material properties

For the fabrication of the connections, a CHS with a nominal size of 168 x 4.8mm was

used and it was cold-formed Class C material with a minimum specified yield stress of 350MPa

(CSA 2004). An EHS with a nominal size of 220x110x6.3mm was used and it was hot-finished

with a minimum specified yield stress of 355MPa (EN 10210-1, CEN 1994). Plates with 25mm

and 32mm thickness were required for the fabrication of the gusset plates; these plates had a

minimum specified yield stress of 300MPa (CSA 2004). A group of test coupons was fabricated

from tubes and plates in order to determine their material properties. Seven test coupons were

taken from the CHS with two of them cut from the Heat Affected Zone (HAZ). A 25mm plate was

used in the fabrication of the CHS connections and two test coupons were cut from this plate.

Four test coupons were cut from the EHS and three 32mm plates were used in the fabrication of

these connections so a total of six coupons were tested from these plates. The size and location

of these coupons were made according to ASTM (2003).

During testing, the engineering stress-strain relationship was acquired before the coupon

test developed a neck. Afterwards, the clip gauge was removed from the test coupon. In some

test coupons from the CHS, it was possible to acquire information beyond the formation of the

neck but eventually the clip gauge had to be removed. In all the cases, the load and maximum

elongation at rupture were determined for each coupon test. The engineering stress-strain

curves from the materials are shown in Figures 3.1 to 3.4 and their measured material

x

SLOTTED END CONNECTIONS TO HOLLOW SECTIONS, CH 3: EXPERIMENTAL PROGRAM

3-2properties are given in Table 3.1. Additional information from the tube and gusset plate material

is given in Appendix A.

Figure 3.1 Coupon tests for CHS

Figure 3.2 Coupon tests for 25 mm plate


3-3

Figure 3.3 Coupon tests for EHS

Figure 3.4 Coupon tests for 32 mm plate


3-4

a) Properties determined by the average measurements from several tensile coupon tests.b) Using the 0.2% offset method, as material was cold-formed.

3.1.1 Stub column tests

In addition to the test coupons, a stub column test was performed on both the CHS and

the EHS to determine their properties under compression load. The specimen size and the

testing procedure were as recommended by SSRC (Galambos 1998). Before testing, four strain

gauges were placed around the tube’s circumference at the mid-height (see Figure 3.5). This

allowed the generation of an average σ-ε relationships for the tube materials. Results from the

tests are given in Table 3.2

a) Measured area obtained by weighing a tube segment and using a density of 7850 kg/m3.b) Average length measured with a caliper.c) Csc= Stub column ultimate compressive strength.

Table 3.1 Measured material properties

E(GPa) a) Fy(MPa) a) Fu(MPa) a) εu(%) a)

CHS 196 498 b) 540 25.9

EHS 216 421 530 34.7Plate (tp=25.7mm) 201 358 482 28.0

Plate (tp=32.0mm) 214 356 472 30.0

Table 3.2 Stub column properties and test results

Length (mm) Weight (Kg) Area a) (mm2) Cscc) (kN)

CHS 150 b) 2.91 2471 -1213

EHS 104.7 b) 2.51 3053 -1393

Figure 3.5 Strain gauges on stub columns


3-5Using the data acquired through the test of the CHS (see Figure 3.6), the calculation of

average Young’s Modulus agreed with the value previously determined from the tensile test

coupons. A similar conclusion was achieved from the computation of the average yield stress.

For the EHS, the average Young’s Modulus (see Figure 3.7) also agreed well with the value

previously determined by tensile test coupons. However, an increase of 8% was observed when

the EHS stub column yield stress was compared to the tensile test coupons. This difference has

been attributed to the uneven manner in which the EHS stub column changed its shape through

the test, which likely resulted in a higher value.

Figure 3.6 Stub column response of CHS


3-6

3.2 Test specimens and instrumentation

A total of six connection types were examined throughout this experimental program (see

Figure 3.8). Connection type A was fabricated with a slotted CHS which was connected to a

25mm thick gusset plate by longitudinal fillet welds. Connection type B was originally fabricated

as connection type A, however, the slot was filled in when the weld return was included. This

connection type eliminates the reduction in the gross cross-sectional area of the tube due to

slotting. For connection type C, a 25mm thick gusset plate was slotted so the CHS gross cross-

sectional area remained unaffected. For this connection type, the tube and the gusset plate

were connected by longitudinal fillet welds too.

For the CHS tension tests, two specimens were fabricated for each connection type (A, B

and C) and the main difference between specimens (from a similar connection type) was their

weld length. Hence, they were labelled in a progressive order as the weld length increased.

Additionally, specimens from the connection types A and C were fabricated and tested under

compression loading.

Five EHS specimens were fabricated for tensile testing. In order to avoid confusion

amongst the EHS connections, these were simple labelled in a progressive order (E1 to E5)

0

50

100

150

200

250

300

350

400

450

500

0 0.0002 0.0004 0.0006 0.0008 0.001 0.0012 0.0014 0.0016 0.0018 0.002

Strain (mm/mm)

Str

ess

(MP

a)

SGE SGW

SGN SGS

Figure 3.7 Stub column response of EHS


3-7depending on their connection type and weld length. The connection types E1 and E2 were

fabricated with a slotted EHS, with the gusset plates oriented to give a large eccentricity and

only longitudinal weld lengths were used to transfer the load. Connection type E5 was similar to

these connections, however the orientation of the gusset plate was changed to give a smaller

eccentricity. In general, the connection types E3 and E4 were similar to connection type C, but

the EHS was oriented to produce a large eccentricity.

In all cases, the test specimens had a Lw/w ratio within the range from 0.60 to 0.90 which

guaranteed the presence of the shear lag phenomenon during the tests. All gusset plates and

welds were dimensioned so as not to be critical. Fillet welds had a nominal size of 10 or 15 mm

and they were fabricated using E480XX electrodes (CSA 2003). The tube lengths were 1.5 and

2.0 metres for the CHS and EHS respectively. In order to facilitate the tests, two identical

connections were fabricated at each tube end, which allowed the testing of two connections

with very similar weld lengths simultaneously (see Figure 3.9). The average dimensions and

properties of the specimens are shown in Table 3.3. Additionally, all measured dimensions from

the test specimens are given in Appendix A.

Figure 3.8 Connection types examined


3-8

a) Measured area calculated by weighing a piece of HSS and using a density of 7850 kg/m3

All the specimens were loaded in quasi-static axial tension to failure in a universal testing

machine and displacement control was used throughout each test. Four LVDTs (linear variable

differential transformers) were placed on each specimen to measure deformations during the

test. The tube deformation reported herein corresponds to the average deformation measured

by two LDVTs from the centre of the tube to the gusset plate. Each specimen was also equipped

with 10 strain gauges to establish the strains in the connection region (see Figure 3.10). All this

information was acquired with a computer during the tests and the use of white-washing

allowed the identification of regions with high strain concentration that in most cases induced an

early fracture in the tube material.

For the compression tests performed on specimens A3C and C3C, a minimum free

distance of 2tp was provided in the gusset plate between the machine clamps and the tube

ends. In addition to the instrumentation used in the tension tests, a fifth LVDT was placed at the

test specimen mid-height to measure its out-of-straightness during the test.

Table 3.3 Measured dimensions and geometric properties of test specimens

Specimen Tubeal

(mm)Lw

(mm)w

(mm)Lw/w (mm)

tp (mm)

Wp (mm)

A1

CHS168.5x4.89

Aa)=2471 mm2

10 156

238

0.65

25.7

197A2 10 192 0.80 198

A3C 10 206 0.86 197B1 9 169 0.71 197B2 9 208 0.87 198C1 14 162

2390.67 2 x 74.3

C2 14 195 0.81 2 x 75.5C3C 14 200 0.83 2 x 74.3E1

EHS110.9x221.2x5.94

Aa)=3054 mm2

13 145234

0.61

32.0

161E2 14 182 0.77 161E3 15 146

2370.61 2 x 94.0

E4 15 175 0.73 2 x 93.8E5 15 185 234 0.79 270


3-9

Figure 3.9 Experiment setup for tests

Figure 3.10 Location of strain gauges on test specimens


3-103.3 Experimental test results

During the tests, the connection strength was determined principally by one of the

following failure mechanisms: a) a tear out failure (TO) where the crack initiated at the weld

termination then propagated through the tube base material near the weld toes, b) a

circumferential fracture (CF), where the crack initiated at the weld termination then propagated

around the tube circumference, and c) a combination of both failure modes (TO-CF). In the

latter, both failure mechanisms occurred simultaneously at the connection end. The four LVDTs

installed on each test specimen measured the overall elongation from the mid-length of the test

specimen to the gusset plates. Even though two connections were fabricated alike for each test

specimen (one at each tube end), failure was generally concentrated at one end. This

behaviour has been attributed to variations in actual weld lengths and imperfections included

during fabrication. The load-deformation response shown for the test results corresponds to the

failed connection. In general, all the connections exhibited an uneven strain distribution along

the connection and around the tube circumference. From the data acquired during the test, the

strain distribution in the connections is only presented for a stage near the end of the

connection elastic response. The rest of the strain readings are given in Appendix C.

3.3.1 Slotted CHS connection - slot end not filled (type A)

The use of this connection type is advantageous since the fabrication tolerance for the

slot makes assembly of the parts easier. However, the presence of an open slot end can affect

the overall connection behaviour, as seen by the tests. In general, the behaviour of these

connections can be described in several stages. Initially, the connections showed an elastic

response with an equivalent constant stiffness. Afterwards, the strain concentration in the slot

region (due to the presence of the shear lag phenomenon) induced yielding of the tube material

there, thus modifying the overall connection stiffness. The magnitude of the shear lag (affecting

each connection), which is determined by the weld length, increases as the weld length

decreases, and the weld length was the only distinction between the two test specimens.

(Figure 3.11 shows a superior performance for the test specimen with the longer weld length,

A2). At this yielding stage, whitewash flaking confirmed the strain concentration taking place in

the tube base material near the weld start (in the slot region). The strain gauge readings from


3-11the test specimens also confirmed this, as they showed an uneven strain distribution around the

tube circumference and along the connection (see Figure 3.12).

As the test specimens elongated, deformation was concentrated in the slot region

producing a gradual change of the tube shape (inducing the formation of a neck there). In

addition, the uneven strain distribution at the slot cross-section (due to the shear lag

phenomenon) stimulated a quick increase in the strains at the weld start location, where

straining of the tube material continued until fracture occurred. In general, a longer weld length

allowed a better load transfer over the connection which diminished the connection

deformation, however tube material fracture always governed the connection behaviour. Once

fracture started, the crack continued to propagate gradually from the weld heel to its toe. Then,

depending of the load level and the strain distribution in the connection, the crack would

continue to propagate over the weld length (TO) or around the tube circumference (CF).

Specimen A1 showed both failure modes and specimen A2 only CF (see Figure 3.13). The

maximum load and deformation attained by these test specimens are shown in Table 3.4.

Figure 3.11 Load-deformation response for connections type A


3-12

Table 3.4 Ultimate capacity for connections type A

Weld Length (mm)

Test Load Nux (kN)

Deformation @ Max Load

(mm)

FailureMode

Nux/AnFu

Specimen A1 156 1032 8.8 TO-CF 0.87Specimen A2 192 1154 8.8 CF 0.97

Figure 3.12 Strain distribution for test specimens A1 and A2 at 800kN

Figure 3.13 Failure in test specimens A1 and A2


3-133.3.2 Slotted CHS connection - slot end filled (with a weld return) (type B)

The addition of a weld return to these test specimens eliminated the possibility of a failure

through the tube net area. Moreover, it allowed the attainment of the maximum load with small

deformations (see Figure 3.14). In both tests, the load increase produced a strain concentration

that was located at the weld return region (specifically at the weld toe). This behaviour has been

attributed to the difference in the ductility of the return welds, since these were loaded at 90º

with respect to their longitudinal axis which creates a region of high stiffness.

Whitewash flaking confirmed the strain concentration taking place at the weld return

region as the tube material yielded there at an early stage of the tests. Moreover, the readings

of the strain gauges always exhibited very uneven strain distributions around the connections.

Figure 3.15 shows the strain distribution around the tube and along the connection length, at the

end of the elastic response.

The strain gauge readings around the tube circumference showed an improvement

compared to the strain distribution from connections type A. However, the strains experienced

an increase right at the weld return region, relative to connections type A (see Figures 3.15 and

3.12 at z=+50mm). For specimens B1 and B2, the strain distribution presented a dependency

Figure 3.14 Load-deformation response for connections type B


3-14on the weld length and the specimen B1 (which had the smaller weld length) showed higher

strain at z=+50mm (see Figure 3.15).

Once the overall connection stiffness noticeably changed, any load increment was

associated with a gradual increase of the strains in the weld return region and a change in the

tube cross-section shape. The maximum load was limited by the propagation of a crack in the

tube material near the weld return toe. This crack spread gradually at a 45 degree angle from

the gusset plate. Finally, specimen B1 showed a TO failure and specimen B2 a CF (see Figure

3.16). The maximum load and deformation attained by these test specimens are shown in Table

3.5.

Table 3.5 Ultimate capacity for connections type B

Weld Length (mm)

Test Load Nux (kN)


(mm)

FailureMode

Nux/AnFu(An=Ag)

Specimen B1 169 1087 6.1 TO 0.91Specimen B2 208 1211 6.1 CF 1.02

Figure 3.15 Strain distribution in test specimens B1 and B2 at 800kN


3-15

3.3.3 Slotted EHS connection - slot end not filled (gusset plate oriented to give a large

eccentricity)

The behaviour of these connections emulated the response of specimens type A.

However, some differences occurred herein which have been associated mainly to the tube

geometry. During these tests, the overall connection response can be described by several

stages. At first, the test specimens had a similar elastic stiffness, while strain concentrations

developed at the slot region (specifically in the tube near the weld start). This eventually caused

tube material yielding at that location and affected the overall connection response. In general,

the magnitude of this strain concentration was directly determined by the weld length. As a

consequence, the elastic response of specimen E1 had an early ending (relative to specimen

E2) as it had the shorter weld length (see Figure 3.17).

Figure 3.16 Failure in test specimens B1 and B2


3-16

At a load of 600 kN (near the end of the elastic response), the strain gauge readings

around the tube circumference showed an uneven strain distribution (as expected for this

connection type). In both tests, the maximum strain along the longitudinal weld took place at the

weld beginning and a much lower value was recorded at the slot open end (see Figure 3.18). At

this load level, considerable differences were observed between the readings from specimens

E1 and E2 in the weld region. E2 had higher local strains than E1, despite having a longer weld

length, which initially represented an inconsistency with the results from other connections

(where the strain concentration decayed as the weld length increased). A further examination of

specimen E1 revealed that during the fabrication of specimen E1 the tube was over-slotted, with

a slot length of 268 mm. This dimension far exceeded the required weld length which was only

145 mm. Moreover, the weld fabrication started near the slot end leaving a considerable portion

of the slotted tube free behind the welds (see Figure 3.19). Hence, the progressive deformation

of connection E1 was accompanied by a bowing outwards of the free slotted tube portion as the

load increased. This may have positively affected the strain distribution in the connection since

it modified the strain concentration at the slot end. The bowing in the slotted tube E1 did not

eliminate the shear lag phenomenon, but was sufficient to change the connection strain

distribution.

Figure 3.17 Load-deformation response for connectionstype E1 and E2


3-17

Once the tube material started to yield, the connection deformation began to concentrate

near the open slot region (adjacent to the beginning of the welds). This local straining was

combined with gradual propagation of material yielding in surrounding areas, illustrated by

flaking of the whitewash along the connection. In addition, yield lines emanated from the slot

into the tube. In both test specimens, these yield lines were neatly depicted on the tube surface

(this contrasted with the CHS connections where material yielding was mainly exemplified by a

region rather than lines). This different behaviour has been attributed to the EHS tube material

properties, which exhibited a clear yield plateau unlike the CHS material. Finally, close to the

attainment of the maximum load, the tube started to neck at the open slot region, slowing the

load increase. Then, the connection distortion stopped as the tube material fractured (see

Figure 3.19).

Figure 3.18 Strain distribution for test specimens E1 and E2 at 600kN


3-18

The crack continued propagating around the tube circumference (CF) in both specimens

until complete tube rupture. Although the maximum load in specimen E2 nearly reached the

tube gross cross-sectional area yield load (AgFy =1286 kN), the capacity was still limited by the

uneven strain distribution induced by shear lag. Finally, the maximum load and deformation

attained by these test specimens are shown in Table 3.6.

Table 3.6 Ultimate capacity for connections type E1 and E2

3.3.4 Slotted EHS connection - slot end not filled (gusset plate oriented to give small

eccentricity)

The change in the gusset plate orientation significantly improved the behaviour of this test

specimen relative to its counterpart with a large eccentricity (see Figure 3.20).

Weld Length (mm)

Test Load Nux (kN)


(mm)

FailureMode

Nux/AnFu

Specimen E1 145 1109 9.9 CF 0.81Specimen E2 182 1236 11.1 CF 0.90

Figure 3.19 Failure in test specimens E1 and E2


3-19

At a load of 1040 kN (near the end of the elastic response), the strain gauge readings

around the tube circumference showed a very uneven strain distribution, illustrated by Figure

3.21 (as was observed previously in specimens E1 and E2). Along the parallel welds, the strain

distribution again reached its maximum value at the beginning of the weld as before.

Figure 3.20 Load-deformation response for connection type E5

Figure 3.21 Strain distribution in test specimen E5 at 1040 kN


3-20As part of the transition from an elastic response to a distinct yield plateau, the connection

deformation began to concentrate at the open slot region and shear yield lines (visible due to

the whitewash flaking) also emanated from this region towards the tube mid-length. The low

connection eccentricity significantly improved the load transfer from the EHS to the gusset

plate, relative to its counterpart with a large eccentricity. This also decreased the strain

concentration occurring at the beginning of the weld, thus allowing the attainment of the yield

stress across the tube net section. At this load level, shear yield lines continued to propagate

but now over the entire tube length, increasing the overall deformation from 12 to almost 27mm.

In contrast with test specimens E1 and E2 (where the overall deformation was mainly

concentrated at the slot region), the total deformation here was a combination of the

deformation at the slot region plus the overall tube elongation due to material yielding. In order

to continue increasing the load, the material at the net section started to strain harden. The

uneven strain distribution taking place at the open slot, aggravated by the shear lag

phenomenon, eventually caused tube fracture there (see Figure 3.22).

Once material fracture began, the load decreased as a consequence of the crack

propagation around the tube circumference (CF), until complete tube rupture. Even though the

tube material reached strain hardening, the maximum connection efficiency (Nux/AnFu) was

restrained to only 94%. Nevertheless, this connection did allow the attainment of complete tube

yielding (AgFy=1286 KN) which may represent an advantage of this structural shape over the

CHS. The maximum load and deformation attained by this test specimen is shown in Table 3.7.

Table 3.7 Ultimate capacity for connection type E5Weld

Length (mm)

Test Load Nux (kN)


(mm)

FailureMode

Nux/AnFu

Specimen E5 185 1282 31.8 CF 0.94


3-21

3.3.5 Slotted gusset plate to tube connections in tension

This connection type avoids any loss of the tube cross-sectional area and its potential

effect on connection strength. Even though this may be considered its principal advantage over

slotted tube connections, the slot in the gusset plate can negatively affect the connection

stiffness, leading to excessive deformation of the gusset plate and consequently to the tube

cross-section (as was observed during the tests).

3.3.5.1 Slotted gusset plate to CHS connection (type C)

A strain concentration took place at the beginning of the welds (in the CHS) and interior

corners of the gusset plate. Close to 600 kN, the gusset plate yielded and caused flaking of the

whitewash there and a change in the overall connection stiffness (see Figure 3.23).

Figure 3.22 Failure in test specimen E5


3-22

For both test specimens, this happened at a lower load level than for the slotted tube

connections. At this load stage, the strain gauge readings around the tube circumference

showed an uneven strain distribution and the maximum strain value took place at the beginning

of the welds (see Figure 3.24). In addition, the minimum value (near zero) was detected for the

strain gauge located at 90º (see Figure 3.24), as for slotted tube connections. Moreover, close

to attainment of the maximum load the readings at 90º switched to negative values (indicating

compressive strains). This initially-unexpected behaviour was attributed to the excessive

distortion of the tube cross-section, due to the gusset plate bowing and the necking of the tube.

The readings along the parallel welds also showed typical variations, with the highest strain

concentration occurring at the beginning of the weld (see Figure 3.24). Of the two tests, the

higher strains were registered in specimen C1 which has the shorter weld length.

Figure 3.23 Load-deformation response for connections type C


3-23

Beyond the elastic response, each load increment resulted in increasing distortion of the

tube cross-section. Moreover, the bowing outwards of the gusset plate introduced out-of-plane

strains at the tube surface which are believed to have induced a triaxial state of stress at the

beginning of the weld. This behaviour continued throughout the tests until the material fractured

(see Figure 3.25).

Once the fracture started (at the beginning of the welds), the crack continued propagating

around the tube circumference (CF) in both tests. These tests again corroborated how the

presence of shear lag can affect the strain distribution in such connections. Nevertheless, the

magnitude of this strain concentration (which triggers the material fracture) is a consequence of

factors such as: magnitude of the shear lag, bowing of the gusset plate, tube cross-section

distortions and tube necking. Based on these two tests, it seems necessary to consider the

influence that the gusset plate dimension may have on the connection strength in a further

parametric analysis. Moreover, the potential need to limit the maximum load based on the tube

cross-section distortion has arisen herein since large distortions were observed before the

attainment of the connection maximum load. Finally, the maximum load and deformation

attained by these test specimens are shown in Table 3.8.

Figure 3.24 Strain distribution in test specimens C1 and C2 at 595kN


3-24Table 3.8 Ultimate capacity for connections type C

3.3.5.2 Slotted gusset plate to EHS connection (gusset plate oriented to give a large

eccentricity)

These two connections exhibited a strain concentration taking place in the EHS near the

beginning of the welds and also in the gusset plate adjacent to the end of the welds (at the tube

end). Near 850 kN, the materials in specimen E3 commenced yielding in these regions (as was

confirmed by whitewash flaking) producing a change in the overall connection stiffness (see

Figure 3.26).

Weld Length (mm)

Test Load Nux (kN)


(mm)

FailureMode

Nux/AnFu(An=Ag)

Specimen C1 162 1107 13.9 CF 0.83Specimen C2 195 1196 16.8 CF 0.90

Figure 3.25 Failure in test specimens C1 and C2


3-25

Once yielding started, it continued propagating from there towards the tube mid-length

while the strain concentration continued increasing near the weld region. In addition, shear yield

lines (visible from whitewash flaking) emanated from the connection region to the tube mid-

length. Even though the response of specimen E4 remained elastic at this load level (850 kN), it

followed a similar behaviour afterwards.

In order to transition from an elastic response to a distinct yield plateau, both connections

required a large elongation (approximately of 20 mm). These slotted gusset plates exhibited

smaller deformations throughout this transition, and even during the incursion of the EHS

material into the plastic range, in comparison to their CHS test counterparts (connection type

C). This better behaviour has been related to the higher moment of inertia of the gusset plates

for these connections (E3 and E4). Gusset plates with a 32mm thickness were used herein,

whereas 25.7mm gusset plate were used for connections type C (see Table 3.3).

At a load of 800 kN, the strain gauge readings around the tube circumference showed the

typical uneven strain distribution observed previously. An increase in the weld length (for

Figure 3.26 Load-deformation response for connectionstype E3 and E4


3-26specimen E4) only produced a marginal reduction in the strain magnitude (see Figure 3.27).

Moreover, close to attainment of the maximum load the strain readings at the gusset plate

corners reported a considerable increase which has been attributed to the tube ovalization in

the connection region.

Beyond the yield load, the connections entered the strain hardening range, but this load

increase was associated with a gradual increase in the bowing outwards of the gusset plate.

This distorted the tube cross-section and induced a triaxial state of stress near the beginning of

the weld. This gradual degradation of the connection continued until the material fractured (see

Figure 3.28).

Figure 3.27 Strain distribution for test specimens E3 and E4 at 800kN


3-27

Once the fracture started, the crack continued propagating around the tube circumference

(CF) in both tests. In a similar manner to the CHS connections, these tests illustrated the need

to consider the influence that the gusset plate dimensions may have on the connection strength

and the need to potentially limit the maximum load based on the tube cross-section. Finally, the

maximum load and deformation attained by these test specimens are shown in Table 3.9.

Table 3.9 Ultimate capacity for connections type E3 and E4

3.3.6 Connections under compression load

In order to comprehend the influence that the shear lag phenomenon may have on

connections under compression load, two connection types were studied during this

experimental program; the first connection was a slotted CHS to a gusset plate, and the second

was a slotted gusset plate to a CHS. In both cases, the governing failure mechanism

corresponded to a local bucking (LB) of the CHS at the connection region. Nevertheless, the

Weld Length (mm)

Test Load Nux (kN)


(mm)

FailureMode

Nux/AnFu(An=Ag)

Specimen E3 146 1336 43.9 CF 0.82Specimen E4 1175 1400 53.4 CF 0.86

Figure 3.28 Failure in test specimens E3 and E4


3-28occurrence of this failure mechanism in each connection was influenced by factors

characteristic of each connection type.

3.3.6.1 Slotted CHS to gusset plate connection - slot end not filled

The behaviour of this connection type may be explained by several stages. Initially, the

connection exhibited an elastic response that allowed the attainment of almost 70% of its

maximum load (see Figure 3.29). In a similar manner to the tension tests, a strain concentration

developed at the slot region but, due to the difference in the load condition herein (compression

loading), this strain concentration induced tube local buckling at the slot region rather than

material straining to fracture.

The tube local buckling started near the beginning of the welds and then gradually

extended to the entire slotted tube cross-section. As result of this, the tube deformation was

concentrated at that part of the connection. With increasing loads, the slot local buckle grew and

the gusset plate moved towards the CHS wall, thus reducing the slot length. From LVDT

readings, it could be ascertained that the tube out-of-straightness increased slowly before the

attainment of the maximum load. At the maximum load, the LVDT indicated an out-of-

straighness of 1.3 mm. After this load, full contact between the gusset plate and the tube wall

Figure 3.29 Load-deformation response for connection A3C


3-29occurred and this triggered an increase on the tube out-of-straightness and the rotation of the

gusset plate free length (the distance left between the CHS and the machine clamps). This

behaviour was associated with the start of overall bucking of the CHS. At this point, the test was

stopped in order to avoid large moments being applied to the testing machine.

In a similar manner to the tension tests, an uneven strain distribution existed throughout

the test. Near the end of the elastic response (at 800 kN), the compressive strain around the

tube circumference reached its maximum at the beginning of the welds in a comparable manner

to the tensile test specimens (see Figure 3.30). The strain distribution along the connection also

reproduced the behaviour seen during the tension tests (see Figure 3.30).

Even though the maximum load attained corresponded to 93% of AgFy, this required

considerable deformation at the slot (see Figure 3.31). Besides the weld length, the slot length

is likely to have a considerable influence on the connection capacity.These two parameters will

hence be considered in further parametric analysis. For a long member (such as a brace in a

regular building) an efficiency as high as 0.93 AgFy would never be required, as the brace

capacity will be governed by the member slenderness ratio. The maximum load and

deformation attained for the test specimen is shown in Table 3.10.

Figure 3.30 Strain distribution in test specimen A3C at 800kN.

z (see Figure 3.10),


3-30Table 3.10 Ultimate capacity for connection type A3C

3.3.6.2 Slotted gusset plate to CHS connection

In a similar manner to the tension test, this connection type had an initial elastic range of

almost 70% of its maximum load. Again, an uneven strain distribution formed at the connection

region with a high strain concentration near the beginning of the welds, resulting in local

buckling of the tube at the location. Even though an un-slotted tube would generally require a

high load to reach local buckling, the bowing inwards of the gusset plate exacerbated the tube

local instability. Near the end of the elastic response (600 kN), whitewash flaking was visible at

interior corners of the gusset plate. Then, the bowing inwards of the gusset plate began to affect

the connection deformation which, became a combination of the deformation of the tube and

the gusset plate. LVDT readings showed that the tube out-of-straightness slowly increased until

the attainment of the maximum load, at which time the out-of-straightness reached a value of

2.0 mm. After the maximum connection deformation progressed in a stable, ductile manner (see

Figure 3.32).

Weld Length (mm)

Test Load Nux (kN)


(mm)

FailureMode

Nux/AgFy

Specimen A3C 206 -1145 4.8 LB 0.93

Figure 3.31 Failure in test specimen A3C


3-31

In the same way as the tension connection type C, similar uneven strain distributions were

present throughout the test. The connection strain distributions (at 600 kN) recorded are shown

in Figure 3.33. Nevertheless, the lack of a slot reduced the strain concentration in front of the

weld region (at z = 50mm in Figure 3.33).

Figure 3.32 Load-deformation response for connection C3C

Figure 3.33 Strain distribution in test specimen C3C at 600kN.


3-32The bowing inwards of the gusset plate negatively affected the connection behaviour,

causing premature the tube local bucking at the weld region and a subsequent increase in the

tube distortion (see Figure 3.34). Because of this, the maximum load attained here

corresponded to only 70% of AgFy. This illustrates the need to limit the ultimate connection

capacity by a distortion limit, as has been suggested previously for the comparable tension test.

The influence of weld length and gusset plate dimensions will thus be considered in further

parametric analysis. The maximum load and deformation attained for the test specimen is

shown in Table 3.11.

Table 3.11 Ultimate capacity for connection type C3C

3.4 Summary of this experimental program

Even though all the tensile test specimens were designed to avoid failure by modes other

than circumferential fracture (CF), this was only accomplished by the test specimens A2, B2,

C1, C2, E1 and E5. Specimen B1 failed by block shear tear out (TO) of the base material along

the weld and specimen A1 presented a combination of both failure modes (CF-TO). In addition,

failure of the specimens tested under compression loading was due to local bucking at the

connection end (LB). By means of this program, it was found that several factors influenced

Weld Length (mm)

Test Load Nux (kN)


(mm)

FailureMode

Nux/AgFyAn=Ag

Specimen C3C 200 - 869 6.0 LB 0.70

Figure 3.34 Failure in test specimen C3C


3-33connection failure. In general, all hollow sections showed ovalization of their cross-section

before failure and this was especially pronounced in connections with a slotted gusset plate

(C1, C2, C3C, E3 and E4). Moreover, strain gauge readings always showed a very non-uniform

strain distribution during the tests. (Only some of the strain gauge readings have been shown in

this chapter; the remainder can be found in Appendix C. There, these readings are compared

with readings from FE models of these connections). Finally, a summary of the ultimate

capacities and failure modes for all tests is given in Table 3.12. These results are evaluated

against current design provision in the next chapter.

Table 3.12 Ultimate capacity of test specimens

Test Specimen Failure Mode

Test Load Nux (kN)

A1 CF-TO 1032A2 CF 1154

A3C LB -1145B1 TO 1087B2 CF 1211C1 CF 1107C2 CF 1196

C3C LB -869E1 CF 1109E2 CF 1236E3 CF 1336E4 CF 1400E5 CF 1282


4-1

CHAPTER 4: EVALUATION OF EXPERIMENTS AGAINST DESIGN PROVISIONS

A single governing limit state has not been established for slotted end connections, based

on geometrical properties such as the weld length (Lw), distance between welds (w) or the ratio

between these parameters (Lw/w). On the contrary, the philosophy behind current design

provisions (AISC 2005, CSA 2001 & CEN 2005) demands the check of several limit states

(such as yielding of the gross cross-sectional area, circumferential tensile fracture through the

net area and tear out failure) in order to determine the predicted connection strength. This

chapter hence shows a comparison between the tests results from research undertaken on

tension-loaded slotted end connections to hollow sections (see Table 4.1) and these provisions,

to verify their accuracy.

Table 4.1 Current research on slotted end connection to hollow sections

a) Failure of gusset plate (GP) or through the bolts (BL).b) Failure through the welds (WF).c) Connections fabricated with very high strength tubes (VHS).d) Local buckling (LB) of the connection since it was tested in compression.

For these comparisons, all the resistance factors ( ) and partial safety factors ( ) have

been equated to 1.0. (This includes the “partial resistance factor” corresponding to fracture

(0.85) in CSA). Only the experimental data corresponding to tear out failure (TO),

Research No. of tests Hollow section used Fabrication Type Failure Mode

British Steel (1992) 24 CHS, RHS & SHS Slotted tube to gusset plate

CF (16)GP or BLa) (8)

Korol et al. (1994) 18 RHS & SHS Slotted tube to gusset plate

CF (7)TO (11)

Zhao and Hancock (1995) 48 RHS & SHS Slotted tube to gusset plate TO (48)

Cheng et al. (1996) 9 CHS Slotted tube to gusset plate

CF (2)Necking (7)

Zhao et al. (1999) 24 RHS & SHS Slotted tube to gusset plate TO (24)

Wilkinson et al. (2002) 8 RHS Slotted tube to gusset plate

WFb) (5)TO (3)

By the author at U of T 13 CHS & EHS Slotted tube to gusset plate and Slotted gusset plate to tube

CF (9)TO (1)

TO-CF (1)LBd) (2)

Ling (2005) 45 CHS with VHSc) Slotted gusset plate to tubeTO (25)

TO-CF (4)CF (16)

φ γM

SLOTTED END CONNECTIONS TO HOLLOW SECTIONS, CH 4: EVALUATION OF EXPERIMENTS

4-2

circumferential tensile fracture (CF) of the HSS or overall tube necking are considered. All test

data pertaining to the tests performed by other researchers are contained in Appendix B.

4.1 Experimental program by British Steel (1992)

British Steel tested 24 specimens fabricated with CHS, RHS and SHS under tension

loading. These specimens were fabricated with a weld return which provided reinforcement to

the connection region (Hence, Ag=An). These specimens had Lw/w ratios ranging from 0.91 to

1.04; 15 specimens failed by CF, one specimen was considered to have developed a neck since

it attained 100% of AnFu, and eight failed through their gusset plate or the bolts. In addition to

these tests, a duplicate set of these 24 specimens was tested under compression loading.

However, the use of a three dimensional ring (to provide pinned ends to the specimens) and the

inclusion of an initial eccentricity due to the support configuration caused failure of the

specimens at lower loads than predicted. Thus, these compression test results have not been

included herein. Since the report from British Steel lacks information on the weld sizes, TO

failure predictions have been included herein using a weld leg size of zero.

Table 4.2 shows a comparison between test results and the predictions from current

codes or guides. From this table, one can appreciate that whereas for CSA (2001) and Packer

& Henderson (1997) the predicted failure mechanism was CF (considering that the connection

strength for all specimens was defined by CF), for AISC this was not the case (since 6 out of 16

specimens were predicted to fail by a TO failure) which does not reproduce accurately what was

seen during the test.

The use of the American design provision (AISC 2005), on average, provides a better

connection strength prediction (with a mean = 1.07) than the Canadian design provision (CSA

2001) when a CF failure is expected. Moreover, the use of a reduced eccentricity ( ) in the

AISC equation can improve these predictions even more. Even though AISC provides a better

overall CF prediction than CSA, it fails to consistently predict the correct failure mode (CF) since

the predicted block shear strength (TO) is sometimes lower. The use of the efficiency factors

suggested by Packer & Henderson (1997) results in very conservative design predictions.

x'


4-3

Table 4.2 Actual and predicted connection strength for British Steel (1992) data

a) CF represents circumferential fracture and NK represents necking of the HSS;b) As slot end was welded, it is likely appropriate to assume that An = Ag.

4.2 Experimental program by Korol et al. (1994)

Korol et al. tested 18 specimens, with Lw/w ratios ranging from 0.40 to 1.0 and fabricated

with slotted RHS and SHS. Since these test specimens were assembled with only longitudinal

welds, the strain concentration was located at the start of the welds (near the slot at the weld

heel). The tests failed by CF through the HSS net area in seven specimens while the other 11

failed by TO failure. The HSS steel was reported as CSA Grade 40.21- 350W, however its

measured yield stress and tensile strength were not reported. Because of this, the TO failure

predictions have been omitted in Table 4.3 and the CF comparison has been made based on

the efficiency factors (U) reported by Korol and those predicted by design provisions.

Although the CSA design provision provided a better predicted connection strength than

AISC, and even with a reduced eccentricity, the efficiency attained for specimens failing by CF

always surpassed these predicted values. The causes producing these results have been

related to specific factors affecting each design provision. Even though CSA generally provides

more conservative values than AISC, this was not evident herein. As specimens 1a to 3b had

Lw/w ratios close to 1.0, CSA assigned them an efficiency factor near 0.90. (CSA assigned a

SpecimenTest

Capacity Nux [kN]

Failure Mode a)

Nux/AnFu (An=Ag)b)

Circumferential Tensile Fracture Block Shear

AISC (2005)

using Nu [kN]

Nux/ Nu

AISC (2005)

using Nu [kN]

Nux/ Nu

CSA (2001) Nu [kN]

Nux/ Nu

Packer & Henderson

(1997) Nu [kN]

Nux/ Nu

AISC (2005) &

CSA (2001)Nu [kN]

Eurocode (CEN 2005)

Nu [kN]

C-Sep-1 256 CF 0.90 285 0.90 285 0.90 191 1.34 177 1.45 232 225C-Sep-2 326 CF 0.90 362 0.90 362 0.90 246 1.32 224 1.45 315 305C-Sep-3 371 CF 0.89 416 0.89 416 0.89 285 1.30 312 1.19 355 344C-Sep-4 522 CF 0.93 561 0.93 561 0.93 363 1.44 348 1.50 420 406C-Sep-5 652 CF 0.85 763 0.85 763 0.85 499 1.31 473 1.38 583 565C-Sep-6 795 CF 0.84 952 0.84 952 0.84 631 1.26 590 1.35 811 784S-Sep-2 274 CF 0.94 223 1.23 251 1.09 188 1.46 181 1.52 291 282S-Sep-3 505 CF 0.96 393 1.29 462 1.09 336 1.50 325 1.56 524 508S-Sep-4 478 CF 0.85 435 1.10 472 1.01 365 1.31 348 1.37 475 460S-Sep-5 833 CF 0.94 686 1.21 759 1.10 587 1.42 549 1.52 890 861S-Sep-6 949 CF 0.90 805 1.18 911 1.04 680 1.40 650 1.46 935 906R-Sep-3 475 CF 0.89 421 1.13 492 0.96 347 1.37 332 1.43 576 559R-Sep-5 384 CF 0.97 292 1.32 329 1.17 258 1.49 245 1.57 406 393R-Sep-8 711 CF 0.72 815 0.87 914 0.78 633 1.12 610 1.17 970 939R-Sep-9 913 CF 0.87 867 1.05 1012 0.90 665 1.37 786 1.16 1085 1053R-Sep-10 560 NK 1.00 409 1.37 437 1.28 363 1.54 346 1.62 509 492

Mean 1.07 0.98 1.37 1.42COV (%) 17.3 13.8 7.7 10.0

x x'


4-4

low efficiency to only specimen 5a which had a Lw/w ratio near 0.60). On the other hand, AISC

does not allow full efficiency attainment for RHS and SHS, as it does for CHS. Because of this,

the efficiency factors by AISC are lower than CSA values, for ratios Lw/w >1.0. Even though

consideration of the plate thickness can improve the calculation of AISC efficiency factors, the

AISC U’ values generally remained below the CSA U values. The influence of the gusset plate

orientation on connection efficiency (Ureal) of RHS was found to be very minor in the test, but

AISC predicted efficiency differs when efficiency factors for specimens 1a vs 3b and 1b vs 3a

are compared.

Table 4.3 Actual and predicted connection efficiency for Korol et al. (1994) data

a) CF represents circumferential fracture.

4.3 Experimental program by Zhao and Hancock (1995)

Zhao and Hancock tested 48 slotted tube connections fabricated using SHS and RHS

with wall thicknesses varying from 1.6 to 3.0 mm. This program was aimed at assessing the

accuracy of then-current design equations for welds in sections less than 3mm thick and

determining the influence that the gusset plate orientation and weld length may have on the

failure. Hence, the weld length in all the specimens was chosen to prevent CF. (These

specimens had Lw/w ratios ranging from 0.37 to 0.55). Even though all the test specimens failed

by TO failure, Table 4.4 shows a comparison between the test results and the predicted CF and

TO capacities according to current codes. Packer and Henderson (1997) have also suggested

that TO failure will govern for ratios Lw/w < 0.6.

Specimen

Test Capacity

Nux [kN]

Failure Modea) Ureal

Circumferential Tensile Fracture

AISC (2005)

using [U]

Ureal/ U

AISC (2005)

using [U']

Ureal/ U’

CSA (2001)

[U]

Ureal/ U

Packer & Henderson

(1997)

[U]

Ureal/ U

1a 811 CF 0.98 0.75 1.31 0.80 1.23 0.89 1.10 0.75 1.311b 836 CF 1.04 0.74 1.40 0.80 1.31 0.87 1.19 0.62 1.682a 664 CF 0.95 0.79 1.20 0.84 1.13 0.86 1.10 0.62 1.532b 725 CF 1.00 0.80 1.26 0.85 1.18 0.89 1.13 0.75 1.333a 845 CF 1.05 0.86 1.22 0.91 1.15 0.86 1.23 0.62 1.693b 854 CF 1.05 0.87 1.21 0.92 1.15 0.88 1.19 0.75 1.405a 612 CF 0.88 0.66 1.33 0.74 1.19 0.75 1.17 0.62 1.42

Mean 1.28 1.19 1.16 1.48COV(%) 5.8 5.2 4.3 10.6

x x'


4-5

As shown in Figure 4.4, AISC (2005) predicts TO as the governing failure mechanism for

42 out of 48 specimens. Moreover, the consideration of the plate thickness in the AISC model

would predict solely TO failures. On the other hand, the small ratios Lw/w values herein (ranging

from 0.37 to 0.55) negatively impact the CSA predictions for CF. As a result, CSA predicts CF

as the governing failure mechanism for all tests, which clearly contradicts the test results.

Since the weld sizes are not reported by the authors, TO failure predictions have been

included herein using a weld leg size of zero. Among North American and Eurocode design

provisions, where all have underestimated the connection strength for the check of TO failure, a

slightly better prediction can be obtained by AISC and CSA compared to Eurocode (slightly

lower mean and similar COV). The use of a hypothetical weld leg size of 3mm in all the

specimens would improve these predictions to produce a mean of 1.18, but still conservative

relative to the real test values.


4-6

Table 4.4 Actual and predicted connection strength for Zhao and Hancock (1995) data

a) TO represents block shear tear-out failure along the weld;b) As slot end was welded, it is likely appropriate to assume that An = Ag.

Spec

imen Test

Capacity Nux [kN]

Failure Mode a) Nux/AnFu Nux/AgFu


AISC (2005) using Nu [kN]


CSA (2001)

Nu [kN]

AISC (2005) & CSA (2001) Nu [kN]

Nux/ Nu

Eurocode (CEN 2005)

Nu [kN]

Nux/ Nu

R1ENS 228 TO 0.64 0.59 195 224 128 177 1.29 171 1.34R1ENL 264 TO 0.74 0.68 225 248 157 217 1.21 209 1.26R2ENS 187 TO 0.58 0.54 174 200 113 159 1.18 153 1.22R2ENL 198 TO 0.62 0.57 187 211 124 175 1.13 168 1.18R3ENS 168 TO 0.71 0.63 119 150 85 116 1.45 111 1.51R3ENL 173 TO 0.73 0.64 133 160 96 131 1.32 126 1.37R4ENS 140 TO 0.67 0.59 106 134 74 105 1.33 102 1.38R4ENL 161 TO 0.77 0.67 128 150 93 133 1.21 128 1.26R5ENS 102 TO 0.61 0.54 85 107 58 82 1.24 79 1.29R5ENL 132 TO 0.79 0.69 105 121 76 109 1.22 104 1.26R6ENS 78 TO 0.61 0.54 64 81 43 61 1.28 58 1.34R6ENL 78 TO 0.61 0.54 71 86 49 69 1.14 66 1.18R7ENS 185 TO 0.61 0.56 136 167 104 148 1.25 143 1.30R7ENL 209 TO 0.69 0.63 164 190 125 179 1.17 173 1.21R8ENS 153 TO 0.63 0.57 110 135 83 116 1.31 112 1.36R8ENL 188 TO 0.77 0.70 141 160 108 152 1.24 146 1.29R9ENS 123 TO 0.60 0.55 93 114 68 94 1.31 90 1.36R9ENL 131 TO 0.64 0.58 105 124 77 106 1.24 102 1.29R7WNS 199 TO 0.66 0.60 175 207 104 148 1.34 143 1.39R7WNL 243 TO 0.81 0.73 197 223 125 179 1.35 173 1.41R8WNS 173 TO 0.71 0.64 142 167 83 116 1.49 112 1.54R8WNL 220 TO 0.90 0.82 165 185 108 152 1.45 146 1.51R9WNS 127 TO 0.62 0.56 119 141 68 94 1.35 90 1.41R9WNL 139 TO 0.68 0.62 129 148 77 106 1.32 102 1.37R1EYS 262 TO 0.68 b) 211 243 128 205 1.28 198 1.32R1EYL 309 TO 0.80 b) 244 269 157 245 1.26 237 1.31R2EYS 235 TO 0.68 b) 189 217 113 184 1.28 178 1.32R2EYL 251 TO 0.72 b) 203 228 124 199 1.26 192 1.31R3EYS 203 TO 0.76 b) 136 171 85 146 1.39 141 1.44R3EYL 224 TO 0.83 b) 151 183 96 161 1.39 156 1.44R4EYS 174 TO 0.73 b) 121 152 74 131 1.32 127 1.37R4EYL 209 TO 0.87 b) 145 170 93 159 1.31 154 1.36R5EYS 133 TO 0.70 b) 96 121 58 103 1.29 100 1.33R5EYL 169 TO 0.89 b) 119 138 76 129 1.31 125 1.36R6EYS 103 TO 0.72 b) 72 91 43 76 1.36 74 1.40R6EYL 111 TO 0.77 b) 81 97 49 84 1.32 81 1.37R7EYS 215 TO 0.65 b) 150 185 104 177 1.21 172 1.25R7EYL 243 TO 0.73 b) 182 210 125 208 1.17 201 1.21R8EYS 190 TO 0.71 b) 121 149 83 139 1.36 135 1.41R8EYL 228 TO 0.85 b) 156 177 108 175 1.31 169 1.35R9EYS 154 TO 0.68 b) 102 125 68 113 1.36 109 1.41R9EYL 170 TO 0.75 b) 116 137 77 125 1.36 121 1.41R7WYS 240 TO 0.72 b) 194 228 104 177 1.35 172 1.40R7WYL 268 TO 0.81 b) 218 246 125 208 1.29 201 1.33R8WYS 184 TO 0.69 b) 143 175 74 127 1.45 123 1.49R8WYL 248 TO 0.92 b) 182 204 108 175 1.42 169 1.47R9WYS 165 TO 0.73 b) 131 155 68 113 1.46 109 1.51R9WYL 173 TO 0.77 b) 140 162 76 123 1.41 119 1.46

Mean 1.31 1.35COV (%) 6.6% 6.5%

x x'


4-7

4.4 Experimental program by Cheng et al. (1996)

Cheng et al. tested nine specimens fabricated with CHS and all but one were reinforced

with a transverse weld (thus, An became equal to Ag). Failure of two of these specimens was by

CF at the connection region and the others developed a neck at their mid-length. The CHS that

developed a neck had Lw/w ratios ranging from 1.00 to 1.14 (as well as a Lw/D>1.3) plus a

transverse weld. One specimen, pwc1, which lacked a transverse weld and had a slot, had a

ratio of Lw/w=1.14 and failed by CF through its net section (see Table 4.5). Nevertheless, this

specimen was able to develop its full efficiency (100% of AnFu). The addition of a transverse

weld on specimen spec2 allowed the attainment of its full efficiency (100% of AgFu). However,

the relatively low ratio for Lw/w of 0.85 increased the presence of shear lag in this specimen

triggering fracture near the weld region and consequently a CF failure mode.

Table 4.5 shows a comparison between the test results and predictions from current

codes/guides. The predicted failure mechanism based on current design provisions is CF, since

the predicted values for TO failure always surpassed the CF predictions. This result is due to

the connection type tested herein and the large Lw/w and Lw/D ratios for these test specimens.

During the CF check by the AISC (2005) design provision, an efficiency factor of U=1.0

has been used for specimens with Lw/D > 1.3. Because of this, the only improvement based on

the use of a reduced eccentricity ( ) can be seen in specimen spec2. Nevertheless, in both

cases the AISC design provision underestimates the connection strength for this specimen. On

the other hand, the predicted connection strength by CSA (2001) was always much more

conservative (with a mean of 1.36). Finally, the use of the efficiency factors suggested by

Packer & Henderson (1997) results in extremely conservative predictions.

For the TO failure predictions, Eurocode, AISC and CSA design provisions showed

similarities in their predicted values and the main difference is in the calculation of the maximum

shear stress, since it is calculated as Fy/ by Eurocode and 0.6 Fy by AISC and CSA.

x'

3


4-8

Table 4.5 Actual and predicted connection strength for Cheng et al. (1996) data

a) CF represents circumferential fracture and NK represents necking of the CHS;b) As slot end was welded, it is likely appropriate to assume that An = Ag.

4.5 Experimental program by Zhao et al. (1999)

In a similar manner to the study previously undertaken by Zhao and Hancock (1995),

Zhao et al. (1999) tested 24 slotted tube connections fabricated with RHS but utilizing a high-

strength tube material (with a nominal Fy= 450 MPa) and for a second time wall thickness below

3mm. During this study, the influence of the gusset plate orientation, loading condition and the

connection detailing was studied. For all the test specimens, weld lengths (resulting in Lw/w

ratios ranging from 0.35 to 0.50) were designed to prevent a CF resulting in a generalized TO

failure. Nevertheless, these test results have been evaluated against the CF failure mode as

well. Table 4.6 shows a comparison between the test results and predictions from current

codes. Since small ratios Lw/w were used herein, Packer and Henderson (1997) would also

suggest that TO governs.

For these tests, AISC predicts CF failure as the governing failure mechanism (for 18 out

of 24 specimens were governed). Moreover, the small Lw/w values herein (ranging from 0.35 to

0.50) negatively affected the CSA predictions. CSA predicts CF as the governing failure

mechanism for all specimens. In both cases, these predictions clearly contradict the test results.

On the other hand, the consideration of the plate thickness in the AISC model improve its

predictions considerably, resulting in solely TO failure predictions.

Spec

imen Test

Capacity Nux [kN]

Failure Modea)

Nux/ AnFu

Nux/ AgFu


AISC (2005)

using Nu [kN]

Nux/ Nu

AISC (2005)

using Nu [kN]

Nux/ Nu

CSA (2001)

Nu [kN]

Nux/ Nu

Packer & Henderson

(1997) Nu [kN]

Nux/ Nu

AISC (2005) &

CSA (2001) Nu [kN]

Eurocode (CEN

2005) Nu [kN]

pwc1 830 CF 1.06 0.98 781 1.06 781 1.06 666 1.25 586 1.42 1081 1043pwc2 869 Neck 1.02b) 849 1.02 849 1.02 666 1.30 637 1.36 1138 1100pwc3 849 Neck 1.00b) 849 1.00 849 1.00 666 1.28 637 1.33 1138 1100pwc4 875 Neck 1.03b) 849 1.03 849 1.03 666 1.31 637 1.37 1138 1100pwc5 645 Neck 1.03b) 624 1.03 624 1.03 459 1.40 468 1.38 693 670pwc6 634 Neck 1.02b) 624 1.02 624 1.02 459 1.38 468 1.36 693 670pwc7 631 Neck 1.01b) 624 1.01 624 1.01 459 1.37 468 1.35 693 670spec1 2160 Neck 1.01b) 2141 1.01 2141 1.01 1598 1.35 1606 1.35 2339 2258spec2 2157 CF 1.01b) 1596 1.35 1674 1.29 1325 1.63 1327 1.63 1904 1839

Mean 1.06 1.05 1.36 1.39COV (%) 10.4 8.6 8.2 6.5

x x'


4-9

The use of small Lw/w ratios produced a comparable outcome as previously seen by Zhao

and Hancock (1995). The best prediction for TO failure (using a weld leg size of zero as weld

sizes are not reported by the authors) corresponded to North American design provisions with a

mean of 1.30. Once more, the hypothetical use of a weld leg size of 3mm in all the specimens

would improve these predictions to a mean of 1.16. Nevertheless, these predictions would still

remain below the real test values in all cases.

Table 4.6 Actual and predicted connection strength for Zhao et al. (1999) data

a) TO represents block shear tear-out failure along the weld;b) As slot end was welded, it is likely appropriate to assume that An = Ag.

SpecimenTest

Capacity Nux [kN]

Failure Modea)

Nux/ AnFu

Nux/ AgFu



AISC (2005)

using Nu [kN]

CSA (2001) Nu [kN]

AISC (2005) &

CSA (2001) Nu [kN]

Nux/ Nu

Eurocode (CEN 2005)

Nu [kN]

Nux/ Nu

G1EY1 183 TO 0.73 b) 133 165 82 146 1.25 142 1.29

G1EY2 185 TO 0.74 b) 133 165 82 146 1.26 142 1.30

G1EY3 192 TO 0.77 b) 133 165 82 146 1.31 142 1.35

G1EN1 140 TO 0.63 0.56 117 145 82 119 1.18 115 1.22

G1EN2 161 TO 0.73 0.64 117 145 82 119 1.35 115 1.41

G1EN3 148 TO 0.67 0.59 117 145 82 119 1.24 115 1.29

G2EY1 171 TO 0.57 b) 102 139 76 134 1.27 130 1.31

G2EY2 177 TO 0.59 b) 102 139 76 134 1.32 130 1.36

G2EY3 167 TO 0.56 b) 102 139 76 134 1.24 130 1.28

G2EN1 143 TO 0.53 0.48 93 126 76 109 1.31 105 1.37

G2EN2 141 TO 0.52 0.47 93 126 76 109 1.30 105 1.35

G2EN3 141 TO 0.52 0.47 93 126 76 109 1.30 105 1.35

G2WY1 174 TO 0.58 b) 149 186 76 134 1.30 130 1.34

G2WY2 175 TO 0.59 b) 149 186 76 134 1.30 130 1.35

G2WY3 169 TO 0.57 b) 149 186 76 134 1.26 130 1.30

G2WN1 140 TO 0.52 0.47 135 169 76 109 1.29 105 1.34

G2WN2 139 TO 0.51 0.47 135 169 76 109 1.28 105 1.33

G2WN3 140 TO 0.52 0.47 135 169 76 109 1.29 105 1.34

G3EY1 216 TO 0.76 b) 150 186 93 165 1.31 160 1.35

G3EY2 231 TO 0.82 b) 150 186 93 165 1.40 160 1.44

G3EY3 221 TO 0.78 b) 150 186 93 165 1.34 160 1.38

G3EN1 184 TO 0.74 0.65 132 163 93 134 1.37 129 1.43

G3EN2 192 TO 0.77 0.68 132 163 93 134 1.43 129 1.49

G3EN3 183 TO 0.74 0.65 132 163 93 134 1.36 129 1.42

Mean 1.30 1.35

COV(%) 4.2 4.3

x x'


4-10

4.6 Experimental program by Wilkinson et al. (2002)

Wilkinson et al. tested eight specimens fabricated with RHS having a nominal size of

100x50mm. These specimens had differences in their weld lengths, weld sizes, gusset plate

orientation and wall thicknesses (having thickness of 4 and 6 mm). They were also fabricated

with end return welds, hence An=Ag. Tests had either weld metal failure or a base metal failure.

These connections had very small Lw/w ratios (ranging from 0.33 to 0.64) that clearly

encourage the presence of a TO failure. However, a lack of proper fusion between the weld

material and the HSS in these tests led to a failure mechanism involving the welds. Even though

this premature failure mode recurred during the tests, the strength from several connections

exceeded the predicted load by prediction equations, which may suggest that those equations

are too conservative. Since the failure mechanisms seen during this experimental program do

not correspond to a CF or a clear TO failure, a comparison of this data with current design

provisions (such as AISC, CSA and Eurocode) specific to these failure mechanisms has been

omitted herein.

4.7 Experimental program by the Authors

The test specimens fabricated for this experimental program were the largest scale done

to date and thus more representative of real structural members. In addition, the weld lengths

used in these specimens produced Lw/w ratios ranging from 0.62 to 0.87 which ensured the

presence of shear lag during the tests and predominantly CF failures.

The predicted governing failure mechanism based on current codes/guides (AISC, CSA

and Packer & Henderson) was always a CF, since the predicted values for TO failure always

exceeded the CF values (see Table 4.7). Nevertheless, this contradicts the result from test

specimen B1 where a clear TO failure occurred. The CF failure mode results showed better

agreement with the predictions made by AISC than CSA. Moreover, the use of a reduced

eccentricity with the AISC model improved these predictions (since the mean actual-to-

predicted value was decreased from 1.32 to 1.17 and the COV also decreased). Despite this

upgrading, the modified AISC ( ) model still underestimates the real CF connection capacity. In

any case, the use of the AISC provision offers a better result than the CSA provision which

underestimated the connection capacity (resulting in a mean actual-to-predicted value of 1.61).

The use of the efficiency factor by Packer & Henderson (1997) provides a better prediction than

CSA. This can be explained by the manner in which the efficiency factors are calculated. While

x'

x'


4-11

CSA (2001) provides a variable efficiency factor (based on the Lw/w ratio), which may be

excessively conservative for ratios near to 0.60, Packer & Henderson (who follow an earlier

CSA code version) use a constant efficiency factor in this range. Hence, codes/guides failed

herein to predict accurately the connection strength and their corresponding failure mechanism.

Table 4.7 shows a comparison between the test results and predictions from current codes/

guides. For TO failure, the predictions by all three design provisions were similar and

reasonably accurate. AISC and CSA showed a mean actual-to-predicted value of 0.95 and

Eurocode a mean value of 0.99.

Even though the presence of the shear lag phenomenon in connections under

compression loading is completely neglected by design provisions, the specimens tested in

compression (both having very similar weld lengths) failed by local buckling at the connection

region under the influence of this phenomenon. Hence, a more extensive study of shear lag

under compression loading needs to be undertaken in order to provide a clear understanding of

this behaviour.

Table 4.7 Actual and predicted connection strength from University of Toronto data

a) CF represents circumferential fracture, TO represents block shear tear-out failure along the weld and LB represents localbuckling at the connection end;

b) Specimen was loaded in compression;c) As slot end was welded, it is likely appropriate to assume that An = Ag.

Spec

imen

Test

Cap

acity

N

ux [k

N]

Failu

re

Mod

ea) Nux/ AnFu

Nux/ AgFu


AISC (2005)

using Nu [kN]

Nux/Nu

AISC (2005)

using Nu [kN]

Nux/Nu

CSA (2001)

Nu [kN]

Nux/Nu

Packer & Henderson

(1997)Nu [kN]

Nux/Nu

AISC (2005) & CSA (2001) Nu [kN]

Nux/Nu

Eurocode (CEN

2005) Nu [kN]

Nux/Nu

A1 1032 CF,TO 0.87 0.77 782 1.32 880 1.17 618 1.67 739 1.40 1017 1.01 983 1.05

A2 1154 CF 0.97 0.86 859 1.34 939 1.23 760 1.52 739 1.56 1228 1185

A3C -1145 LB Nux/Csc=0.94 b)

B1 1087 TO 0.81c) 911 1013 669 827 1219 0.89 1181 0.92

B2 1211 CF 0.91c) 990 1.22 1073 1.13 824 1.47 827 1.46 1446 1401

C1 1107 CF 0.83 893 1.24 999 1.11 642 1.73 827 1.34 1230 1195

C2 1196 CF 0.90 968 1.24 1055 1.13 772 1.55 827 1.45 1423 1380

C3C -869 LB Nux/Csc=0.72 b)

E1 1109 CF 0.81 0.69 766 1.45 914 1.21 685 1.62 835 1.33 1034 1001

E2 1236 CF 0.90 0.76 898 1.38 1019 1.21 859 1.44 849 1.46 1269 1227

E3 1336 CF 0.83 925 1.44 1102 1.21 689 1.94 1004 1.33 1267 1234

E4 1400 CF 0.86 1040 1.35 1188 1.18 826 1.69 1004 1.40 1441 1401

E5 1282 CF 0.94 0.79 1069 1.20 1187 1.08 874 1.47 841 1.52 1299 1257

Mean 1.32 1.14 1.61 1.42 0.95 0.99

COV (%) 6.9 4.4 9.6 5.7 9.1 9.3

x x'


4-12

4.8 Experimental program by Ling (2005)

An experimental program on CHS to slotted gusset plate connections utilizing very high

strength tubes (VHS) has been undertaken by Ling. For this program, 45 connections were

fabricated with Lw/w ratios ranging from 0.17 to 1.37. In addition to the longitudinal fillet welds, a

weld return was applied to all the test specimens (hence, Ag=An). Of these specimens, 25 failed

by TO, 16 by CF and 4 had a combination of both (TO-CF).

The predicted failure mode based on current codes/guides (AISC, CSA and Packer &

Henderson) is always a CF, since the predicted values for TO failure always exceeded the CF

values (see Table 4.8). However, this is not in accord with the test results where 25 out of 45

specimens had TO failure. For TO failures, all design provisions have overestimated the

connection strength. Thus, codes/guides have failed herein to accurately predict the connection

strength and their corresponding failure mode. Table 4.8 shows a comparison between test

results and the predictions from current codes/guides.

In general, the results from these specimens have a lower efficiency when they are

compared to other experimental programs. Moreover, despite several connections having a Lw/

w ratio greater than 1.0, their maximum connection strength never exceeded a value of

0.70AnFu (where An=Ag), and thus the attainment of necking was not reported. These results

are related to the tube material properties used herein (which have extremely high Fy and Fu

values) and the undermatched weld metal, which likely diluted the material properties in the

Heat Affected Zone (HAZ). This resulted in premature material fracture there and hence a

generally lower connection efficiency.


4-13

Table 4.8 Actual and predicted connection strength for Ling (2005) data

a) CF represents circumferential fracture and TO represents block shear tear-out failure along the weld.b) As slot end was welded, it is likely appropriate to assume that An = Ag.

Spec

imen

Test

Cap

acity

N

ux [k

N]

Failure Modea)

Nux/AnFu

b) (An=Ag)


AISC (2005)

using Nu [kN]

Nux/Nu

AISC (2005) using

Nux/Nu

CSA (2001)

Nu[kN]

Nux/Nu

Packer & Henderson

(1997)Nu [kN]

Nux/Nu

AISC (2005) &

CSA (2001) Nu [kN]

Nux/Nu

Eurocode (CEN 2005)

Nu [kN]

Nux/Nu

TV1P15a 101.2 TO 0.44 75 152 55 156 0.65 153 0.66TV1P15b 103.8 TO 0.43 80 161 58 166 0.63 163 0.64TV1P30a 137.8 TO, CF 0.57 160 0.86 200 0.69 114 1.21 150 0.92 245 0.56 239 0.58TV1P30b 122.7 CF 0.53 154 0.80 193 0.64 111 1.11 144 0.85 237 231TV1P45a 151.3 CF 0.65 233 0.65 233 0.65 154 0.98 175 0.86 317 308TV1P45b 154.0 CF 0.65 238 0.65 238 0.65 158 0.98 179 0.86 324 315TV2P20a 126.0 TO 0.48 130 196 84 164 209 0.60 205 0.62TV2P20b 125.2 TO 0.49 126 190 81 159 202 0.62 198 0.63TV2P35a 162.6 CF 0.63 183 0.89 220 0.74 143 1.14 160 1.02 291 283TV2P35b 162.0 TO, CF 0.61 187 0.86 225 0.72 147 1.10 164 0.99 299 0.54 291 0.56TV2P50a 170.2 CF 0.66 258 0.66 258 0.66 177 0.96 194 0.88 378 367TV2P50b 167.5 CF 0.63 264 0.63 264 0.63 181 0.92 198 0.85 389 378VT3P25a 149.8 TO 0.50 176 236 119 184 261 0.57 255 0.59VT3P25b 151.1 TO 0.51 176 235 119 183 261 0.58 255 0.59TV3P40a 184.9 CF 0.57 243 0.76 284 0.65 209 0.89 202 0.92 392 382TV3P40b 180.4 CF 0.61 221 0.82 258 0.70 190 0.95 222 0.81 357 347TV3P55a 177.3 CF 0.58 305 0.58 305 0.58 220 0.81 229 0.77 464 450TV3P55b 175.4 CF 0.60 294 0.60 294 0.60 212 0.83 221 0.79 448 435TV4P25a 120.7 TO 0.40 154 214 98 186 222 0.54 217 0.56TV4P25a 130.1 TO 0.45 148 206 94 179 213 0.61 208 0.63TV4P40a 162.2 TO, CF 0.54 208 0.78 246 0.66 157 1.03 185 0.87 306 0.53 297 0.55TV4P40b 170.4 TO 0.57 208 245 156 185 304 0.56 295 0.58TV4P55a 188.2 CF 0.64 292 0.64 292 0.64 197 0.96 219 0.86 377 366TV4P55b 195.4 CF 0.66 294 0.66 294 0.66 199 0.98 221 0.88 382 370TV5P25a 146.5 TO 0.48 158 219 101 191 231 0.63 225 0.65TV5P25b 142.8 TO 0.47 155 215 99 187 226 0.63 220 0.65TV5P40b 182.5 TO 0.59 214 252 161 191 318 0.57 309 0.59TV5P55a 213.2 CF 0.69 309 0.69 309 0.69 209 1.02 232 0.92 404 392TV6P35a 193.5 TO 0.54 232 283 164 221 334 0.58 325 0.60TV6P35b 194.5 TO 0.56 227 276 160 215 326 0.60 317 0.61TV6P50a 230.5 CF 0.66 348 0.66 348 0.66 229 1.01 261 0.88 423 411TV6P50b 229.1 CF 0.66 347 0.66 347 0.66 229 1.00 261 0.88 423 411TV6P65a 238.8 CF 0.69 347 0.69 347 0.69 251 0.95 260 0.92 521 505TV7W20a 113.7 TO 0.26 38 147 75 187 0.61 183 0.62TV7W20b 113.2 TO 0.26 39 149 76 191 0.59 187 0.61TV7W35a 141.7 TO 0.34 202 262 127 260 0.54 253 0.56TV7W35b 155.2 TO 0.35 212 275 133 273 0.57 266 0.58TV7W50a 178.4 TO 0.41 276 319 186 269 349 0.51 339 0.53TV7W50b 188.4 TO, CF 0.43 275 0.68 319 0.59 185 1.02 269 0.70 347 0.54 337 0.56TV8W18a 113.2 TO 0.20 65 170 0.66 167 0.68TV8W18b 109.0 TO 0.19 66 172 0.63 169 0.65TV8W28a 135.5 TO 0.25 81 180 101 219 0.62 214 0.63TV8W28b 136.8 TO 0.25 79 178 101 219 0.62 214 0.64TV8W38a 159.2 TO 0.28 210 285 140 277 0.58 269 0.59TV8W38b 160.6 TO 0.29 207 281 138 273 0.59 266 0.60

Mean 0.71 0.66 0.99 0.87 0.59 0.60COV (%) 13.1 6.1 9.8 8.1 6.5 6.3

xx'


4-14

4.9 Summary of Chapter 4

In general, the application of current code TO failure equations can provide a good

prediction of connection capacity when the test specimen Lw/w ratio moves towards a value of

0.6, but a clear overestimation will occur when this ratio approaches 1.0. On the other hand, the

current code CF prediction equations have shown better correlation to tests when the

connection Lw/w ratio is near to 1.0, and a gradual underestimation of connection capacity as

this ratio is decreased.

In most cases, the predicted failure mechanism (and its corresponding maximum

connection strength predicted by these design provisions) did not agree well with the

experimental results in the Lw/w range from 0.6 to 1.0. Moreover, the results from Zhao and

Hancock (1995) and Zhao et al. (1999) may encourage the idea that a TO failure always

governs for Lw/w ratios below 0.6, but the CF failure check must be avoided in that range since

it may falsely be found to control. Hence, it seems necessary to develop a comprehensive

model that can provide a better connection strength prediction over a wide parameter range.

Finally, it is noted that the use of the experimental results from Ling (2005) may not be

valid for regular strength structural hollow section connections since these results have shown a

considerable scatter relative to the rest of the test programs.


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