Experimental Work on Cold-Formed Steel Elements for Earthquake Resilient Moment Frame Buildings

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Experimental work on cold-formed steel elements for earthquake resilient

moment frame buildings

Alireza Bagheri Sabbagh a,, Mihail Petkovski a, Kypros Pilakoutas a, Rasoul Mirghaderi b

a Department of Civil and Structural Engineering, University of Sheffield, Sheffield, UKb School of Civil Engineering, College of Engineering, University of Tehran, Tehran, Iran

a r t i c l e i n f o

Article history:

Received 23 November 2011

Revised 17 April 2012

Accepted 26 April 2012

Available online 7 June 2012

Keywords:

Cold-formed steel sections

Beam-column connections

Moment-resisting frames

Seismic behaviour

a b s t r a c t

This paper presents an experimental investigation on the use of thin-walled cold-formed steel (CFS) sec-

tions as energy dissipative elements for earthquake resistant moment frame multi-storey buildings. The

tests were performed on six bolted beam-to-column connections, using through plates and curved flange

beams with different types of out-of-plane stiffeners in the connection region. The hysteretic behaviour

of the CFS connections shows high seismic energy dissipation capacity and sufficient ductility to satisfy

code requirements for seismic design. The use of out-of-plane stiffeners inside the beams in the connec-

tion region results in improvement of the momentrotation behaviour of the connection by up to 35% in

strength and 75% in ductility. Mobilising connection slip after the elastic cycles provides highly stable

hysteretic behaviour and an increase of up to 240% in energy dissipation capacity. The tested connections

can be classified as rigid with partial or full strength depending on the connection stiffeners.

2012 Elsevier Ltd. All rights reserved.

1. Introduction

Experimental work on monotonic and cyclic behaviour of com-

ponents and elements of CFS moment-resisting frames (MRFs) is

very limited [14]. Premature local failures are prevalent in com-

mon CFS sections because of their thin-walled elements. One solu-

tion to avoid premature local failures and to provide ductility in

bolted CFS connections is to mobilise slip and bearing action of

bolts while beams and columns remain elastic [1]. This limited

source of ductility restricts such structures to one-storey dwellings

[1]. For seismic design of multi-storey buildings, there is a need to

dissipate large amount of energy through plasticity in the beams

rather than just yielding the material around the bolt holes.

There are research studies [2,3] showing that by using appropri-

ate connection details for CFS beam-column connections, such as

gusset plates, relatively high moment resistance can be developed

in CFS double back-to-back channel sections. In this type of beam-

to-column connection however, no ductile capacity was achieved

after reaching the peak bending moment. The general assumption

is that CFS beams with thin-walled elements cannot develop plas-

tic hinges, thus cannot be used for high seismicity areas [57].

In a recent study by the authors [4] conventional double back-

to-back channel beam sections integrated within topping concrete

were shown to possess a degree of ductile capacity in dissipating

seismic energy by achieving rotations larger than 0.04 rad; satisfy-

ing the requirements for special moment frames [7]. However,they did not satisfy the required width/thickness limits of design

codes [57] which aim to delay the local buckling after yielding.

This research also showed encouraging results for through plate

type of CFS beam-to-column connections [4].

If CFSbeams are designedas themainenergy dissipation compo-

nents in seismic resistant MRFs, the ductility capacity of the beams

with thin-walledelements must be improved. Thefirst step is to de-

lay local buckling as much as possible to enable plastic deforma-

tions. Curved flange beam sections were developed by the authors

[810] by introducing more bends in the flanges (Fig. 1a), a step-

by-step process that ultimately led to significant increase in mo-

ment resistance, stiffness and ductility, compared with flat flange

beams.

In this study a web bolted moment resistant type of connection

is used for CFS beam-column connections. This type of connection

has already been examined as lapped connections in portal frames

[11,12]. The main components of the beam-column MR connection

are welded cross through-plates which can be bolted to separate

beam and column sections, as shown in Fig. 1b. Previous research

[810] has shown that web bolted through-plate beam-to-column

connection produced a lower level of ductility and strength than

that of a theoretical fixed-end beam. Web buckling adjacent to

the first line of bolts at the beam-through plate connection was

identified as the main reason for premature loss of strength [810].

An optimum combination of vertical and horizontal out-of-

plane stiffeners has been identified for the web bolted CFS connec-

0141-0296/$ - see front matter 2012 Elsevier Ltd. All rights reserved.http://dx.doi.org/10.1016/j.engstruct.2012.04.025

Corresponding author. Address: Department of Civil and Structural Engineering,

University of Sheffield, Sir Frederick Mappin Building, Mappin Street, Sheffield S1

3JD, UK. Tel.: +44 (0)114 222 5724; fax: +44 (0)114 222 5700.

E-mail address: [email protected] (A. Bagheri Sabbagh).

Engineering Structures 42 (2012) 371386

Contents lists available at SciVerse ScienceDirect

Engineering Structures

j o u r n a l h o m e p a g e : w w w . e l s e v i e r . c o m / l o c a t e / e n g s t r u c t
http://dx.doi.org/10.1016/j.engstruct.2012.04.025mailto:[email protected]://dx.doi.org/10.1016/j.engstruct.2012.04.025http://www.sciencedirect.com/science/journal/01410296http://www.elsevier.com/locate/engstructhttp://www.elsevier.com/locate/engstructhttp://www.sciencedirect.com/science/journal/01410296http://dx.doi.org/10.1016/j.engstruct.2012.04.025mailto:[email protected]://dx.doi.org/10.1016/j.engstruct.2012.04.025


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tions to increase both the strength and ductility [810]. Minimum

number of vertical stiffeners has also been identified for easier

implementation and for connections where ductility demand canbe met by a reduced number of stiffeners rather than a full (opti-

mum) set of stiffeners [9,10].

The experimental study presented in this paper investigates the

concept of CFS thin-walled curved flange sections as seismic en-

ergy dissipative elements for moment frame buildings using full

and minimum sets of out-of-plane stiffeners. It also aims to exam-

ine if curved flange CFS sections can produce full plastic moment

(Mp) sustained at large rotations similar to Class 1 cross sections

in Eurocode 3 [6] and larger than 0.04 rad required for special mo-

ment frames in AISC Seismic Provisions [7].

2. Testing arrangement and specimen details

Two specimen types with different thickness (A) 3 mm, and (B)

4 mm, with three different out-of-plane stiffener configurations

(A1, A2, A3, and B1, B2, B3) were used in the experimental investi-

gation (Table 1). The nominal dimensions of the components of the

test specimens are shown in Fig. 2.

According to the FE analysis presented previously by the

authors [810], the connections with full stiffeners (used for Spec-

imens A3 and B3) produce a significant increase in both strength

($40%) and ductility ($100%) in comparison with the connections

without stiffeners (Specimens A1 and B1) which were used for

bench marking and comparison purposes. The difference between

the boundary conditions of FE models and the test specimens is

that hot-rolled back-to-back channels were used in the tests in-

stead of CFS columns. This was easier to manufacture and installinto the testing rig. This change was supported by the results of

the FE analyses, which showed no yielding or large deformation

in the CFS columns [9,10].

For all specimens, the connections were designed using therequirements for slip-critical joints given in AISC Specification for

Structural Joints [13]. The design slip resistances of the farthest

bolts (Rn) in the connections were calculated by assuming slip coef-

ficient ofl = 0.5 for uncoated blast-cleaned steel [13] and applyinga pretension force Tm = 67 kN for the beam-to-through plate (BT)

and Tm = 53 kN for the through plate-to-column (TC) connections,

approximately equal to 42% of the tensile strength of the bolts (60%

of the 70% tensile strength, given in the Specification [13]). This re-

sulted in design resistance (Rn), higher than the required resis-

tances (Rreq) for specimens type A, but just below the required

resistances (Rreq) for specimens type B. Therefore, it was expected

that if the beamplastic moment was mobilised in the tests of spec-

imens type B, slipping at the connections would be triggered.

2.1. Test set-up

Fig. 3 shows a drawing and a photo of the test set-up. A brief

description of the design specifications for the test set-up compo-

nents are given in Appendix A.

2.2. Instrumentation

Strain gauges (SGs), inclinometers and LVDTs were placed at

different locations of the through plates, beams and columns as

(a) (b)

Fig. 1. (a)Step-by-step development of curvedflange sections and(b) CFSbeam-columnconnections: diamond column, cross through-plates andcurved flange beam[810].

Table 1

The specimens configurations.

Specimens Beam thickness

(mm)

Connection stiffeners Connection

type

A1 3 No stiffeners Slip-critical

A2 3 Partial (minimum)

stiffeners

Slip-critical

A3 3 Full (optimum) stiffeners Slip-critical

B1 4 No stiffeners Slip-critical

B2 4 Partial (minimum)

stiffeners

Slip-critical

B3 4 Full (optimum) stiffeners Slip-critical

Fig. 2. Dimensions and configuration of the test specimens.

372 A. Bagheri Sabbagh et al./ Engineering Structures 42 (2012) 371386


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shown in Fig. 4 to measure stresses and deformations at all the

critical parts of the specimens.

2.3. Loading protocol

Cyclic loading was applied through a hinge connection at the

end of the beam (Fig. 3) using a loading protocol (Fig. 5) given in

section S6.2 of AISC Seismic Provisions [7] for qualifying beam-col-umn moment connections in special and intermediate moment

frames. The centre of the plastic hinge region, used for calculating

the bending moment, M, and the rotation, h of the beams, is as-

sumed to be at the end of the through plate (Fig. 5). The distance

between this section and the loading point (1811 mm) was used

to determine the displacement of the actuator (Fig. 5) for a given

value ofh.

3. Test results

The momentrotation (Mh) behaviour, connection rigidity and

strain distribution results for all the specimens are presented in the

following subsections. The normalised moment (M/Mp) is shown

against the applied h. Mp is the nominal plastic moment of the

beam sections: 67 kN m for Specimens A and 90 kN m for Speci-

mens B, all assumed with nominal yield stress fy = 275 MPa. The

actual yielding stresses of the beams based on the tensile test re-

sults for Specimens A and B were 310 MPa and 320 MPa, respec-

tively. Therefore, the actual plastic moment strength of the

beams is expected to be 75 kN m for Specimens A and 105 kN m

for Specimens B. It should be noted that in the design of beam-col-

umn connections this difference is accounted for by using an over-

strength factor [7].

3.1. Specimens A13 and B1: connections dominated by rotation in the

beams

In Specimens A13 and B1 the rotation behaviour is dominated

by flexural and local buckling deformation in the beams. From the

M/Mph curves (Fig. 6), different regions corresponding to different

aspects of behaviour can be identified: (i) Elastic region (AB): Points

B and B correspond to the beginning of the inelastic region, (ii)

Inelastic region (BC): Points C and C correspond to the maximum

bending moments in the beam, (iii) Postbuckling region (CD): Points

DandD correspond to 80% of the maximum moment and (iv) Fail-

ure region (DE): Points E and E correspond to connection failure as

Fig. 3. Test set-up.

A. Bagheri Sabbagh et al. / Engineering Structures 42 (2012) 371386 373


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seen in Fig. 7. For a better understanding of the rotational behav-

iour of the connections of the specimens, videos of the connections

(Videos 14) that were made by framing the photos at every 60 s

are also presented.

The ductility factors based on the rotation ratios at points B and

D are 2.7, 4.0, 4.7 and 4.7 for Specimens A13 and B1, respectively.

The beams sustained 80% of the nominal plastic moment at a

rotation (at Point D) larger than 0.04 rad (for Specimens A23and B1) which is the rotation required for special moment frames

(SMFs) [7].

3.1.1. Hysteretic momentrotation behaviour of Specimen A1

The web buckling in Specimen A1 beam was initiated at

h = 0.03 rad rotation at the end of the through plate and extended

to the flange at h = 0.04 rad (see Video 1). This web buckling was

also identified by the authors through a set of FE analyses [9] as

the main reason for an abrupt loss of strength. This confirms that

vertical stiffeners at this location should be used (similar to Spec-

imens A2 and B2).

In the failure region (DE) the web and flange buckling pro-

gressed further and the flanges at the end of the beam opened

up as seen in Fig. 7. The large failure deformation of the beamflanges and webs highlights the need for vertical stiffeners.

The local rotations and deformations at the connection region

versus the actuator load (F) are shown in Fig. 8. The total rotation

of the beam at the connection region was very small, as shown in

Fig. 8a. This was verified by measuring the deformations at the top

and bottom of the through plate (horizontal transducers 11 and 12

in Fig. 8b). Slight slip in the through plate-to-column (TC) connec-

tion during the last cycles (see Fig. 8b) can be due to out-of-plane

bending of the through plate (this is verified by strain measure-ments shown later, in Section 3.4.1). A small vertical deformation

of the column was also recorded (Fig. 8c) and this is attributed to

shear deformation of the column web.

It can be concluded that the rotation recorded at the end of the

beam (Fig. 6A1) represents the beam rotation capacity, without

significant contributions from the connection elements.


In the test of Specimen A2, the connection slipped at Point C, in

the reverse loading cycle of h = 0.03 rad, dropping from M/

Mp =1.03 to M/Mp =0.81 (Fig. 6A2). In the loading cycle of

h = 0.04 rad, the connection slipped at M/Mp = +0.54. By repeating

this cycle, the web buckling was triggered outside the connection

adjacent to the vertical stiffener and the flange buckling was initi-ated between the vertical stiffener and the lateral bracing (see Vi-

Fig. 4. Instrumentation sketches of the specimens.

Fig. 5. Loading cycles.

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deo 2). The maximum loads were slightly reduced due to the initi-

ation of local buckling in this cycle.

In the repeating cycle of 0.06 rad, the maximum load was re-

duced to M/Mp = 0.81, corresponding to Points D and D in

Fig. 6. Points E and E correspond to the rotation of 0.07 rad. In this

region the web and flange buckling deformation extended at both

sides of the vertical stiffener at the through plate end as shown in

Video 2 and Fig. 7.

The local rotations and deformations in the connection region

versus the applied load (F) are shown in Fig. 9. As expected, the

rotation at the column side of the connection was very small

(Fig. 9a). This confirms that the fixing of the specimens to the reac-

tion frame worked well.

The total rotation of the beam-to-through plate (BT) and the

through plate-to-column (TC) connections (identifying the first

slip and maximum points corresponding to those in Fig. 6A2) is

shown in Fig. 9b. The rotation of the TC connection was also mea-

sured by the horizontal transducers 11 and 12 (Fig. 9c). These

measurements show that the TC connection contributed substan-

tially in the total rotation of the connection. The total rotation

showed elastic cycles up to Point C where the first slip occurred.

The total rotation of the connection elements (Fig. 9b) shows a

significant contribution (0.0173 rad) to the global beam-end rota-

tions (Fig. 6A2). Most of the slip occurred at the stabilised slip

load, Fs, at around 20 kN (corresponding to M/Mp = 0.54).

Fig. 6. Momentrotation curves of Specimens A13 and B1.

Fig. 7. Failure deformations of Specimens A13 and B1 correspond to Point E.

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The maximum total rotation of the connection (h6 = 0.0173 rad)

and the maximum rotation of the TC connection, 0.0135 rad, were

close to the value of 0.0167 rad, calculated as rotation due to clear-

ance. This value was calculated by assuming a uniform rotation

and 2 mm clearance for the bolt holes divided by the farthest dis-

tances at each of the connections (2 mm/300 mm BT connec-

tion + 2 mm/200 mm TC connection = 0.0167 rad). This showsthat very little bearing action was activated in the connections.

The slip of the bolts is not symmetric as can be seen in Fig. 9b.

This leads to different maximum strain values at the top and

bottom flanges.


The flange buckling of Specimen A3 was initiated in theh = 0.05 rad cycle, at the edge of the flanges between the 2nd, 3rd

Fig. 8. Specimen A1: local deformations in elements of connection.




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and the 4th vertical stiffeners, and extended in the subsequent cy-

cles (see Video 3).

At h = 0.07 rad, the maximum load dropped to M/Mp = +0.84 and

0.86 corresponding to Points D and D, respectively. At Points E

and E, corresponding to rotation of h = 0.08 rad, the specimen

failed due to rupture of the flanges following large beam deforma-

tion (see Video 3 and Fig. 7).

During the test, the vertical stiffeners were deformed out-of-

plane due to restraining forces from the horizontal stiffeners.

Therefore, both the vertical and horizontal stiffeners were effective

in restraining the web and flange buckling and resisted the local

buckling deformations.

The total rotation of the BT and the TC connections, measured

by inclinometer 6 is shown in Fig. 10a. The maximum total rotation

was 0.005 rad, part of which was due to the rotation in the TC

connection (shown in Fig. 10b). The other part was the rotation

in the BT connection. The rotation measured for the TC connec-

tion by the horizontal LVDTs 11 and 12 can be due to both slip of

the bolts and out-of-plane bending of the through plate (especially

towards the end of the test). No sudden slip was observed in the

connection (see Video 3), but clicking sounds were heard during

the test due to local slip of the bolts. This slip was mainly in the

first and last line of bolts of the BT connection. In this test, the slip

deformation was not sudden or significant. Therefore, the slip af-

fected neither the global behaviour (Fig. 6A3), nor the overall con-

nection behaviour (Fig. 10a).

3.1.4. Hysteretic momentrotation behaviour of Specimen B1

The web buckling in Specimen B1 was initiated in the beam at

h = 0.05 rad, at the critical section at the through plate end and ex-

tended in the subsequent cycles (see Video 4). At rotation of

h = 0.07 rad the maximum load dropped to M/Mp = 0.76 corre-

sponding to Points D and D (Fig. 6). During this cycle the web

buckling was intensified and it interacted with the flange buckling.

The specimen failed at Points E and E corresponding to rotation of

h = 0.08 rad (Fig. 6) due to extensive web and flange buckling

accompanied by a large deformation of the beam (Fig. 7).The total connection rotation measured by inclinometer 6 is

shown in Fig. 11a. The rotation of the TC connection (measured

by the horizontal transducers 11 and 12, Fig. 11b), was small. Dur-

ing the last cycles significant deformation occurred in the TC con-

nection mainly due to the out-of-plane deformation of the through

plate. The total rotation of the connection was mainly due to the

contribution of the BT connection. Most of the slip of the BT con-

nection occurred in the last cycles after the peak moment was

reached (Point C) and after the development of web buckling. This

may be a result of reduction of the frictional contact surfaces of the

first line of bolts due to the adjacent web buckling. Therefore, ver-

tical stiffeners at this location (similar to Specimens A2 and B2)

should be used to protect the frictional contact surfaces of the first

line of bolts which are the farthest from the centre of rotation and

more susceptible to slip.

Relatively small slip of the bolts was due to the fact that the

deformation demand was mostly met by the beam. The rotation of

the connection elements (Fig. 11) shows that their maximum con-

tribution to the beam-end rotation (Fig. 6B1) was only 0.015 rad.

3.2. Specimens B2 and B3: Connections dominated by rotational

behaviour produced by connection slip

From the momentrotation (M-h) curves (Fig. 12B2 and B3), dif-

ferent regions corresponding to different aspects of behaviour can

be identified: (i) Elastic region (AB), (ii) Prebuckling-slip region (BC)

and (iii) Postbuckling region (CD). The ductility factors, based on

the deformations at Point D (Specimen B2) or Point C (Specimen

B3) relative to Point B are 6.0 and 6.7, respectively, although the

load has not degraded at the end of the tests. The beams sustained

maximum strength up to a rotation significantly higher than the

rotation of 0.04 rad required for special moment frames (SMFs)

[7]. At that stage of these tests, since a high ductility factor had al-

ready been achieved, before doing further damage to the beam, it

was decided to unload and stop the test and to increase the preten-

sion forces of the connection bolts in order to limit the deforma-

tions due to slip. The pretension forces were increased from 42%

of the tensile strength of bolts to 68% and 56% (0.95 and 0.80 of

70% tensile strength) for BT connections of Specimens B2 and

B3, respectively, and to 60% (0.85 of 70% tensile strength) for TC

connections of both specimens. In the retest of these specimens,

only the inelastic cycles from 0.03 rad onward were applied

(Fig. 12B2 (retest) and B3 (retest)). The behaviour patterns for

the retests are similar to those of the specimens dominated by

rotation in the beams (Section 3.1). The cyclic behaviour indicating

connection slip and progressive failure (Videos 58 and Fig. 13) for

each of these tests and retests is discussed below.

3.2.1. Hysteretic momentrotation behaviour of Specimen B2The first slip at the through plate-to-column connection of

Specimen B2 occurred at the reverse cycle of h = 0.02 rad

(see Video 5) when the load dropped from M/Mp =0.8 (at

h = 0.015 rad, Point B) to M/Mp =0.56 (Fig. 12). During the sec-

ond cycle of 0.02 rad rotation, the connection slip load was reduced

to M/Mp = +0.48 at the positive direction. The slip load levels for all

subsequent rotations remained around M/Mp = 0.4.

The edges of the flanges between the vertical stiffener at the

through plate end and the lateral brace were slightly deformed

at h = 0.07 rad (see Video 5). This may be due to the initiation of

lateraltorsional buckling of the beam which, however, was pre-

vented by the lateral braces.


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At h = 0.08 rad, the maximum loads were M/Mp = 1.07 (Points C

and C, Fig. 12). At the end of this cycle web buckling was initiated

beyond the vertical stiffeners inside the beam.

At h = 0.09 rad the maximum moment was M/Mp = +1.0 (Point

D, Fig. 12). The flange buckling that was initiated in the previous

cycles was more noticeable (Fig. 13) and accompanied by large

connection slip and beam deformation.The total rotation of the BT and the TC connections measured

by inclinometer 6 is shown in Fig. 14a. Both the TC connection

(Fig. 14b) and the BT connection (Fig. 14c and d) contributed to

the total rotation of the connection. The start of the slip is identi-

fied by Point B on the connection rotation curves (Fig. 14a and

b). This point indicates that the slip of the connection started at

the TC connection (see Video 5). BT slip was activated from

the cycle at h = 0.05 rad (see Video 5 and Fig. 14c and d).

Most slip occurred at the stabilised slip load, Fs, at around

20 kN (corresponding to M/Mp = 0.40). The maximum rotation va-

lue of the connection (h6 = 0.055 rad) was higher than 0.0167 rad

calculated by assuming 2 mm bolt holes clearance, suggesting that

elongation of the bolt holes.

It can be concluded that the local rotations of the connectionelements (Fig. 14) made a large contribution (maximum

h = 0.055 rad) to the global rotation of 0.09 rad at the beam-end

(Fig. 12B2).

3.2.2. Hysteretic momentrotation behaviour of Specimen B2 during

retest

The torque on the bolts of the beam-to-through plate (BT) andthe through plate-to-column (TC) connections was increased

from 240 N m to 380 N m and from 170 N m to 240 N m, respec-

tively. The new increased pretension forces were calculated such

that the resistance of the farthest bolts was greater than the largest

applied forces (Rreq/Rn = 0.71 and 0.8 for the BT and the TC con-

nections, respectively). It should be noted that both curved flanges

of the beam were already slightly buckled from the previous test.

The specimen failed at Points E and E at a rotation of 0.08 rad

(see Fig. 12 and Video 6). Rupture in the bottom flange took place

at the bottom of the first bolt line of the BT connection following

large web and flange buckling in the beam (shown in Fig. 13).

The total rotation of the BT and the TC connections measured

by inclinometer 6 (Fig. 15a) and the rotation of the TC connection

(Fig. 15b) confirm that the connection slip was minimised by theincrease in the pretension force of the bolts. Therefore, the slip of

Fig. 11. Specimen B1: local deformations in elements of connection.

Fig. 12. Momentrotation hysteretic curves of Specimens B2 and B3.

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the connection elements (Fig. 15) made only a small contribution

to the global rotation at the beam end (Fig. 12).

3.2.3. Hysteretic momentrotation behaviour of Specimen B3

In the test of Specimen B3, the first slip was triggered ath = 0.03 rad, in the BT connection (see Video 7) at the load of M/

Mp = 0.76 (Fig. 12). At h = 0.05 rad rotation, the slip load level re-

duced to M/Mp = 0.58. The reduction in slip loads indicates either

loss of friction or pretension force in the bolts (or both). The slip

load levels for all the subsequent rotations remained around M/

Mp = 0.4 which corresponds to the same load level as for Speci-men B2. The maximum loads increased to M/Mp = 1.19 at Points

Fig. 13. Connection slip and failure deformations of Specimens B2 and B3 at the last cycle.


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C and C corresponding to h = 0.1 rad (Fig. 12). This confirms that

bolt slip eventually leads to significant bearing action, enough to

develop the full plastic moment capacity of the beam section.

The slip and elongation of the material around the bolt holes

(see Fig. 16) show that the first line of bolts in the BT connection

experienced the largest deformation. The last line of bolts (near the

face of column) and the second line of bolts also experienced some

slip and bearing action. This indicates that the centre of BT rota-

tion was not at the centre of the bolt group contrary to what was

assumed in its design.

The total rotation of the BT and the TC connections measured

by inclinometer 6 is shown in Fig. 17a. The rotation of the TC con-

nection (Fig. 17b) shows that a part of the total rotation was pro-

vided by the TC connection. The other part of the connectionrotation was provided by the BT connection (Fig. 17c and d).

The initial connection rotation was mainly provided by the BT

connection (see Video 6). The first line of bolts of the BT connec-

tion slipped first (as expected) as the rotation of the last line of

bolts was not initially significant (Fig. 17d). The slipof the TC con-

nection was mainly activated from the 0.06 rad cycle onwards (see

Video 6 and Fig. 17b).

It can be concluded that the local rotation of the connection ele-

ments (Fig. 17) made a large contribution (maximum h = 0.053 rad)

to the global rotation of 0.1 rad at the beam end (Fig. 12B3).

3.2.4. Hysteretic momentrotation behaviour of Specimen B3 during

retest

For the retest of Specimen B3 the torque on the beam-to-through plate and the through plate-to-column connections bolts

was increased from 240 N m to 320 N m and from 170 N m to

240 N m, respectively. The new increased pretension forces were

calculated such that the resistance of the farthest bolts was

greater than the largest applied forces (Rreq/Rn = 0.83 and 0.8

for the BT and the TC connections, respectively). It should be

noted that both curved flanges of beam already experienced

massive plastic strains (in excess of 35,000 le) from the previous

test.

Slip of the connections was triggered at h = 0.03 rad cycle (see

Video 8) at the load of M/Mp =0.89 which was higher than the

M/Mp = 0.76 from the first test (Fig. 12). In the second cycle of this

rotation, the slip load dropped to M/Mp = +0.78 and0.72.

At h = 0.06 rad cycle, flange local buckling was initiated be-

tween the 2nd and 3rd vertical stiffeners (see Video 8). Ath = 0.07 rad cycle, the slip loads occurred at M/Mp = +0.78 and

0.7 and the maximum loads reached M/Mp = 1.09 (Points C

and C, Fig. 12). At h = 0.09 rad, at the second cycle the maximum

loads dropped to M/Mp = +0.83 and 0.91 (Points D and D).

At h = 0.1 rad cycle, the load dropped significantly to Point E due

to rupture in the flanges which extended vertically through the

bolt holes in the webs, as shown in Figs. 13 and 18.

The total rotation of the BT and the TC connections ( Fig. 19a)

shows that connection slip was reduced in comparison with the

first test (Fig. 17). The slip of the TC connection (Fig. 19b) was

activated from the first cycle (see Video 8).

By increasing the pretension forces, most of the deformation

was provided in flexure by the beamand this led to very large plas-

tic strains on top of the plastic strains already developed during thefirst test. The fracture was a result of low cycle fatigue.

Fig. 15. Specimen B2 (retest): local deformations in elements of connection.

Fig. 16. Specimen B3: elongation of the material around the bolt holes in the BT connection after the test.

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Fig. 18. Specimen B3 (retest): Rupture at h = 0.1 rad cycle (Point E).

Fig. 19. Specimen B3 (retest): local deformations in elements of connection.



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3.3. Connection rigidity

The initial stiffness value ofSj,ini = KbLb/EIb proposed in Eurocode

3 part 18 [14] was used to classify the rigidity of the beam-to-

column connections: Sj,ini < 0.5Lb/EIb for simple connections and

Sj,ini > 25Lb/EIb for rigid connections, where Lb and EIb are the length

and bending rigidity of the beam, respectively.

For all the specimens, the connections can be placed in the cat-

egory of rigid connections. The initial slope of the curves, Mc/

(h6 h5), is larger than 25Lb/EIb as shown in Fig. 20 for Specimen

A1, where Mc is the bending moment at the column face and the

relative beam-column rotation (h6 h5) was taken as the differ-

ence of the values measured by inclinometers 5 and 6.

3.4. Strain distributions

This section presents the maximum strain values at the critical

sections of the through plate and the beam flanges normalised by

the proof strain, ey = 0.2%.

3.4.1. Strain distribution in through platesThe normalised strains for different cycles at the face of the col-

umn and at the middle of the inclined edges of the through plate

are shown in Fig. 21a and b, respectively for Specimen A1 typical

of specimens dominated by flexural and local buckling deforma-

tion in the beams (Specimens A13 and B1). At rotations exceeding

0.03 rad, following buckling initiation, the strain values begin to

drop. It is clear from Fig. 21 that, as desired, the through plate

didnot yield. The unsymmetrical results can be due to out-of-plane

bending of the through plate, which means there is a potential for

lateral failure mode if the beam is not braced properly.

For the specimens in which the deformations were dominated

by connection slip (Specimens B2 and B3), the through plate strain

values remained steady without noticeable change at large rota-

tions (Fig. 22).

3.4.2. Strain distribution in beam flanges

The typical normalised strains at the crest of the flanges (SG61-

71) for different cycles and for both loading directions are shown in

Fig. 23a and b for Specimen A1. The strain values show that e/ey = 1

occurred in the critical sections (SG63-69) immediately after0.015 rad corresponding to Point B in Fig. 6. At h = 0.03 rad-

0.04 rad rotations, the strain values in the critical sections in-

creased significantly due to buckling in compression (region CD

in Fig. 6). Due to extensive flange buckling after 0.04 rad, the strain

values in the critical sections increased massively and plastic

strains were locked in such a way that the local strains remained

compressive even when the flange was in tension (region DE in

Fig. 6).

For the specimens in which the deformations were dominated

by connection slip (Specimens B2 and B3) the strain increased in

a symmetric and gradual manner both on the tension and com-

pression flanges up to the last cycle of 0.1 rad rotation (Fig. 24

for Specimen B3) without any evidence of significant local buckling

(contrary to what was seen in the previous type of specimens).

4. General discussion

Based on the test results, this section presents a summary and

discussion on ductility, strength, energy, slip-bearing action and

design of the web bolted connections.

4.1. Ductility and strength

The use of the two stiffener configurations in the connections

(minimum: 2 pairs of vertical stiffeners, full: four pairs of vertical

and two pairs of horizontal stiffeners) resulted in a significant in-

crease in ductility for Specimens A2 and A3 by 50% and 75% and

for Specimens B2 and B3 by 28% and 43%, respectively, relativeto the specimens without stiffeners (A1 and B1; see Fig. 25a). Cor-

respondingly, the moment strength increased for Specimens A2

and A3 by 29% and 35% and for Specimens B2 and B3 by 10% and

23% (see Fig. 25b), respectively. The stiffeners constrained both

the flanges and the webs and increased the buckling resistance of

the elements.

4.2. Energy dissipation

The cumulative energy dissipation (E) derived from the hyster-

etic curves (at each rotation level) for all the specimens is shown in

Fig. 26.

Fig. 20. Specimen A1: connection rigidity Sj,ini = Mc/(h6 h5).

(a) (b)

Fig. 21. Specimen A1: maximum through plate normalised strains (ey = 0.2%) (a) at the face of the column (b) at the middle of the inclined edges.



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(a) (b)

Fig. 22. Specimen B2: maximum through plate normalised strains (ey = 0.2%) (a) at the face of the column (b) at the middle of the inclined edges.

Fig. 23. Specimen A1: maximum normalised strains (ey = 0.2%) at the crest of the flanges.

Fig. 24. Specimen B3: maximum normalised strains (ey = 0.2%) at the crest of the flanges.



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Specimen A1 (3 mm beam thickness) without any stiffeners dis-

sipated the lowest amount of energy (Fig. 26a). The use of partial

stiffeners for Specimen A2 increased the hysteretic energy by

$30%. The use of full stiffeners for Specimen A3 increased the hys-

teretic energy further by $90% and allowed the specimen to reach

larger rotation cycles (Fig. 26a). By increasing the beam thickness

to 4 mm in Specimen B1, the hysteretic energy increased by

$122% compared to Specimen A1 (Fig. 26a).

In specimens B2 and B3, the use of partial or full stiffeners in

conjunction with connection slip enabled the development of both

the nominal plastic moment capacity and stable hysteretic cycles

at large rotations. This led to a significant increase of the hysteretic

energy (e.g. 73% for B3 in Fig. 26b relative to B1). By limiting the

connection slip in the retest of specimens B2 and B3, the cumula-

tive hysteretic energy increased even further (240% for B3-retest in

Fig. 2.26b relative to B1) until the eventual failure due to flange

rupture. Connection slip can be beneficial in severe earthquakes

by delaying failures in beams and thus producing highly stable en-

ergy dissipating elements.

4.3. Slip-bearing action

Slip-bearing at the connections can be a highly nonlinear action

especially under cyclic loads. The observed slip of the bolts is char-

acterised by sudden changes (jumpinggripping), thus hardening

in the Mh curve could be either due to friction (gripping) or bear-

ing of the bolts. This depends on the frictional resistance between

bolts and plates and the contact resistance between bolts and the

material around the bolt holes. In the tested specimens the plate

surfaces were shot blasted to increase the roughness and slip

resistance. The applied pretension forces were expected to some-

what flatten the surface and reduce the roughness of the area of

the plates in contact with the bolt washers. Once the slip resistance

of the bolts was reached, the slipping bolts (and washers) had toflatten the rough surface of the adjacent area and as a result

experience higher frictional resistance which lead to higher loads.

On the other hand the position of the bolts relative to the plate

holes is not known. Therefore, the bolt shafts may come in contact

with the plate at any stage after slip. This contact can be in any

direction and it can be a full contact or just partial contact. Hence,

at the beginning of bolt slip the bearing contribution is not easily

determined. However, eventually at large rotations bearing action

will be fully mobilised.

The connection slip-bearing actions at high levels of load and

deformation (without any local buckling) were reached in Speci-

mens B2 and B3 (as expected). When bearing action was activated

after slip of the bolts, the bending moment in the beam increased

to values greater than the nominal Mp. Therefore, after reaching

large deformations in the connections, the deformation demand

was again shifted to the beam. This allowed the beams to sustain

their bending moment resistance at very large deformations with-

out any local buckling. The contribution of the connections in dis-

sipating seismic energy by slip-bearing actions can be utilised to

reduce the damage on the main structural members during severe

earthquakes.

4.3.1. Stabilised slip resistance load

After the initial slip cycles, the contact area within the slip zone

becomes smoother, thus leading to a reductionof the slip resistance

load at subsequent cycles, as shown in Fig. 27 for Specimens B2 and

B3. Another reason for this reduction may be the loss of pretension

forces in the bolts. In the case of specimens B2 and B3 the slip load

stabilised at around M/Mp = 0.4. These results can be used in up-

dated FE models to simulate the slip resistance [15]. Simple exper-

iments can also be devised to determine the cyclic slip force.

4.3.2. Design implications of slip-bearing actions

In Specimens A13 the required resistance of the farthest bolts

was lower than the design resistance [13] (Rreq/Rn = 0.83 for BTconnection and Rreq/Rn = 0.85 for TC connection). In Specimens

Fig. 25. (a) Ductility factors and (b) normalised bending moment strength of all specimens.

(a) (b)

Fig. 26. Hysteretic energy dissipation curves of all specimens.



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B13 the required resistance of the farthest bolts was higher than

the design resistance [13] (Rreq/Rn = 1.11 for BT connection and

Rreq/Rn = 1.13 for TC connection). Therefore, no connection slip

was expected in Specimens A1- 3. However, the test results

showed that slip was not completely absent in Specimens type A,

with noticeable slip in Specimen A2. The main reason for this isthe assumption of uniform rotation in connection design, which

was found not to be valid in these experiments as the first line of

bolts was more susceptible to slip than the other bolt lines. A thor-

ough investigation is needed to determine the precise location of

the centre of rotation.

Specimen B1 showed no noticeable connection slip contrary to

what was expected from the design calculations. The reason was

that only 0.97Mp was mobilised at the critical section of the beam

instead of the full design moment of 1.21Mp. Therefore, the re-

quired/available slip resistance ratio was reduced to 0.97/

1.21 1.13 = 0.91. In Specimens B2 and B3 the mobilised moments

were 1.065Mp and 1.19Mp, respectively, so the required/available

slip resistance ratios were reduced to 1.065/1.21 1.13 = 1.0 and

1.19/1.21 1.13 = 1.11, respectively.For Specimens B2 and B3 a large part of the deformation de-

mand was met by slip and bearing actions in the connection prior

to the occurrence of any local buckling in the beams. To achieve

this behaviour, it is recommended to design the slip of the connec-

tion at the level of nominal bending strength of the beam without

accounting for the strain hardening and material overstrength and

by applying a modification factor if the full moment strength can-

not be activated.

5. Conclusions

This study has shown that CFS curved flange beams cannot only

exceed the nominal plastic moment capacity (Mp), but also sustain

this capacity at large rotations similar to Class 1 cross sections inEurocode 3 part 1-1 [6] and larger than 0.04 rad required for spe-

cial moment frames in AISC Seismic Provisions [7].

The tested connections can be classified as rigid, and either par-

tial strength with M/Mp < 1 (Specimens A1 and B1) or full strength

with M/Mp > 1 (Specimens A23 and B23) according to Eurocode

3 part 18 [14].

In the beams of Specimens B2 and B3, large plastic strains were

delayed and occurred at very large rotations after mobilisation of

the bearing action in the connections. This enabled the CFS bolted

connections to produce highly stable hysteretic behaviour which

can be utilised to improve the seismic performance of buildings.

The use of a minimum configuration (for Specimens A2 and

B2) and a full configuration (for Specimens A3 and B3) of connec-

tion stiffeners resulted in a significant increase (relative to Spec-imens A1 and B1 without stiffeners) in (i) ductility (up to 75%),

(ii) moment capacity (up to 35%) and (iii) hysteretic energy dissi-

pation capacity (up to 240%) also due to the activation of connec-

tion slip.

Acknowledgements

The authors would like to express their gratitude to the Earth-quake Research Group in the Department of Civil and Structural

Engineering at The University of Sheffield and Corus Research,

Development & Technology for their financial support.

Appendix A

The design specifications of the auxiliary (test set-up) compo-

nents are the following:

Connections: The force distribution in the beam-to-through

plate and through plate-to-column connections was based on

the assumption of uniform rotation in the connections. The col-

umn-to-reaction frame connection was designed by assumingcoupling forces such that to provide overstrength in this

connection.

Reaction frame: This comprised two very large (and stiff) col-

umns seated on a large bottom beam and a supporting top

beam. The test column was connected to the lower part of

one of the reaction frame columns. Lateral restraints were con-

nected to the bottom beam and the loading actuator was

hanged from the top beam of the reaction frame.

Column: Back-to-back hot-rolled channels were used with

flange width of 150 mm, thickness of 16 mm, and web depth

of 300 mm and thickness of 10 mm.

Lateral restraints: Two lateral brace frames were used to avoid

premature global instability. According to the AISC Seismic Pro-

visions [7], both beam flanges should be laterally braced nearplastic hinges and regions with concentrated forces. Four brack-

ets and PTFE sheets were used between the bracing frames and

the beam at both sides to minimise friction.

Loading plates: A loading plate connected by 6M16 was placed

between the webs of the beam sections. This plate was con-

nected to the head of the actuator by the detail shown in

Fig. 3. The reason of using this complicated detail was the

height limitation of the supporting frame that lowered the actu-

ator head below the beam.

Stiffeners: The end of the beams at the loading point was stiff-

ened to avoid premature local failure. For this purpose, three

pairs of stiffeners were welded inside the beams near the load

application point (Fig. 3). Four pairs of stiffeners and continuity

plates were welded inside the webs of the column and the reac-tion frame (Fig. 3), to avoid any premature local failure in the

connection between the column and reaction frame.

Filler plates: To protect the strain gauges in the connection panel

zone area and to simulate the actual boundary conditions of the

through plate in tube columns (without restraint inside the col-

umn), two filler plates were placed on each side of the through

plate at the locations of the through plate-to-column connec-

tion bolts (Fig. 3).

Appendix B. Supplementary material

Supplementary data associated with this article can be found, in

the online version, at http://dx.doi.org/10.1016/j.engstruct.2012.04.025.

Fig. 27. Slip resistance at the loading cycles of Specimens B2 and B3.

http://dx.doi.org/10.1016/j.engstruct.2012.04.025http://dx.doi.org/10.1016/j.engstruct.2012.04.025http://dx.doi.org/10.1016/j.engstruct.2012.04.025http://dx.doi.org/10.1016/j.engstruct.2012.04.025


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Documents

Experimental Work on Cold-Formed Steel Elements for Earthquake Resilient Moment Frame Buildings