ORIGINAL ARTICLE

Fire response of steel column-tree moment resisting frames

Abbas Rezaeian • Mahmood Yahyai

Received: 1 January 2014 / Accepted: 7 February 2014

� RILEM 2014

Abstract The column-tree moment resisting frames

are common steel construction design in many coun-

tries. Very limited research has been carried out on

such systems at elevated temperatures. This paper

presents experimental investigations of the perfor-

mance of beam and its bolted connections in steel

column-tree MRF under fire conditions. Six full-scale

steel sub-frames with different link-to-stub beam

connections were tested under ISO 834 standard fire.

The effects of factors, including load level, splice

plate’s size, bolt size and bolt grade were studied. The

thermal and structural fire behaviors as well as failure

modes were investigated. The link-to-stub beam con-

nections failed at temperatures beyond 750 �C, while

beam underwent large deflections of more than span/

20. It was observed that, bolt shear fracture generally

governs the failure of flange connection, whereas bolt

hole bearing controls the failure of web connection at

elevated temperatures. The results show that the use of

stronger bolts in flange splice can significantly enhance

the resistance and rotational capacity of the link-to-

stub beam connection at high temperatures.

Keywords Column-tree � Link-to-stub

connection � Fire test �Moment resisting frame �Steel beam � Elevated temperature

1 Introduction

In recent years, major research endeavors have been

devoted to better understanding of structural behavior

under fire conditions, development of rational design

approaches for evaluating fire resistance of structures,

accompanied by the development of sophisticated

codes of practice such as the Eurocode [4] and AISC’s

steel design manual [1]. Observations from full-scale

fire tests confirm that connections play an important

role on the resistance time of structural components in

fire. Because of the high cost of elevated temperature

tests, adequate experimental data on a broad range of

connections are not available. One type of such

connections is the link-to-stub beam connections

(splice connections) in column-tree moment resisting

frame (MRF).

The column-tree MRF is a common steel construc-

tion design in seismic risk regions due to its better

quality control in stub beam-to-column shop welding

and erection efficiency of link-to-stub field bolting

during construction [2]. As has been demonstrated by

the collapse of World Trade Center 5 (WTC 5) building,

the inadequate design of link-to-stub connections in

column-trees could cause the premature failure in fire

conditions which might result in successive collapse of

the structure (Fig. 1) [5]. Also, the National Institute of

Standards and Technology, which has published the

results of its investigation of the WTC disaster, identifies

structural connections under fire exposures as a vital

area for further study [14].

A. Rezaeian (&) � M. Yahyai

Civil Engineering Department, K.N. Toosi University of

Technology, Tehran, Iran

e-mail: rezaeian_a@scu.ac.ir; yahyai@kntu.ac.ir

Materials and Structures

DOI 10.1617/s11527-014-0271-1

Considering the importance of connection perfor-

mance in steel beams under fire condition, some

experimental and numerical studies were carried out

to investigate the behavior of connections in steel

buildings. Most of the fire tests to date have been

performed on single isolated joints [15–17, 20, 22]. On

the other hand, observations from real fires show that,

on some occasions, the accumulative effects of a

number of factors (including hogging bending moment,

tension field action in shear and high cooling strain or

pulling in effect at large deflections of the connected

beam) could make the tension components of the joints

fracture [6, 13]. Therefore, limited number of research-

ers utilized sub-frame assembly to test the beam and its

connections in fire conditions [11, 12, 18, 21]. This

assembly can simulate the realistic forces applied to

beam’s connections in fire tests. Very limited numerical

researches have been carried out on column-tree MRF

under fire conditions [10]. However, there has been no

report of full-scale experimental study on column-tree

MRF under fire loading.

The aim of this work is to investigate the behavior of

beam and splice connections in column-tree MRF at

elevated temperatures through experimental study. In

this research six full-scale steel sub-frames with different

beam splice components are tested under standard fire.

2 Test setup and specimens

2.1 Test setup

A rectangular fire testing furnace having internal

dimensions of 5,970 9 2,930 9 2,960 mm was used.

Both sides of the furnace were equipped with eight

gas-fire burner nozzles (Fig. 2). To achieve the

maximum heat efficiency of these burners, the internal

faces of the furnace were covered with 75 mm thick

ceramic fiber wool. Six K-type thermocouples were

Fig. 1 Internal collapse area in WTC 5 building [5]

Fig. 2 Plan view of

schematic arrangement of

the furnace

Materials and Structures

used to record the furnace temperatures. Their average

time–temperature relationship was intended to follow

the standard fire condition, mentioned in ISO 834 [7].

In total, six fire tests were conducted to investigate

the behavior of beam and splice connections under fire

loading in column-tree MRF. The effect of load level,

splice plates size, bolt size and bolt grade on speci-

mens performance were studied. The arrangement of

the tests was in the form of a sub-frame as shown in

Fig. 3. The columns and bottom girder, both having

Fig. 3 Elevation view of test setup

Table 1 Level of loading for specimens

Specimen number Load ratio Applied

load P (kN)

S-1 0.7 20.58

S-2 0.5 14.72

S-3 0.7 20.57

S-4 0.7 20.61

S-5 0.7 20.60

S-6 0.7 20.62

Fig. 4 Location of

measurement devices on a

specimen

Materials and Structures

dimensions of 440 9 250 9 12 9 20 mm section,

together with the lateral restraint system were fire-

protected by 50 mm thick ceramic fiber blanket, so

that could be reused in all the tests. The lateral

movement of the beam was restrained at three points;

mid-span point, and at 1/5 of the span on either side.

Two concentrated mechanical loads were symmet-

rically applied on 500 mm away from either side of the

beam mid-span (Fig. 3). The load ratio of about 0.7

was taken here for five tests as given in Table 1, since it

is common for most steel beams, designed on the basis

of full strength at ambient temperature. Also to observe

the load ratio effect, one specimen was tested under

Fig. 5 Detail of link-to-stub beam splices (dimension in mm)

Table 2 Summary of specimen details

Specimen number Beam section Splice plates size (mm) Bolt size Bolt grade Distances (mm)

Flange Web Flange Web d1 d2 d3 e1 e2

S-1 IPE 200 240 9 90 9 10 2(140 9 100 9 4) M12 M12 8.8 50 36 50 23 20

S-2 IPE 200 240 9 90 9 10 2(140 9 100 9 4) M12 M12 8.8 50 36 50 23 20

S-3 IPE 200 240 9 100 9 10 2(140 9 100 9 6) M12 M12 8.8 50 36 50 23 25

S-4 IPE 200 260 9 90 9 10 2(140 9 100 9 4) M14 M12 8.8 50 42 50 21 20

S-5 IPE 200 290 9 90 9 10 2(140 9 100 9 4) M16 M12 8.8 54 48 46 22 22

S-6 IPE 200 240 9 90 9 10 2(140 9 100 9 4) M12 M12 10.9 50 36 50 23 20

Table 3 Material properties of the specimens

Material Grade Yield

stress Fy

(MPa)

Ultimate

stress Fu

(MPa)

Modulus of

elasticity

E (MPa)

Beam S235 242 420 2.06 9 105

Plate S235 296 442 2.06 9 105

Bolt 8.8 737 963 2.00 9 105

Bolt 10.9 1,052 1,136 2.00 9 105 0100200300400500600700800900

1000

0 20 40 60 80 100 120 140

Tem

per

atu

re (

°C)

Time (minute)

ISO 834S-1S-3S-2S-4S-5S-6

Fig. 6 ISO 834 standard fire and the average temperatures of

furnace

Materials and Structures

load ratio of about 0.5. The load ratio is defined as the

ratio of the applied load during the fire test to the load-

carrying capacity of the beam at room temperature.

The testing procedure comprised two sequential

steps: step (1) the load was applied to reach a

predetermined level; step (2) the heating was applied

on. The mechanical load was kept constant and the

thermal load was increased according to the standard

fire condition, mentioned in ISO 834, until the

connection failure occurred. The specimens were

continuously monitored using a digital camera placed

in front of the observation hole.

0

100

200

300

400

500

600

700

800

900

Tem

per

atu

re (

°C)

Time (minute)

(S-1)

0

100

200

300

400

500

600

700

800

900

Tem

per

atu

re (

°C)

Time (minute)

(S-5)

0

100

200

300

400

500

600

700

800

900

Tem

per

atu

re (

°C)

Time (minute)

(S-6)

0

100

200

300

400

500

600

700

800

900

1000

Tem

per

atu

re (

°C)

Time (minute)

(S-2)

furnacetop flangewebbot flangetop platetop boltwebboltwebplatebot platebot bolt

0

100

200

300

400

500

600

700

800

900

Tem

per

atu

re (

°C)

Time (minute)

(S-4)

0

100

200

300

400

500

600

700

800

900

0 10 20 30 40

0 10 20 30 40 0 10 20 30 40

0 10 20 30 40

0 10 20 30 400 10 20 30 40

Tem

per

atu

re (

°C)

Time (minute)

(S-3)

Fig. 7 Temperature

distribution in specimens

and furnace temperature

Materials and Structures

2.2 Instrumentation

In order to monitor the temperature distribution in the

structure, several K-type thermocouples were installed

on the beam, splice plates, bolts and supporting frame,

as shown in Fig. 4. The displacement transducers

(LVDT) were designed on the beam mid-span and

either side of the splice connection zones. To

minimize the unwanted effects of elevated tempera-

ture, all displacement transducer were placed outside

the furnace, and displacements were measured via

coated ceramic rods inserted through the fiber lining of

the furnace. These measurements as well as the

thermocouple temperatures were recorded through a

computerized data recording system.

2.3 Test specimens

In all specimens, two stub beams with the length of

500 mm were welded to the columns and then a link

beam having 2,980 mm length was fully bolted to the

ends of stub beams using web and flange splice plates

(Fig. 3). The flange splice plates were configured as a

single plate with single shear bolts as shown in Fig. 5.

Details of various components of connection are given

in Table 2. The cross-section of beams was European

profile IPE 200. All the bolts and nuts were Grade 8.8

except for specimen S-6 to investigate the effect of

bolt grade. Standard holes created for the bolts, were

2 mm greater than the nominal bolt diameter accord-

ing to the AISC [1].

2.4 Material properties

For all specimens, the mechanical properties of steel

members at ambient temperature were measured using

standard tensile coupon tests, and cross sectional

dimensions were recorded prior to testing in the

furnace. The material properties are reported in

Table 3.

3 Test results

3.1 Temperature distribution

The average measured temperatures in the furnace

during the tests are compared with ISO 834 standard

fire temperature curve in Fig. 6. It can be clearly seen

0

50

100

150

200

250

300

350

0 100 200 300 400 500 600 700 800 900 1000

Bea

m m

id-s

pan

def

lect

ion

Beam bottom flange temperature (°C)

S-1

S-3

S-2

S-4

S-5

S-6

(mm

)

Fig. 8 Beam mid-span temperature–deflection

Table 4 Failure criteria for flexural members as per BS-476

Member

section

Member

dimensions

BS-476 failure criteria [meet

either (1) or (2)]

(1) (2)

L (mm) d (mm) L/20

(mm)

L/30

(mm)

L2/(9,000d)

(mm/min)

IPE 200 4,000 200 200 133.3 8.9

Table 5 Temperature, deflection and rotation at failure

Specimen

number

Failure of top flange splice

Beam

bottom

flange

temp. (�C)

Top

bolts

temp.

(�C)

Beam mid-

span

deflection

(mm)

Splice

rotation

(millirads)

S-1 755 763 212 211

S-2 919 914 259 253

S-3 787 779 224 221

S-4 806 765 270 269

S-5 840 a 323 311

S-6 843 832 293 285

a Thermocouple failure

0100200300400500600700800900

1000

0 50 100 150 200 250 300 350

Rotation (milirad)

S-1

S-2

S-3

S-4

S-5

S-6Bea

m b

ott

om

fla

ng

e te

mp

erat

ure

(°C

)

Fig. 9 Splice connection temperature–rotation

Materials and Structures

that the furnace temperature closely follows the

standard curve. The temperature histories of the

specimens were obtained during the tests. Measure-

ments of the temperature in the beam mid-span cross

section were taken on the web and flanges, as shown in

Fig. 4. Also, the temperatures of the plates and bolts

were measured in the beam splices. Since the results

on both sides of the specimen follow the same pattern,

only temperatures on the left-hand side splice con-

nection are presented in Fig. 7.

To be precise, during the heating, all the compo-

nents follow a similar trend in their temperature

profiles. Top and bottom flanges, as well as splice

components show negligible temperature differences.

Average temperatures of the protected columns did

not rise beyond 167 �C. Hence, material properties of

the columns and performance of the specimens were

not affected.

3.2 Beam deflection and splice connection

rotation

The temperature–deflection curves of the beam at mid-

span are represented in Fig. 8. Measurements were

recorded until the beam splice failed. The beam

deflections occurred in three phases. During the early

stage of fire exposure, the beam started to bend and the

Table 6 The relations between tests

Specimen

number

Specimen

number

Relation between tests

S-1 S-2 Effect of load ratio

S-1 S-3 Effect of splice plates size

S-1 S-4, S-5 Effect of bolt size

S-4 S-5 Effect of bolt size

S-1 S-6 Effect of bolt grade

0

50

100

150

200

250

300

350

Bea

m m

id-s

pan

def

lect

ion

(m

m)

Beam bottom flange (a)

S-1

S-2

0

50

100

150

200

250

300

350Bea

m m

id-s

pan

def

lect

ion

(m

m)

Beam bottom flange (b)

S-1

S-3

0

50

100

150

200

250

300

350Bea

m m

id-s

pan

def

lect

ion

(m

m)

Beam bottom flange

(d)

S-1

S-6

0

50

100

150

200

250

300

350

0 200 400 600 800 1000 0 200 400 600 800 1000

0 200 400 600 800 10000 200 400 600 800 1000

Bea

m m

id-s

pan

def

lect

ion

(m

m)

Beam bottom flange

(c)

S-1

S-4

S-5

temperature (°C) temperature (°C)

temperature (°C) temperature (°C)Fig. 10 Comparisons of

temperature–deflection

curves: a load ratio, b splice

plates size, c bolt size, d bolt

grade

Materials and Structures

deflection increased slowly. When the temperature of

beam bottom flange reached about 600 �C, drastic

increase was observed in the beam deflection due to

sudden reductions in strength and stiffness of steel,

leading to a progressive runaway of the beam deflection.

Finally, the deflection rate decreased at about 750 �C,

where large deflection was observed in the beam.

Excessive deflection and deflection rate were

defined in accordance with BS-476 [3]. According to

this standard, flexural failure occurs when either:

(1) The deflection exceeds unsupported length

(L) divided by 20, or

(2) The deflection exceeds unsupported length

(L) divided by 30 and the deflection rate exceeds

L2 divided by 9,000 times the depth (d).

Table 4 summarizes these failure criteria for the

IPE200 beam in the specimens. For instance, the beam

in S-1 failed after 29 min of heating with a mid-span

deflection of 151 mm and deflection rate of 11.1 mm/

min, which were greater than those mentioned in

second failure criterion of BS-476. The beam deflection

continued to increase and reached a maximum value of

212 mm after 32.5 min of heating, where the splice

connection failed through fracture of bolts. Summary of

temperatures, deflections and rotations at connection

failure is presented in Table 5. As can be seen in all fire

tests, link-to-stub splice connection failed after the first

flexural failure criteria of BS-476 was satisfied.

The experimental temperature–rotation curves of

left-hand beam splice at elevated temperatures are

plotted in Fig. 9. These were derived from the relative

displacement of either side of the splices. Results

showed that the rotations of beam splices in each

specimen are almost symmetrical. The behavior was

almost linear at the beginning. Since the expansion of

the beam was restrained by the supports, the horizontal

gaps between the bottom flanges in beam splices were

closed. Thereafter, the splice connection entered a

non-linear stage upon progressive runaway of the link

beam deflection, simultaneously with the onset of

plastic shearing of flange bolts and bearing

0

100

200

300

400

500

600

700

800

900

1000

Bea

m b

ott

om

fla

ng

e te

mp

erat

ure

(°C

)

Rotation (milirad)

(a)

S-1

S-2

0

100

200

300

400

500

600

700

800

900

1000

Bea

m b

ott

om

fla

ng

e te

mp

erat

ure

(°C

)

Rotation (milirad)

(b)

S-1

S-3

0

100

200

300

400

500

600

700

800

900

1000

Bea

m b

ott

om

fla

ng

e te

mp

erat

ure

(°C

)

Rotation (milirad)

(d)

S-1

S-6

0

100

200

300

400

500

600

700

800

900

1000

0 100 200 300 0 100 200 300

0 100 200 3000 100 200 300

Bea

m b

ott

om

fla

ng

e te

mp

erat

ure

(°C

)

Rotation (milirad)

(c)

S-1

S-4

S-5

Fig. 11 Comparisons of

temperature–rotation

curves: a load ratio, b splice

plates size, c bolt size, d bolt

grade

Materials and Structures

deformation of the bolt holes. The bending moment

along with the significant reduction in steel strength

caused failure of the beam splices.

4 Discussion of results

The beam’s ability to survive high temperatures depends

on the ability of connections to resist the tensile force

due to catenary action and hogging bending moment.

The effects of the four parameters, depicted in Table 6,

on structural behavior including beam deflections,

splice rotations and failure modes are discussed.

4.1 Effect of load ratio

Effect of load ratio on the behavior of beam and splice

connection was investigated. For this purpose, two

identical specimens S-1 and S-2 were tested under

different load ratios (Table 1). Temperature–deflec-

tion and temperature–rotation curves of these speci-

mens are compared in Figs. 10a and 11a, respectively.

The rate of beam deflection and splice rotation in S-2

was decreased in comparison to S-1. Also, specimen

S-2 experienced larger mid-span deflection and splice

rotation before failure. The beam bottom flange

temperature in specimen S-2 reached about 900 �C

before failure, which was the maximum tolerable

temperature in all tests as shown in Table 5. It was

164 �C higher than that of S-1. In other words, since

the applied load combination consisted of mechanical

and thermal loads, reduction of mechanical load

allowed the specimen to tolerate higher temperatures.

4.2 Effect of splice plates size

The only difference between specimens S-1 and S-3

was beam splice plates size. The thickness of web

plates and width of flange plates were increased in S-3

as per Table 2. The temperature–deflection and tem-

perature–rotation curves of these specimens are com-

pared in Figs. 10b and 11b respectively. The results

show that, splice plates size had negligible influence

on the behavior of the specimens. In fact, premature

failure of bolts did not allow the capacity of flange

plates to be fully used. In other words, since the failure

is controlled by fracture of the bolts, using greater

splice plates do not always increase the resistance of

splice connection in fire.Fig. 12 Observations of beam splices in specimen S-1 during

fire test (t = time in minute)

Materials and Structures

4.3 Effect of bolt size and bolt grade

Bolts are one of the most important elements that

affect the behavior of connections and connected

beam in fire condition. Earlier studies have shown that

the tensile and shear strength of bolts deteriorates

dramatically at temperatures between 300 and 700 �C

[8, 9]. Specimens S-1–S-3 using grade 8.8 M12 bolts

showed that the shear resistance of the top flange bolts

is lower than the bearing resistance of the splice plate

at elevated temperatures. Therefore, the remaining

three specimens were modified to check whether an

Fig. 13 Overall deformation shape of specimen S-6 after fire test

Fig. 14 Failed splice connection after fire test in specimens S-1, S-2, S-5 and S-6

Materials and Structures

increase of the bolt resistance would change the failure

mode and enhance the tying resistance of the connec-

tion in fire. In specimens S-4 and S-5 the bolt diameter

was increased as per Table 2, and for S-6 the bolt

grade was increased to 10.9.

The beam mid-span deflection and splice rotation of

these specimens are compared in Figs. 10c, d and 11c,

d. As can be seen, using thicker bolts or higher grade

bolts increased the fire resistance of splice connections

and connected beam significantly at elevated temper-

atures. The specimen S-6 tolerated 843 �C at bottom

flange before failure. It was the maximum tolerable

temperature in tests that conducted under same load

ratio. This temperature was 88 and 37 �C higher than

that in S-1 and S-4 respectively (Table 5). This was

because of the higher strength of 10.9 bolts at elevated

temperature. The failure was still controlled by bolt

shear, but significant bearing deformations to the bolt

holes were caused before top bolts shear. Moreover,

the failure of the specimens S-5 was due to tensile

fracture of the net area of top splice plate in contrast to

that of other specimens.

5 Failure modes

During the tests, the expansion of beam assembly led

to closure of the gaps between the link beam and stub

beams, and subsequently forced the bolts into the

flanges and webs. Furthermore, the rigidity of spec-

imen decreased with increasing the temperature and

the beam deflected significantly. This deflection

caused the end of link beam to rotate and bottom

flange of the link to push the stub beam. As the end of

link beam continued to rotate in response to mid-span

deflection, the top flange splice plate was extremely

pulled and finally the connection failure took place

(Fig. 12).

Table 7 Summary of splice connection failure modes

Specimen

number

Top flange

splice

Web splice Bottom flange

splice

S-1, S-3 Bolts shear Bolt holes

bearing

Bolts shear

S-2 Bolts shear Bolts shear Bolts shear

S-4 Bolts shear Bolt holes

bearing

No fracture

S-5 Tensile fracture

of net areas

of plate

Bolt holes

bearing

No fracture

S-6 Bolts shear Bolt holes

bearing

Bolts shear

Fig. 15 Details of failure mode of connection at specimen S-1:

a top flange splice, b web splice, c bottom flange splice

Materials and Structures

The beams experienced large deflections before

connection failure as shown in Fig. 13. The post-test

failed shapes of the link-to-stub splice connection in

specimens S-1, S-2, S-5 and S-6 are presented in

Fig. 14. The failures occurred sequentially in top flange

splice, web splice and bottom flange splice of the

connection as summarized in Table 7. It was observed

that the bearing deformation of the top bolts increased

with increasing the connection rotation, until complete

fracture. The failure of top flange splice was shear

fracture of bolts connecting splice plate to the link beam

in specimen S-1 (Fig. 15a). In S-6 significant bearing

deformations to the bolt holes were also observed

before top bolts shear as shown in Fig. 16. In the case of

S-5, on the other hand, the failure was due to tensile

fracture of the net area of top splice plate (Fig. 17a),

because much stronger bolts resisted the shear forces

and only experienced intensive bearing deformation.

As can be seen, bolt shear fracture generally

governs the failure of flange connection at elevated

temperatures. This is because design guides such as

Eurocode 3: part 1.8 predict the minimum resistance

based on bearing resistance. The reduction of bearing

resistance at elevated temperature follows that of the

structural steel strength. As shown in Fig. 18, the bolt

strength reduces more rapidly than normal structural

steel strength. Therefore, the bolts in shear become

more critical in comparison to plates in bearing at

elevated temperatures.

In all specimens except S-2, a block-shear failure in

stub web was then developed from the bolt holes

towards the end of stub beams. By this stage, web bolts

also experienced significant bearing deformation as

shown in Fig. 19a. In the case of S-2, the failure was

double shear fracture of web bolts and the stub beam

web underwent obvious bearing deformation in the

bolt holes (Fig. 19b). As can be seen, bolt hole bearing

generally controls the failure of web connection. While

modifications to increase the bearing strength can be

made to the web connection (such as increasing the

distance from bolt hole centerline to the end of the

beam stub [19]), the limit state shifts from bearing to

bolt shear at higher temperatures. In such a scenario,

there is no advantage to modifying the web splice

connection.

Fig. 16 Deformation of top flange splice in specimen S-6

Fig. 17 Details of failure mode of connection at specimen S-5:

a top flange splice, b web splice, c bottom flange splice

Materials and Structures

As shown in Fig. 15c, the final failure occurred due

to the fracture of the bolts in the bottom flange splice

of all specimens except S-4 and S-5 (Fig. 17c). The

bolts adjacent to the splice gap experienced significant

tensile and bearing deformations and showed ductile

necking in the thread before fracture.

6 Conclusions

Experimental results revealed various failure modes of

the splice connections in column-tree MRF. The shear

fracture of bolts or tensile fracture of the net area of

plate in top flange splice occurred at temperatures

beyond 750 �C. Consequently, stub beam web failed at

those temperatures because of block-shear. In the case

which fire resistance of splice connection was about

900 �C, double shear fracture of web bolts occurred.

The fire resistance of link-to-stub beam connections

increased significantly by decreasing the applied gravity

load, while the rotational capacity did not increase

considerably. Since the failure is controlled by fracture

of the bolts, using greater splice plates do not always

increase the resistance of connection in fire. The use of

stronger bolts can significantly enhance the resistance

and rotational capacity of the link-to-stub beam connec-

tion at high temperatures. It is suggested that the bolts

shall be designed stronger than plate in flange splice.

The temperature–deflection and temperature–rota-

tion curves remained in the elastic range until

550–650 �C. Between 650 and 750 �C, the behavior

would be highly nonlinear plastic. The beams expe-

rienced large deflections beyond failure criteria of BS-

476 (span/20) before connection failure. In other

words, considering the BS-476 standard for steel

beams in fire, the beam splices would fail after the

beam failure in column-tree moment-resisting frames

in the range of this study.

References

1. AISC (2010) Specification for structural steel buildings

360–10. American Institute of Steel Construction Inc., Chicago

2. Astaneh-Asl A (1997) Seismic design of steel column-tree

moment-resisting frames. Structural Steel Educational

Council, Berkeley

0

0.2

0.4

0.6

0.8

1

1.2

0 200 400 600 800 1000 1200

Red

uct

ion

fac

tor

Temperature (°C)

Yield strength [EC3]Ky,θ = f y,θ/fy

Kb,θ [Kirby]

Fig. 18 Reduction factors for stress–strain relationship of

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Kirby [8] at elevated temperatures

Fig. 19 Failure modes of web splice connection: a specimen

S-1, b specimen S-2

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