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SEISMIC-INDUCED FIRE ANALYSIS OF STEEL-CONCRETE COMPOSITE BEAM-TO-COLUMN JOINTS: BOLTED SOLUTIONS
Oreste S. Bursi
Department of Mechanical and Structural Engineering, University of Trento Trento, Italy
Fabio Ferrario Department of Mechanical and Structural Engineering, University of Trento
Trento, Italy [email protected]
Raffaele Pucinotti
Department of Mechanics and Materials, Mediterranean University of Reggio Calabria Reggio Calabria, Italy
Riccardo Zandonini Department of Mechanical and Structural Engineering, University of Trento
Trento, Italy [email protected]
ABSTRACT In this paper, a multi-objective advanced design methodology is proposed for steel-concrete composite moment-resisting frames. The research activity mainly focused on the design of beam-to-column joints under seismic-induced fire loading together with the definition of adequate structural details for composite columns. Thermal analyses of cross sections were performed in order to obtain internal temperature distribution; structural analyses were then carried out on the whole frame to assess the global behavior under the combined action of static and fire loadings. Furthermore, results of numerical analyses were used in order to derive information about the mechanical and numerical behavior of joints. In this paper the experimental program carried out on four beam-to-column joint specimens under seismic loading is described and results are presented and discussed together with the outcomes of numerical simulations owing to seismic and fire actions. Experimental tests demonstrated the adequacy of the seismic design. Numerical simulations showed a satisfactory performance of joints under seismic-induced fire loading.
INTRODUCTION Steel-concrete composite structures are becoming increasingly popular around the world due to the favorable performance regarding stiffness, strength and ductility of composite systems under seismic loading, and also due to the speed and ease of erection. Moreover, such
structures exhibit better fire resistance characteristics compared to bare steel structures. Considering the high probability of a fire after a seismic event, the use of steel–concrete composite structures in seismic areas potentially represents a fairly effective design solution. In fact, the adoption of such a solution is in general more efficient from both a structural and a constructional viewpoint when compared to bare steel structures. The benefits increase, if the probability that an earthquake and a fire can occur in sequence is high. In current Eurocodes for structural design, seismic and fire safety are accounted for separately, and no sequence of seismic and fire loading is taken into account. In reality, the risk of loss of life increases if a fire occurs in a building after an earthquake. In the Kobe Earthquake (1995) many people died due to the collapse of buildings exposed to a fire that followed an earthquake; in fact, large sections of the city burned, greatly contributing to the loss of lives. It is obvious therefore that seismic-induced fire is a design scenario that should be properly addressed in any performance-based seismic design. In this situation the traditional single-objective design is not adequate, and a multi-objective advanced design has to be adopted. This makes it possible to take into account of: (i) seismic safety with regard to accidental actions; (ii) fire safety with regard to accidental actions (iii) seismic-induced fire safety on a structure characterized by stiffness deterioration and strength degradation owing to seismic actions. As a result, the fire design applied to a structure with reduced capacity owing to seismic damage will permit a simultaneous fulfillment of the requirements associated with structural, seismic and fire safety, and with structural fire safety and structural seismic safety, separately, considered as accidental actions. In order to characterize the seismic behaviour of composite joints, a research project were carried out by the University of Trento with the objective of developing a design procedure for composite joints under the ‘combined’ action of earthquake and post-earthquake fire. The research activity was mainly concerned with the design of bolted beam-to-column joints with CFT columns with circular hollow steel sections. This solution derives from a parent welded design solution developed in a European project (Bursi et al. 2008) and is aimed at ensuring easiness of assembly and avoiding problems related to on-site welding. The results of the experimental programme devoted to the evaluation of the cyclic behavior of beam-to-column bolted joints are reported in the following. In order to better understand the activation of all the transfer mechanisms proposed in Eurocode 8 (UNI EN 1998-1, 2005), a numerical finite element (FE) model of the slab has been developed, together with FE model of the joint subject to fire.
SEISMIC AND FIRE DESIGN OF REFERENCE FRAMES A reference composite steel-concrete office building was considered, made up of three five storeys moment resisting frames, placed at a distance of 7.5 m each in the longitudinal direction and braced in the transverse direction. The storey height was equal to 3.5 m. Two moment resisting frames, having the same structural typology but different slab systems, were analyzed: i) a composite steel-concrete slab with structural profiled ribbed steel sheeting; ii) a concrete slab composed of electro-welded lattice girders. Since the two slab systems had different load bearing capacities, a different distance between the secondary beams was adopted for the two solutions. The distance slightly affected the frame geometry. The first solution (steel sheeting slab), is depicted in Figure 1a while the second one, with the slab on prefabricated lattice elements, is shown in Figure 1b. The elevation is shown in Figure 1c (Bursi et. al 2008). Two different types of composite beams were hence designed according to Eurocode 4 (UNI EN 1994-1-1, 2005); the steel section was maintained in both cases as an IPE400 of steel grade S355 (A615 grade 50). The slab, 150 mm (5.90 in.) deep, was designed in accordance with the relevant specifications of Section 9 of Eurocode 4 [1994-1-1. 2005]; moreover design procedure indicated by the producer of the prefabricated lattice elements was followed in the second case.
All connections between the steel beam and the slab were made by Nelson 19 mm stud connectors made of steel with an ultimate tensile strength fu=450 MPa.
a) b) c)
Figure 1 - The configuration of the building and the longitudinal five-storey MR Frame; slab with a) profiled ribbed steel sheeting; b) slab with prefabricated lattice girder; c) elevation
The frames were designed to have a dissipative structural behavior according to Eurocode 8 (UNI EN 1998-1, 2005). In detail two following structural types were considered.
1. Moment Resisting Frames: in which dissipative zones were mainly located to the end of the beams, in the vicinity of beam-to-column connections, where plastic hinges could develop (UNI EN 1998-1, 2005, 7.1.2 b);
2. Concentrically braced frames: in which the horizontal forces were mainly resisted by members subjected to axial forces with dissipative zones located on the tension diagonals only (UNI EN 1998-1, 2005, 7.1.2 c).
The effective width of the slab in composite beams was determined according to Eurocode 4 (UNI EN 1994-1-1, 2005, 5.4.1.2) for static and fire analyses, and to Eurocode 8 (UNI EN 1998-1, 2005, 7.6.3) for seismic analyses. Beam-to-column connections resulted to be rigid according to Eurocode criterion (UNI EN 1994-1-1, 5.1.2). The design of structural elements were performed considering static and seismic design situations. Moreover, numerical simulations on two-dimensional (2D) frames, were first performed by means of the SAFIR program (Franssen J.-M. 2000), in order to study different fire scenarios acting in the reference buildings and to evaluate the performance of different elements, i.e. composite beams, composite columns, and beam-to-column joints, under fire load for different times of exposure to fire. Five fire scenarios were considered. In the first one (FS1), fire acted only into a span in the first floor as illustrated in Figure 2a.
a) FS1 b) FS2 c) FS3 d) FS4 e) FS5 Figure 2 - Fire scenarios considered in thermal analysis
In the second one (FS2), fire acted on the whole first floor; this means that the first level columns and the first level beams were heated. In the third one (FS3), fire acted only into a span on the last floor. In the fourth one (FS4), fire acted on the whole fifth floor. Finally, in the fifth one (FS5) fire acted on the whole frame. The fire load followed the ISO 834 fire law. Figure 3 shows the evolution of the bending moment and of the axial force as a function of time at
various locations in the FE model for the case FS1. In detail the bending moment in column C shows an inversion in the sign when axial forces in the beams changes sign too. In fact, the increase in temperature causes first an increment of axial load in the beam (compression) to about 18 min owing to the presence of the column restraint. Afterwards, the reduction of stiffness of the columns subject to fire prevails and the axial force in the beam turns to tension, being similar to a “catenary” structure (with large displacement and deflection). The elongation of the beam due to increase in temperature and the different “restraint” effect offered by the column also causes a reversal of the sign of bending moment at mid-span of the beam which from sagging becomes hogging and then sagging again. The results of the analysis shows that the frame collapses because of the formation of a beam mechanism in the longest span involving the formation of three plastic hinges located at mid-span and at both beam ends near supports. Figure 4 shows column temperature distributions in the fire case FS1; it is possible to notice that the steel tubular section offers a practically negligible protection to internal concrete; therefore temperatures of steel elements directly exposed to fire are equal to that of the external atmosphere.
B C110
101
151 155 160 161 165 170
A
Column C - first floor
-1500.00
-1000.00
-500.00
0.00
500.00
1000.00
0 5 10 15 20 25 30
Time (min)
M (k
Nm
)
elem. 101elem. 110
Beams A & B - first Floor
-800
-600
-400
-200
0
200
400
600
0 5 10 15 20 25 30
Time (min)
Axi
al L
oad
(kN
)
elem. 165
elem. 155compressione
Beam B - first floor
-800.00
-400.00
0.00
400.00
0 5 10 15 20 25 30
Time (min)
M (k
Nm
)
elem. 161elem. 165elem. 170
Figure 3 – Fire Case FS1: Bending Moment and Axial Load in beams and columns of the first
storey
10’ 30’ 40’ Figure 4 - Temperature distributions inside the column at different time. Fire Case FS1
The thermal distribution, for the composite slab with steel sheeting, evaluated after 10 min, 30 min and 40 min of ISO fire, by means of the thermal FE model created by SAFIR, are presented in Figure 5, while Figure 6 shows the thermal distribution for the case of the composite slab with the prefabricated slab. As expected, the temperature increment was higher in the composite slab with steel sheeting than in the slab with prefabricated R.C. elements. As a result, the moment resistance of the former slab was lower than in the latter owing to the higher reduction in material strength.
10’ 30’ 40’ Figure 5 – Fire case FS1. Evolution of the temperature distribution inside beam B with a steel
sheeting composite slab.
10’ 30’ 40’ Figure 6 – Fire case FS1. Evolution of temperature distribution inside beam B with prefabricated
elements
DESIGN OF BEAM-TO-COLUMN JOINTS The beam-to-column joint design aimed at ensuring a joint overstrength with respect to the beam. The proposed bolted solution derives from a parent welded design (Bursi et al. 2008) and was conceived to guarantee easiness of assembly and to avoid problems related to welding on site. The joint comprised of two horizontal diaphragm plates and a vertical through-column plate attached to the tube by groove welds as indicated in Figure 7.
Column
18
182,3
14
18
23°
23° 23° 23°
23°
a=8 L=320
36
75 36
= 320 x 1094 x 14
230
660
66 66 66 66 66
= 660 x 1304 x 16
= 660 x 1304 x 16
= 320 x 1094 x 14
660
1717
168
1616
4610
846
a=8 L=320
VERTICAL PLATE
a=8 L=1475
1094
320
2x HORIZONTAL PLATES
502 300 5021304
230
200
36 75 75 75 572 75 75
3612
412
436
A A
1814
300
166
1304
Section A-A
4,3
R228,5
a=8 L=320 a=8 L=320
R228,5
A A
Figure 7 - Steel-concrete composite beam-to-column joint: a) horizontal and vertical plates; b)
Nelson studs in columns
The flanges and the web of each beam were connected to the horizontal plates and the vertical plate by two and three rows of bolts M27 10.9 (1-1/8 A490), respectively. Apart from the slab type, also the joints differed due to the presence of strengthening plates with holes required by Eurocode 8 to avoid brittle fracture in the net sections of bolted joints. Thus, specimens JB-P1 and JB-S1, see table 1, where endowed with extra plates; whereas JB-P2 and JB-S2 had no plate being non-dissipative joints. Beam-to-column joint design was developed using the component method. The joint was simulated by a series of different components in agreement with Eurocode 3 par 1-8 and Eurocode 8 (UNI EN 1993-1-8, 2005, UNI EN 1998-1, 2005), achieving the necessary overstrength to the joint with respect to the connected composite beams. Stiffness and strength of complex components, like top and bottom plates or concrete slab in compression, were defined by means of refined Finite Element (FE) models of the joint including friction between the slab and the column. Depending on the level of friction, the distribution of the compression forces in the slab under sagging bending moment was found to be different: it is localized in front of the column for a friction coefficient equal to 0,35 as indicated in Figure 8a while it spreads over a more extended portion of the slab due to
increasing values of the friction coefficient. In detail, the diffusion angle becomes greater than 80 degrees for a friction coefficient of about 1. The results show that, in order to activate the transfer mechanisms proposed in Eurocode 8 (UNI EN 1998-1, 2005), i.e. Mechanism 1 front mechanism and Mechanism 2 strut and tie mechanism (see Figure 8b), it is necessary to increase the level of friction between the concrete slab and the composite column. As a result, Nelson 19 mm stud connectors welded around the column were used in the tested specimens as indicated in Figure 7. As a result beam-to-column joints were rigid full-strength joints satisfying the relation:
, , ,j Rd ov b pl RdM s Mγ≥ ⋅ ⋅ (1)
in which Mj,Rd is the resisting moment of full-strength beam-to-column joints and Mb,pl,Rd is the resisting moment of the composite beam (UNI EN 1998-1, 2005). Moreover, the ductile behaviour of joints was guaranteed by the following relationships:
beamRdploboltd RsR ,,, ⋅⋅≥ νγ (2)
platesRdplboltd RR ,,, ≥ (3)
RdbRdv FF ,, ≥ (4)
a) b)
80°
F Rd1
1/2F
Rd2
1/2F
Rd2
Figure 8 - Distribution of compression stresses in the slab for: (a) friction coefficient equal to
0,35; b) friction coefficient equal to 1
where { }25.1;min yt ffs = , 1.1=νγ o , boltdR , is the design strength of bolts, beamRdplR ,, is the
plastic design strength of the beam, platesRdplR ,, is the plastic design strength of the plates, while
RdvF , and RdbF , are the design shear strength and design bearing resistance of bolts, respectively. The following conditions were also fulfilled for joints JB-P1 and JB-S1 to avoid brittle fracture in the net sections, i.e.:
02
90.0
M
y
M
unet AffAγγ
≥ and 02
90.0
M
yf
M
unet fAfAγγ
≥ (5)
where fA is the area of the tension flange. Nevertheless, being non-dissipative joints, JB-P2 and JB-S2 did not. A 3D finite element model of the interior joint was implemented in the Abaqus 6.4.1 code (Hibbitt et al. 2000). Then beam-to-column joints subjected to fire load were designed. In particular, the component approach, exclusively applied to the moment-rotation-temperature behaviour, was adopted, in the absence of axial thrust owing to the thermal expansion restraint of the beam. A simplified model was derived, which can predict the moment-rotation-temperature characteristic of the joint. As a result, at high temperatures, the joints were designed to transfer shear forces owing to vertical loads from one beam to the other. Accordingly, vertical through column plate, top horizontal plate together with Nelson stud connectors welded around the column were arranged. In addition, two longitudinal steel rebars were added to the slab to reduce the damage produced by the seismic actions before fire.
EXPERIMENTAL TESTS
The experimental programme consisted of 4 tests under cyclic loading of full-scale substructures representing an interior full-strength bolted beam-to-column joint (Figure 9). Experimental tests were carried out at the Laboratory for Materials and Structures of the University of Trento. Joint specimens were subjected to cyclic loadings up to collapse, according to the ECCS stepwise increasing amplitude loading protocol, modified with the SAC procedure (ECCS 1986, SAC 1997) by using ey=0.005h=17.5 mm where h represents the storey height.
a)
A
A
285028505700
30 229023205002320
2805140280
4012
30
2670
2590
1470
( 310
+ 2
41.3
)55
012
001750
CFT
Ø 4
57 x
12
1344
IPE400IPE400
b)
1750
1750
1200
550
( 310
+ 2
41.3
)14
7080
2590
2670
1230
40
280 5140 2802320 500 2320229030
57002850 2850
A
A
CFT
Ø 4
57 x
12
IPE400 IPE400
1344
Figure 9: Bolted beam-to-column specimens a) slab with electro-welded lattice girders, b)
composite slab with profiled steel sheeting
The slab reinforcement, in the composite steel-concrete beams with steel sheeting, consisted of 3+3φ12 (3+3#4) longitudinal steel rebars in order to carry the sagging moment, and of 4+4φ12@100 mm (4+4#4@4 in.) and 7+7φ16@250 mm (7+7#5@10 in.) transverse steel rebars, in order to enable development of the seismic slab-to-column transfer mechanism as well as the resistance to the shear force. A mesh φ6@200x200 mm (φ6@8x8 in) is also present as shown in Figure 10a. The concrete class was C30/37 (4350 psi) while the steel grade S450 (A615 Grade 60) was adopted for reinforcing steel bars. In the case of the concrete slab prefabricated R.C. elements, the slab reinforcement was made up of 3+3φ12 (3+3#4 in.) longitudinal steel rebars and by 5+5φ12@100 mm (5+5#4@4 in.) plus 8+8φ16@200 mm (8+8#5@8 in.) transverse steel rebars. The same mesh was adopted as for the composite slab (Figure 10b).
a)
mesh Ø 6 / 200x200
880
pos.B 3+3 Ø 12
720
560
pos.C 7 Ø 16 / 250
pos.A 4 Ø 12 / 100
1500150300160
HD
620
mes
h
3Ø12
3Ø
12
2000
b)
mesh Ø 6 / 20x200
880
pos.B 3+3 Ø 12
720
560
pos.C 8 Ø 16 / 200
pos.A 5 Ø 12 / 100
400160 1400150
1Ø8
HD
620
mes
hH
D 1
0/14
/8 h
=9,5
cm
1Ø10
1Ø10
pos.
Dpo
s.H
3Ø12
pos
.Bpo
s.E
1Ø10
pos.
D
1Ø10
pos.
H
2000
Figure 10 – Rebars layout in: a) steel sheeting slab; b) prefabricate lattice girder slab
The columns were concrete-filled columns with a circular hollow steel section with a diameter of 457 mm (18 in.) and a thickness of 12 mm (0.438 in). Steel grade is S355 (A615 Grade 50). The column reinforcement consisted of 8φ16 (8#5) longitudinal steel rebars and of stirrups φ8@150 mm (φ8@6 in.) (Figure 11). The concrete class was C30/37 (4350 psi), while the steel grade S450 (A615 Grade 60) was adopted for the reinforcing steel bars. Actual values better than
nominal ones can be found in (Bursi et al. 2008). A scheme of the test set-up is shown in Figure 12.
40pos.D 8 Ø 16
pos.E Ø 8 / 15 cm
95
325135
50 2490
150150
150120
2670
90
Figure 11 - Column and column reinforcement
LVDTsINCLINOMETERS I
L
L
LL L
L
LI I I I
SG
SG
SG
SG
SG
SG
SG
SG
BEAM
OMEGA STRAING GAUGELVDTs
STRAING GAUGESGOL
LL LL
LLLLL
SG
SG
SG
SGSG
SG SG
OOO
O
BEAM
4
INFERIOR PLATE
SUPERIOR PLATE
SLAB
Figure 12 - Lateral view of the test set-up and main instrumentation on specimens
No vertical actuator was imposed on the column in order to be consistent with tests in other labs. In the tests, the following instrumentation was employed as illustrated in Figure 12: a) 5 inclinometers measured the inclinations: of the column at the top and, in the zone adjacent to the joint, and of the beams near the connection; b) 4 LVDTs detected the interface slip between the steel beam and the concrete slab and between the bottom horizontal joint plate and the flange of beam; c) 2 LVDTs were employed in order to measure the bottom horizontal joint plate deformations; d) 10 LVDTs were utilized in order to measure concrete slab deformations in the zone around the column; e) 4 Omega-shaped strain gauges detected the deformations of the concrete slab; f) 8 strain gauges monitored axial deformations of the reinforcing bars to scrutinise the effective breadth of the reinforcing bars at each loading stage; g) 4 strain gauges monitored deformations of top and bottom horizontal plates; h) 4 strain gauges recorded flange strains in order to estimate internal forces in steel beams; i) 2 load cells were set on the top of pendula and were utilized in order to measure the horizontal and vertical components of the forces; l) 1 digital transducer (Heidenhein DT500) was used in order to measure the top column displacement.
RESULTS AND SIMULATIONS The observed response clearly indicated that all specimens exhibit a good performance in terms of resistance, stiffness, energy dissipation and local ductility. Both the overall force-displacement relationship and the moment-rotation relationships relevant to plastic hinges formed in composite beams exhibited a hysteretic behaviour with large energy dissipation without evident loss of resistance and stiffness, as seen in Figures 13 to 15. Hysteretic loops of moment-rotation relationship are unsymmetrical owing to the different flexural resistance of the composite beam under hogging and sagging moments as can be observed in Figure 15. It was evident that owing to local buckling effects, severe strength degradations appeared to exceed 20 per cent of maximum strength values with rotations of about 40 mrad. The collapse of all specimens was caused by plastic buckling in compression of the bottom beam flange near the horizontal plate as illustrated in Figure 16 or fracture in tension.
Specimen JB-P1
-1000
-500
0
500
1000
-8 -6 -4 -2 0 2 4 6 8
Interstorey Drift [%]
Forc
e [k
N]
Specimen JB-P2
-1000
-500
0
500
1000
-8 -6 -4 -2 0 2 4 6 8
Interstorey Drift [%]
Forc
e [k
N]
Figure 13 - Specimens JB-P1 and JB-P2 – Force-Displacement relationship
Specimen JB-S1
-1000
-500
0
500
1000
-8 -6 -4 -2 0 2 4 6 8
Interstorey Drift [%]
Forc
e [k
N]
Specimen JB-S2
-1000
-500
0
500
1000
-8 -6 -4 -2 0 2 4 6 8
Interstorey Drift [%]
Forc
e [k
N]
Figure 14 - Specimens JB-S1 and JB-S2 - Force-Displacement relationship
Specimen JB-S1
-1000
-750
-500
-250
0
250
500
750
1000
-80 -60 -40 -20 0 20 40 60 80
φ [mrad]
M [k
Nm
]
M (plastic hinge at dx)M (plastic hinge at sx)
Figure 15 - Moment-rotation relationship of the
plastic hinges for the JB-S1 Figure 16 – Specimen JB-S1, plastic hinge in
the composite beam
Table 1 reports some key experimental data for each test: maximum applied displacement d, maximum value F of the force, maximum values of both sagging and hogging moments M, total number Ntot of cycles performed during the tests. Joint behaviour is essentially symmetric in terms of force, while, for all specimens, sagging moments are about 40 per cent larger than hogging moments. Figure 17 shows a comparison between Force-Interstorey drift curves for the specimens with and without plate strengthening of the horizontal diaphragms in both prefabricated slabs and steel sheeting slabs. These reinforcing plates were introduced in order to satisfy the relationships in (5). In this case, in which dissipative zones were located at the end of the beams, beam-to-column joints were rigid full-strength and the introduced reinforcing plates did not influence the specimen behaviour. Moreover, beam-to-column joints subjected to fire load were analyzed. In particular, the component approach, exclusively applied to the moment-rotation-temperature behaviour, was adopted, in the absence of axial thrust owing to the thermal expansion restraint of the beam. A simplified model was derived, which can predict the moment-rotation-temperature characteristic of the joint.The method is capable of predicting the variation of failure modes owing to change in the joint geometrical and material properties, as well as loading conditions. The temperature distribution into the composite joint subjected to three different time-temperature curves, as depicted in Figure 18, was implemented in ABAQUS 6.4.1 software (Hibbitt et al., 2000) and a 3D FE model of joint was studied. Figure 19 shows the mean temperature distribution on the slab as a function of the time of fire exposure obtained from the application of the three inputs. It is evident that the application of the natural fire curve
produces a lower rise in temperature in the concrete slab than the application both of the ISO 834 curve and parametric curve proposed in EC1 (UNI EN 1991-1-2 2004).
JB-P1 vs. JB-P2
-800
-400
0
400
800
-8 -6 -4 -2 0 2 4 6 8
Inter-Storey Drift [%]
Forc
e [k
N]
BJ-P1BJ-P2
JB-S1 vs. JB-S2
-800
-400
0
400
800
-8 -6 -4 -2 0 2 4 6 8
Inter-Storey Drift [%]
Forc
e [k
N]
BJ-S1BJ-S2
Figure 17 - Comparison between Force-Inter-storey drift curves
Table 1: major experimental data
Name Test Method
d [mm]
F [kN]
*M [kNm] Ntot Type of Specimen
JB-P1 Cyclic 210 +655.09 -680.92
+1047.2 -747.75 22
Specimen with electro-welded lattice girders slab and Nelson connectors around the column, with reinforcement plates
JB-P2 Cyclic 210 +666.70 -657.60
+892.60 -760.23 20
Specimen with electro-welded lattice girders slab and Nelson connectors around the column, without reinforcement plates
JB-S1 Cyclic 210 +647.12 -639.66
+760.09 -537.59 20
Specimen with profiled Steel Sheeting slab and Nelson connectors around the column with reinforcement plates
JB-S2 Cyclic 175 +627.19 -634.58
+887.46 -636.37 19
Specimen with profiled Steel Sheeting slab and Nelson connectors around the column, without reinforcement plates
* + Sagging Moment; - Hogging Moment
0200400600800
10001200
0 10 20 30 40 50 60 70 80 90Time [min]
Tem
pera
ture
[°C
]
ISO 834 EC1- Annex A OZone V2.2
050
100150200250300
0 5 10 15 20 25 30 35 40 45 50 55 60Time [min]
Tem
pera
ture
[°C
]
Curve ISO 834EC1-Annex AOZone
Figure 18 - Different time-temperature curves applied as a function of the time of fire exposure
Figure 19 - Mean temperature distribution on the slab as a function of the time of fire exposure
As a result, Figure 20 shows that all the steel parts exposed to fire increase their temperature very fast, reaching a very high temperature after only 15 minutes. Conversely, in the concrete and in the steel component with a concrete cover rebars, as in the horizontal plates and the vertical plate (passing through the column) the temperature does not increase so quickly, remaining near the value of the ambient temperature. On this basis, it was possible to incorporate degradation characteristics of the material into the mechanical component model based on the “Strength Reduction Factor-SRF”, which is actually a strength retention factor present in the current European Design codes (CEN, Eurocode 3-1-2, 2003).
Figure 20 - Temperature distribution in the beam-to-column composite joint with prefabricated
slab
This is basically the residual strength of steel and concrete materials at a particular temperature relative to its basic yield strength at room temperature. Figure 21 shows the reduction of the design moment capacity of the joint as a function of the time of fire exposure.
0200400600800
100012001400
0 0.2 0.4 0.6 0.8 1
M [k
Nm
]
Rotation [rad]
Sagging Moment (Ozone)t=20°C t=15 min -1400
-1200-1000-800-600-400-200
0-3-2.6-2.2-1.8-1.4-1-0.6-0.2
M [k
Nm
]
Rotation [rad]
Hogging Moment (Ozone)t=20°C t=15 min
Figure 21 - Moment-rotation relationship of the joint as a function of the time of fire exposition
An attentive reader can observe how the hogging moment capacity reduces to approximately 90-95 per cent of its initial value after 15 minutes of exposure. The degradation in resistance and stiffness is observed both for the sagging and the hogging bending moment, respectively; nevertheless these numerical results show that beam-to-column joints are able to carry the internal action that beams transfer to columns for a time of 15 minutes without any passive fire protection. This period of time is considered enough to evacuate the building after a severe earthquake. Clearly, these simulations need to be confirmed by fire tests; in this respect, the companion full strength beam-to-column welded solution investigated by Bursi et al. (2008), was shown to be successful.
CONCLUSIONS The paper presented part of the results of a numerical and experimental study aimed at developing performed-based design methods for beam-to-column joints, by considering their structural performance under seismic-induced fire loading. In particular, the paper discussed the experimental program carried out on steel-concrete composite beam-to-column joints together with numerical simulations regarding their seismic-induced fire behaviour. Experimental and numerical results showed how joint details influence beam-column sub-assemblage responses. All sub-assemblages exhibited rigid behaviour for the designed composite joints and favorable performance in terms of resistance, stiffness, energy dissipation and local ductility; as expected plastic hinges developed in beams as a consequence of the capacity design. Moreover, a behaviour factor of about 4 was observed for the composite frames analysed. As a results, the joint can be used in Ductility Class M structures according to Eurocode 8. Numerical fire simulations have shown that the joint is able to carry the internal action for a maximum time of 15 minutes without any passive fire protection: this period of time is enough to evacuate the building after a severe earthquake. Moreover, joints endowed with prefabricated slab exhibit a better behavior compared to joints endowed with composite slabs.
ACKNOWLEDGEMENTS The experimental programme was made possible by the grant provided under the RELUIS National programme for seismic engineering: Task n.5 - "Development of innovative approaches on the design of steel and steel-concrete composite structures.” The authors express their gratitude to Dr: Marco Molinari and the technicians of the Structural Testing Laboratory of Trento University, who skilfully prepared and followed the tests.
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