A multi-objective advanced design methodology dealing with seismic actions ... together with FE simulation demonstrated the adequacy of the advanced design.
A Multi-Objective Advanced Design Methodology of Composite Beam-to-Column Joints Subjected to Seismic and Fire Loads Raffaele Pucinottia, Fabio Ferrariob, Oreste S. Bursib a
Department of Mechanics and Materials, Mediterranean University of Reggio Calabria, loc. Feo di Vito, Reggio Calabria, 89126, Italy b Department of Mechanical and Structural Engineering, University of Trento, via Mesiano 7, Trento, 38050, Italy
Abstract. A multi-objective advanced design methodology dealing with seismic actions followed by fire on steel-concrete composite full strength joints with concrete filled tubes is proposed in this paper. The specimens were designed in detail in order to exhibit a suitable fire behaviour after a severe earthquake. The major aspects of the cyclic behaviour of composite joints are presented and commented upon. The data obtained from monotonic and cyclic experimental tests have been used to calibrate a model of the joint in order to perform seismic simulations on several moment resisting frames. A hysteretic law was used to take into account the seismic degradation of the joints. Finally, fire tests were conducted with the objective to evaluate fire resistance of the connection already damaged by an earthquake. The experimental activity together with FE simulation demonstrated the adequacy of the advanced design methodology. Keywords: Earthquake, Steel-concrete composite joint, full strength joints, seismic actions, fire actions, FE model.
INTRODUCTION Major earthquake in urban areas have often been followed by significant conflagration that have been difficult to control and have resulted in extensive damage to property. Major contributing factors have been identified as accidental ignition due to earthquake shaking, external fire spread through vegetation and inadequate building separation, earthquake damage to building’s fire safety systems, loss of water supplies for fire suppression, and the lack of intervention by fire fighters due to inadequate resources and obstructed access to the site of the fire. Earthquake, then, increases the risk of loss of life if a fire occurs within the building. It is obvious therefore that fire after earthquake is a scenario that should be properly addressed in any performance base design in locations where significant earthquake can occur. The science around predicting the likelihood of such events is so complex that is not considered a design requirement; however given that an important part of Europe is a seismic prone area, the combination of a fire after an earthquake poses a very real danger. Fire and earthquake are accidental actions and have been treated most often as independent events, which each considered separately. However, if the occurrence of a fire does
not increase the probability of an earthquake, the opposite is not true. As Kobe earthquake tragically demonstrated in 1995, a city can be devastated by numerous fires following the earthquake. Post earthquake fire is a low probability event but with high potential consequences, and it is appropriate that it be included in the risk analysis. According to the modern seismic codes, ordinary structures are designed to suffer damage to some extent during strong earthquake, exploiting the structure own ductility to avoid collapse and safeguard human lives. Then, a fire coming soon after an earthquake will find a different, more vulnerable, structure with respect to the initial, undamaged, one. Depending on the extent of damage, the fire resistance rating of the structure could be significantly reduced. In this paper a multi-objective advanced design methodology dealing with seismic actions followed by fire on steel-concrete composite full strength joints with concrete filled tubes is proposed. The data obtained from experimental tests have been used to calibrate a model of joint and a hysteretic law was used to take into account the seismic degradation. Subsequently seismic simulations on several moment resisting frames were performed and fire tests were conducted on both undamaged and damaged specimens.
BEAM-TO-COLUMN JOINT In order to obtain the values of the actions used in the design of the proposed joints, two moment resisting frames having the same structural typology but different slab systems were analyzed: a composite steel-concrete slab with structural profiled steel sheeting and a concrete slab composed of electro-welded lattice girders (Figure 1).
a) b) c) FIGURE 1. Relevant Plains, Structure typology and Elevation of five-storey MR Frame; a) slab with profiled ribbed steel sheeting; b) slab with prefabricate lattice girder, c) elevation.
The composite steel-concrete office-building had 5 floors with 3.5 m storey height. It was made up by three moment resisting frames placed at the distance of 7.5 m each in the longitudinal direction, while it was braced in the transverse direction. In particular, the main moment resisting frame is made up by two bays spanning 7.5 m and 10.0 m in the solution with steel sheeting with a distance between secondary beams equal to 2.5 m; and by two bays spanning 7.0 m and 10.5 m in the solution with lattice steel girders with a distance between secondary beams equal to 3.5 m. The main beams were IPE 400 while the secondary ones were IPE 300. The columns were concrete-filled column with a CFT filled tubular column steel profile with a diameter of 457 mm and a thickness of 12 mm. Buildings were designed to have a dissipative
structural behaviour according to point 7.1.2 of Eurocode 8 [4]. The effective width of the slab in composite beams was assigned according to Eurocode 4 (point, 5.4.1.2) [2, 3] for static and fire analyses and Eurocode 8 (point, 7.6.3 [4]) for seismic analyses. Beam-to-column connections were assumed to be rigid (5.1.2-EC4 [2]). Moreover, the influence of second-order effects on the determination of the action effects have been neglected since the criterion given in 5.2.1(4)-EC3 [1]. This allowed also to neglect member imperfections in composite compression members within global analysis. The welded joints were made by two horizontal diaphragm plates split into two equal halves along the diagonal for fabrication convenience and easiness of assembling. Once each side were properly placed on the pipe the two halves are attached with a full-joint penetration groove weld. In detail, the inferior plates are welded to column in the shop, while the superior plates are welded on site as should be understood from Figure 2. Joint
Horizontal plates
475
5
60
40 R2 28 ,
100
= 125 x 200 x 12 A4
200
25
125
20x20
= 300 x 700 x 14 A6
300
14
5 8,
18
2 R2
4,3
18
4,3
14
700
,5 28 R2
50 150
°
°
23
625 1250
Stiffeners Vertical plate plate
23° 23 ° 23
33°
475
50
1250
23°
300
625 16
300
= 300 x 700 x 14 A6
300 369
19
700
= 290 x 1250 x 16 A 5A
230 290
121,5 16
121,5
FIGURE 2. Detail of welded steel-concrete composite joint. (Dimensions in mm).
EXPERIMENTAL PROGRAMME The experimental programme concerned the execution of 6 tests on full-scale substructures representing the interior welded beam-to-column joint in seismic conditions (Table 1). Specimens (Figure 3) were designed and fabricated according to Eurocode 4 [2, 3] and Eurocode 8 [4] provisions. In all composite specimens the connections between steel beam and desk were made by Nelson 19 mm stud connectors with an ultimate tensile strength fu=450 MPa. Additional Nelson 19 mm studs were localized around the column in order to enforce a better connection between the column and the composite slabs (Figure 2). Experimental tests were carried out at the Laboratory for Materials and Structures of the University of Trento. The scheme of the experimental set-up was detailed in [6] together with the employed instrumentation, while the geometric property specimens are shown in the Figure 3. The joint specimens were subjected to monotonic and cyclic loadings up to collapse, according to the ECCS stepwise increasing amplitude loading protocol, modified with the SAC procedure [4]: ey=17.5 mm. All specimens exhibit a good performance in terms of resistance, stiffness, energy dissipation and ductility. Both the overall forcedisplacement relationships and the moment-rotation relationships relevant to plastic hinges formed in the composite beams exhibit a hysteretic behaviour with large energy dissipation without evident loss of resistance and stiffness (Figures 4, and 5). From the
results of monotonic tests, it was observed that the joints exhibited elasto-plastic behaviour with limited degradation. There was practically no difference in the monotonic responses between the joint with steel sheeting slab and joint with prefabricate slab. A 150
A
IPE 400
CFT Ø 457 x 12
IPE 400 CFT Ø 457 x 12
A
A
a)
b)
FIGURE 3. Welded Beam-to-Column joint specimens; a) endowed with steel sheeting; b) endowed with prefabricate girder Specimen WJ-P1
Specimen WJ-P1 1000
200
-14
-10
-6
-200 -2
2
6
10
14
M [kNmm]
Force [kN]
600
-80
-60
-40
-600 -1000
1000
200 -200 -2
2
6
10
14
M [kNmm]
Force [kN]
600
-6
-80
-60
-40
-600 -1000
1000
200 -6
2
6
10
14
M [kNm])
Force [kN]
600
-10
-80
-60
-40
-600 -1000
1000
200 -200 -2
2
-600 -1000
Displacement [e/ey]
6
10
14
M [kNm]
Force [kN]
600
-6
1000 800 600 400 200 0 -20-200 0 -400 -600 -800
20
40
60
80
Left Hinge Right Hinge
1000 800 600 400 200 0 -20-200 0 -400 -600 -800
20
40
60
80
Left Hinge Right Hinge
Specimen WJ-S2
Specimen WJ-S2
-10
Right Hinge
φ [rad]
Displacement [e/ey]
-14
80
Specimen WJ-S1
Specimen WJ-S1
-14
60
Left Hinge
φ [rad]
Displacement [e/ey]
-200 -2
40
Specimen WJ-P2
Specimen WJ-P2
-10
20
φ [rad]
Displacement [e/ey]
-14
1000 800 600 400 200 0 -20-200 0 -400 -600 -800
-80
-60
-40
1000 800 600 400 200 0 -20-200 0 -400 -600 -800
φ [rad]
20
40
60
80
Left Hinge Right Hinge
FIGURE 4. Cyclic Test results of Specimen WJ: Force vs. Displacement and Moment vs. Rotation curves
No
Name
1
WJ-P1
Test Method Cyclic
2
WJ-P2
Cyclic
3
WJ-S1
Cyclic
4
WJ-S2
Cyclic
5
WJ-PM
Monotonic
6
WJ-SM
Monotonic
TABLE 1. Experimental programme. Type of Specimen Specimens with electro-welded lattice girders slab and no Nelson connectors around the column, without reinforcement plates Specimens with electro-welded lattice girders slab and Nelson connectors around the column, without reinforcement plates Specimens with electro-welded lattice girders slab and no Nelson connectors around the column, without reinforcement plates Specimens with profiled Steel Sheeting slab and Nelson connectors around the column, without reinforcement plates Specimens with profiled Steel Sheeting slab and no Nelson connectors around the column, without reinforcement plates Specimens with profiled Steel Sheeting slab and no Nelson connectors around the column, without reinforcement plates
Specimen WJ-PM
Specimen WJ-PM
1000
Right Plastic Hinge 1000
800
Force [kN]
Lift Plastic Hinge
M [kNm]
600 400
200
-160
200
-120
-80
-40-200 0
40
80
120
160
-600
0 0
6
12
18
Displacement [e/ey]
24
30
-1000
φ [mrad]
Specimen WJ-SM
Specimen WJ-SM Right Plastic Hinge
1000
1000
Lift Plastic Hinge
800
600
600
M [kNm]
Force[kN]
600
400
-160
200
200 -120
-80
-40-200 0
40
80
120 (-M) 160
-600
0 0
6
12
18
Displacement [e/ey]
24
30
-1000
φ [mrad]
FIGURE 5. Specimen WJ: Force vs. Displacement and Moment vs. Rotation curves
COMPUTER FE MODELS FE Model of the Specimens The data obtained from monotonic and cyclic experimental tests have been used to calibrate a model of the joint in order to perform seismic simulations on several moment resisting frames. The model used to perform these analyses is based on two parallel springs at the end of the beam connected to a rigid panel which represents the rigid connection with the column. Each spring is used to match the properties of the beam under sagging and hogging bending moment. Another spring is used to connect this panel with the shear panel of the column (see Figure 6). The main purpose of this spring is to take into account the shear deformation of the connection in order to improve the calibration of the model.
The calibration of the springs used in the model, has been done twice for each type of slab, one for the experiments performed without connectors in the upside of the top plate of the connection and other for the experiments performed with connectors. IDARC program [9] was used to perform the simulation. A hysteretic law has been used to take into account the seismic degradation of the joint according to a modification of the Bouc-Wen model implemented by Silvaselvan-Reinhorn in IDARC [10]. The real measured properties of the concrete and steel have been used in the computer model in order to match as accurately as possible the results of the experimental tests [6]. The behaviour has been validated in two ways: a) the amount of energy dissipated by the whole model and by the plastic hinge of the composite beam during the computer simulations have to match the energy dissipated during the laboratory tests by the same elements; b) the hysteretic behaviour exhibited in the computer simulations has to be similar to that shown in the laboratory tests. 198,25
457 99,125
236,5
473
236,5
99,125
99,125
99,125
a) FIGURE 6. a) Computer model of specimens; b) Joint details of computer model
b)
The hysteretic behaviour exhibited in the computer simulations has to be similar to that shown in the laboratory tests (Figure 7). Once the behaviour of the joints have been calibrated, two frames have been modelled in IDARC [9] for both type of slabs in order to evaluate its dynamic properties. In the application the Park and Ang [7] damage index modified by Chai and Romstad [8] was considered. Confronto Forze Spostamenti Martinetto
Confronto Cerniera Plastica Sinistra
800
1200000 1000000
600
800000 400 600000 400000
-2.50E+02
-2.00E+02
-1.50E+02
-1.00E+02
-5.00E+01
0 0.00E+00
5.00E+01
1.00E+02
1.50E+02
2.00E+02
M (kNmm)
F (kN)
200
200000
-6.00E-02 -200
-800 delta (mm) IDARC
Sperimentale
-2.00E-02
0 0.00E+00
2.00E-02
4.00E-02
6.00E-02
8.00E-02
-200000 -400000
-400
-600
-4.00E-02
IDARC Exp
-600000 -800000 -1000000 Ø (rad) IDARC
Sperimentale
FIGURE 7. a) V- δ curve of the whole model, interior joint, prefabricated slab; b) M-ø curve of the beam, interior joint, prefabricated slab. Sx plastic hinge
FE Model of the Structures The structural behaviour under vertical static loads and under seismic loads has been simulated numerically, utilizing a bi-dimensional model for the principal frame (Figure 1). Incremental dynamic analyses have been carried out with the four frames (two frames for each type of slab). Three incremental time history analyses have been performed with each frame. The accelerograms used to perform the simulation have
a/g
Sa/g
been artificially generated to match Type 1 response spectrum and soils type A, type B and type D ( Figure 8a). Each accelerogram has been used to perform twenty time history analyses with different peak ground accelerations ranging from 0.1g to 2.0g with an interval of 0.1g, thus resulting in an incremental dynamic analysis. Additional simulations were also done to identify the peak ground acceleration when the first yielding occurred and when the 35 mrad plastic rotation was reached. 1.60
1.20
0.50 0.40
TYPE D TYPE B TYPE A
1.40
0.30 0.20
1.00 0.10 0.80
0.00
0.60
-0.10
0.40
-0.20
0.20
-0.30
0.0
2.0
4.0
6.0
8.0
10.0
12.0
14.0
16.0
-0.40
0.00 0.0
0.5
1.0
1.5
2.0
2.5
3.0
3.5
4.0
T [s]
-0.50
T [s]
FIGURE 8. a) Response spectrums of accelerograms matching soil type A, B and D of Type 1 EC8 spectrum.; b) Time history of artificial accelerogram matching soil type A of Type 1 EC8 spectrum.
V/W
The behaviour factor has been calculated using the equation: pgau q dyn = (1) pga y where pgau corresponds to the peak ground acceleration that induces a first plastic rotation in a beam-to-column joint of about 35 mrad; while pgay corresponds to the peak acceleration that produces the first yielding in the structure. Capacity curves and the main behaviour factors obtained from IDA analysis are shown in Figure 9 and in the Table 2 respectively. 1.4 1.2 1.0
1_storey 2_storey 3_storey 4_storey 5_storey Frame
0.8 0.6 0.4 0.2 0.0 0.00%
0.50%
1.00%
1.50%
2.00%
2.50%
3.00%
3.50%
4.00%
Drift [%]
FIGURE 9. IDA capacity curves, with soil Type D, for steel sheeting frame without connectors
Steel sheeting slab Prefabricated slab
TABLE 2. Behaviour factors from IDA analysis Without Connectors With Connectors IDA A IDA B IDA D IDA A IDA B IDA D 3.82 3.57 3.58 4.24 4.10 4.00 3.17 3.42 3.48 3.75 3.94 5.03
Largest values of the damage index for the all studied frames, under a seismic excitation of 0.4 g, are 0.503 for external joints and 0.419 for internal joints. The damage in joints is limited and repairable [12].
FIRE RESPONSE In order to accurately simulate the damage, calculated in the previous session, due to a seismic event, a specified deformation was applied on the specimens. In particular, only one of each type of connections (Steel sheeting slab and prefabricated lattice slab) was subject to the damage test while all specimens were fire tested. The purpose of the fire tests was to determine the residual fire resistance of a damaged connection subject to fire limit state loading. Fire tests were conducted, at Cardington by BRE, with asymmetric loading on the internal connections representing the load in the column of the frame from primary beams of different length (Figure 1).
Fire test results Undamaged Specimen endowed with steel sheeting slab. Figure 10a shows both the west and the east load applied and the average atmosphere temperature evolution, while Figure 10b shows the atmosphere temperatures relative to the standard fire curve; figure 10c shows instead the temperature distribution on the composite beam. The fire test led to full depth cracking of the slab, runaway deflection (Figure 10 d). The test was terminated due to runaway deflection after approximately 40 minutes. West Load EastLoad Temp.
150
2000
1200 1000
1500
100
1000
50
500
800
Burner
600 400 200
0
0 0
20
40
60
80
100
Time [min]
0 0
a)
20
40
60
Time [Min]
80
100
b)
Steel Sheeting Slab-Undamaged
TC7 TC8 TC9 TC10 TC11 TC12 TC13 Av. Atmos
80
Vertical Deflection [mm]
Temp. [°C]
Steel Sheeting Slab-Undamaged 800 700 600 500 400 300 200 100 0
G3 G4 G1 G2 SFC Av. Atmos
Steel Sheeting Slab-Undamaged
Steel Sheeting Slab-Undamaged
Temp. [°C]
Load (KN) Temp. [°C]
200
60
C1061 C1062 C1063 C1064
40 20 0 -20
0
200
400
600
800
+VE Denotes downwards deflection
0
20
40
60
80
100
-40
Temp. [°C] c) d) FIGURE 10. Specimen endowed with Steel sheeting slab, a) The west and the east load applied and the average atmosphere temperature b) Atmosphere Temperatures; c) East Beam Temperature; d) Temperature vs. Vertical Deflection Time [Min]
Damaged Specimen endowed with steel sheeting slab. The test had to be terminated after approximately 34 minutes due to runaway deflection. The atmosphere temperature is shown in Figure 11a while Figure 11b shows the measured column temperatures. Following the fire test the profiled steel sheeting had separated from the slab, the slab had cracking both along the surface and through the depth of the slab and there was extensive buckling of both the lower flange and the web of the composite beam in the area of the connection. Undamaged Specimen endowed with prefabricated lattice slab. For this specimen the test continued through to 60 minutes. At this points the rate of deflection was
increasing but had not yet reached the point where it was not possible to maintain the applied load. There was no permanent deformation and no sign of any significant damage from the fire test. The critical temperatures and the temperature-deformation chracteristics are shown in Figure 12. Steel Sheeting Slab-Damaged
Serie1
1000
Serie3
800
Serie4
600
SFC
400
Burner
200 0 0
10
20
30
40
TC1 TC2 TC3 TC4 TC5 TC6
Steel Sheeting Slab-Damaged
Temp [°C]
Serie2
Temp. [°C]
1200
50
800 700 600 500 400 300 200 100 0 0
5
10
15
20
25
30
35
40
45
50
55
Time [Min] a) b) FIGURE 11. Damaged Specimen endowed with Steel sheeting slab, a) Atmosphere Temperatures; b) Column Temperature;
Time [Min]
Prefabricated Slab-Undamaged 2000
150
1500
100
1000
50
500 0
0 0
20
40
60
80
C1021 C1022 C1023 C1024 C1025 C1026 C1027
1000 800
Temp. [°C]
200
Load (KN) Temp. [°C]
Prefabricated Slab-Undamaged
West Load EastLoad Temp.
600 400 200 0
100
0
Time [min]
20
40
60
80
100
120
Time [Min] a) b) FIGURE 12. Undamaged Specimen endowed with Prefabricated slab, a) The west and the east load applied and the average atmosphere temperature b) East Beam Temperature
Damaged Specimen endowed with prefabricated lattice slab. The behaviour of the prefabricated lattice slab specimen which had been subject to the damage test was very similar to that for the undamaged specimen. Hogging Moment [kNm]
1000 800 600 400
Steel Sheeting Slab temp=20°C Abaqus - after 30' of exposure Experimental - after 30' of exposure - Undamaged Experimental - after 30' of exposure - Damaged
200 0 0.00
2.00
4.00
6.00
Φ [mrad]
8.00
10.00
a) b) FIGURE 13. Specimen endowed with Steel Sheeting Slab a) Comparison among the resistance of undamaged and damaged joint estimated at 20°C and after 30’of fire exposition ; b) Abaqus FE model
With reference to the undamaged and damaged specimens endowed with steel sheeting slab, in Figure 13a the comparison among the resistance estimated at 20°C and after 30 minute is shown. In particular the value of the bending moments were calculated using the component method, considering both the experimental and the preliminary numerical temperature distribution in the joins. In this last case the Abaqus 6.4.1 code [11] was employed to conduct the thermal analysis [6]. The hogging capacity moment becomes approximately 38% of the initial value after 30 minute of fire exposition in the case of the undamaged specimen and 31% in the case of the damaged specimen. Instead, the capacity hogging moment, considering the
temperature distribution obtained by ABAQUS code (Figure 13b), becomes approximately 24% of the initial value.
CONCLUSIONS A multi-objective advanced design methodology dealing with seismic actions followed by fire on steel-concrete composite full strength joints with concrete filled tubes is proposed in this paper. Due to the fact the joint are full-strength and rigid, the energy dissipating mechanisms in moment resisting frames with this type of joint will entirely rely on the formation of plastic hinges at the end of the beams. The study showed an optimum behaviour of the joint components till a damage index of Di=0.5. A behaviour factor of about 4 has been observed for the analyzed composite frames. The fire tests showed that there was no noticeable difference in the performance between the damaged and undamaged specimens; moreover these demonstrated the effectiveness of the strategy based on the protection of steel connection details by means of the exploitation of concrete both in the column and in the slab.
ACKNOWLEDGMENTS This research is developed within the financial support of UE under the project PRECIOUS (RFS-CR-03034) which the writers are grateful.
REFERENCES 1. UNI EN 1993-1-1. 2005. Eurocode 3: Design of steel structures. Part 1: General rules and rules for buildings; 2. UNI EN 1994-1-1. 2005. Eurocode 4: Design of composite steel and concrete structures - Part 1.1: General rules and rules for buildings; 3. UNI EN 1994-1-2. 2005. Eurocode 4: Design of composite steel and concrete structures - Part 1.2: General rules – Structural fire design; 4. UNI EN 1998-1. 2005. Eurocode 8: Design of structures for earthquake resistance - Part 1: General rules, seismic actions and rules for buildings; 5. ECCS 1986. Recommended Testing Procedure for Assessing the Behaviour of Structural Steel Elements under Cyclic Loads. ECCS Publication n° 45; 6. Bursi O. S., Ferrario F., Pucinotti R., (2008) Steel-Concrete Composite full strength Joints with Concrete Filled Tubes: Design and Test Results, 12th International Symposium on Tubular Structures October 8 to 10, 2008 Shanghai, China; 7. Park Y. J. and Ang, A. H. S., (1985) Mechanistic Seismic Damage Model for Reinforced Concrete, Journal of Structural Engineering, vol.111 n.4; 8. Chai Y. H. and Romstad K. M., (1997) Correlation Between Strain-Based Low-Cycle Fatigue and Energy-Based Linear Damage Models, Earthquake Spectra, Vol. 13, No. 2, pp. 191-209; 9. Valles R.E., Reinhorn A.M., Kunnath S.K., Madan A., (1996) IDARC2D Version 4.0: A Program for the Inelastic Damage Analysis of Buildings. Technical Report NCEER-96-0010; 10. Silvaselvan M.V., Reinhorn A.M., (1999) Hysteretic Model for Cyclic Behaviour of Deteriorating Inelastic Structures, Technical Report MCEER-99-0018; 11. Hibbitt, Karlsson and Sorensen 2000. “ABAQUS User’s manuals”. 1080 Main Street, Paw-tucket, R.I. 02860. 12. Williams M. S., Sexsmith R. G., (1995), Seismic Damage Indices for Concrete Structures: A Stateof-the-Art Review., Earthquake Spectra, 11, 2, 319-349]
