Steel-concrete composite full strength joints with concrete filled tubes: design and test results

Bursi O. S., Ferrario F. Department of Mechanical and Structural Engineering, University of Trento, Trento 38050, Italy

Pucinotti R. Department of Mechanics and Materials, Mediterranean University of Reggio Calabria, Reggio Calabria, 89122, Italy

ABSTRACT: In this paper, a multi-objective advanced design methodology dealing with seismic actions fol-lowed by fire on steel-concrete composite full strength joints with concrete filled tubes is proposed instead of a traditional single-objective design where fire safety and seismic safety are achieved independently. Experi-mental tests together with numerical simulations of the fire behaviour were carried out to derive fundamental information about the performance of these joints. In detail a total of six specimens designed according to Eurocode 3, 4 and Eurocode 8 were subjected to monotonic and cyclic loadings up to collapse at the Labora-tory for Materials and Structures Testing of the University of Trento, Italy. These specimens were detailed in order to exhibit a favourable fire behaviour after a severe earthquake. The major aspects of the cyclic behav-iour of composite joints are presented and commented upon together with fire analysis results. Both the ex-perimental activity and the numerical FE simulations demonstrate the adequacy of the proposed joint design.

1 INTRODUCTION

In the design of steel-concrete composite buildings the sequence of seismic and fire loadings are not taken into account. In fact, the seismic safety and the fire safety are considered separately. In reality, the risk of loss of lives increases if a fire occurs within the building after an earthquake. Where significant earthquakes can occur, fire after earthquake is a de-sign scenario that should be properly addressed in any performance-based design.

This approach takes into account seismic safety and fire safety with regard to accidental actions as well as fire safety on a structure characterized by stiffness deterioration and strength degradation ow-ing to seismic actions. The proposed design solution, were developed in an European research (Colombo & Bursi 2006). This research program is intended to develop fundamental data, design procedures and promotion of two types of ductile and fire-resistant composite beam-to-column joints with:

1. partially reinforced-concrete-encased column with I-section;

2. concrete filled tubular column with circular hollow steel section.

The project analyses the scenario in which a fire follows and earthquake, thus defining joint typolo-gies for which, after being damaged by an earth-quake, a residual load-bearing capacity is assured during a fire occurring after an earthquake. The de-sign is performed in the modern context of perform-

ance-based engineering determining both the stiff-ness deterioration and the strength degradation of composite joints after seismic loading.

The present paper presents the results of experi-mental tests under monotonic and cyclic loading to-gether with numerical simulations of the fire behav-iour of steel-concrete composite full strength joints with concrete filled tubes.

2 COMPOSITE BUILDING DESIGN

The actions used in the design of the proposed joints were obtained by the analyses of two moment resist-ing frames having the same structural typology but different slab systems (160 mm composite steel-concrete slab high with structural profiled steel sheeting and 160 mm concrete slab high composed of electro-welded lattice girders).

The composite steel-concrete office-building was endowed with 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 longitudi-nal direction, while it was braced in the transverse direction. A different distance between the secon-dary beams was adopted for the two solutions that takes into account the different load bearing capaci-ties of the two slab systems as well as the need to avoid propping systems during the construction phase. As a result, 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 be-tween 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 secon-dary beams equal to 3.5 m (Figs 1, 2). All slabs were arranged in the direction parallel to the main frames. The main beams were IPE 400 while the secondary ones were IPE 300.

2.1 Composite Beams Two different types of composite beams were checked according to point 6.1.1 of Eurocode 4 (UNI EN 1994-1-1. 2005). In the first one, the beam section was an IPE400 with steel grade S355 while the deck was a composite slab with a prefabricated lattice girder made by the Pittini Group (Fig. 3).

Figure 1. Plan and structure typology; slabs with prefabricate lattice girders. Dimensions in metres.

Figure 2. Plan and structure typology, slabs with profiled steel sheetings. Dimensions in metres.

In this case, the slab reinforcement was per-

formed by 3+3φ12 longitudinal steel bars and by 5+5φ12@100 mm plus 8+8φ16@200 mm transver-sal steel bars. A mesh φ6@200x200 mm completes the slab reinforcement. In the second type of beam

as shown in Figure 4, the same beam section was adopted and a composite slab with profiled steel sheeting was made following the rules given in Sec-tion 9 of Eurocode 4 (UNI EN 1994-1-1. 2005).

The slab reinforcement consisted of 3+3φ12 lon-gitudinal steel bars and of 4+4φ12@100 mm and 7+7φ16@250 mm transversal steel bars. Moreover, the same mesh φ6@200x200 mm was adopted. The concrete class was C30/37 while the steel grade S450 was adopted for the reinforcing steel bars.

1Ø10

pos.

Dpo

s.H

3Ø12

pos

.Bpo

s.E

1Ø10

pos.

D

1Ø10

pos.

H

2000

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

Figure 3. Specimens WJ-P -Slab with prefabricate lattice gird-ers. Dimensions in mm.

HD

620

mes

h

3Ø12

3Ø

12

2000

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

Figure 4. Specimens WJ-S - Slab with profiled Steel Sheetings. Dimensions in mm.

2.2 Composite Columns The columns of steel grade S355 were concrete-filled column with a CFT filled tubular column steel profile with a diameter of 457 mm and a thickness of 12 mm. Column reinforcements consisted of 8φ16 longitudinal steel bars and stirrups φ8@150 mm.

The concrete class of composite columns was C30/37, while the steel grade S450 was adopted for the reinforcing steel bars as illustrated in Figure 5.

pos.D 8 Ø 16pos.E Ø 8 / 15 cm

95

325135

50 2490

150150

150120

2670

9040

Figure 5. Column and column reinforcements.

2.3 Beam-to-Column Composite Joints The following components were considered in the method: concrete slab in compression; upper hori-zontal plate in compression; vertical plate in bending and lower horizontal plate in tension, for sagging moment; reinforcing bars in tension, upper horizon-tal plate in tension; vertical plate in bending and lower horizontal plate in compression for hogging moment. The components concrete slab in compres-sion and upper horizontal plate in tension were iden-tified by means of FE models set with ABAQUS (Hibbitt et al., 2000). Figure 9 shows the FE model of a plate in tension to identify the effective width b.

The joint consisting of a steel concrete composite beam attached to a concrete-filled tubular column with circular hollow steel sections and is composed by two horizontal diaphragm plates and a vertical through-column plate (Figures 6 and 7). The presence of a vertical through column plate shown in Figure 7 allows the transmission of shear forces owing to vertical loads from a beam to the other.

150

A

IPE 400

A

A

CFT

Ø 4

57 x

12

IPE 400

Slab with electro-welded lattice girders

Slab with profiled steel sheetings

A

Section A-A

CFT

Ø 4

57 x

12

Figure 6. Beam-to-Column joint specimens.

230

60

475 300 475

625 6251250

40R22

8,5

290

= 290 x 1250 x 16A 5A

Horizontal plates

Vertical plate

300

700

= 300 x 700 x 14A 6

1250

700

A 6= 300 x 700 x 14

121,5121,5

300

369

1616

300

1950

Joint

R228,5

4,3

14 18

23°

23° 23° 23° 23°

33°

1814

4,3R22

8,5

Stiffeners

20x2

020

0

150

50

100 25

125= 125 x 200 x 12

A 4

plate

Figure 7. Steel-concrete composite beam-to-column joint. Di-mensions in mm.

In the joint design, the transition zone between the column and the beam was assumed to be over-strength compared to the beam, thus forcing plastic hinges formation in adjacent beams (UNI EN 1998-1. 2005). As a result joints were detailed by using the component method as shown in Figure 8 (UNI EN 1993-1-8. 2005).

Figure 8. Mechanical model of an interior joint

a) b) Figure 9. FE model of a plate in tension: (a) elastic stresses; (b) inelastic stresses and width.

3 EXPERIMENTAL TESTS

A total of 6 specimens was designed and fabricated according to Eurocode 4 (UNI EN 1994-1-1. 2005, UNI EN 1994-1-2. 2005) and Eurocode 8 (UNI EN 1998-1. 2005) provisions. The test specimens were interior subassemblage and the joints were made by two horizontal diaphragm plates split into two equal halves along the diagonal for fabrication conven-ience and easiness of assembling.

Once each side is 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 6. Flanges and web of each beam were connected to the horizontal plates and the vertical plate by weld-ing, as depicted in Figure 7.

In all composite specimens the connections be-tween steel beam and desk were made by Nelson 19 mm stud connectors with an ultimate tensile strength fu=450 MPa. The joints differ each other owing to the slab type and to additional Nelson 19 mm studs localized around the column in order to enforce a better force transmission between the column and the composite slabs as indicated in Figure 7. The joint specimens were subjected to monotonic and cyclic loadings up to collapse, according to the ECCS stepwise increasing amplitude loading proto-col, modified with the SAC procedure (ECCS 1986,

SAC 1997): ey=17.5 mm. Hereafter the specimens areindicated as follows:

SG

I I I I I L

LL LL

L

L L LL

LLLLL

SG

SG

SG

SGSG

SG SG

OOO

O

LVDTsINCLINOMETERS

SLAB OMEGA STRAING GAUGELVDTs

STRAING GAUGE

SG-14/ 13

SG-16/ 15

SG-16/ 14

SG-15/ 13BEAM STRAING GAUGESG-10

SG-9

SG-12

SG-11

I L

SGOL

- WJ-P1 – specimens with electro-welded lattice girders slab and no Nelson connectors around the column;

- WJ-P2 – specimens with electro-welded lattice girders slab and Nelson connectors around the column;

- WJ-PM – specimens with electro-welded lattice girders slab and no Nelson connectors around the column;

- WJ-S1 – specimens with profiled Steel Sheeting slab and no Nelson connectors around the col-umn;

- WJ-S2 – specimens with profiled Steel Sheeting slab and Nelson connectors around the column;

- WJ-SM – specimens with profiled Steel Sheeting slab and no Nelson connectors around the col-umn;

3.1 Test Set-Up The test set-up was designed in order to simulate the conditions of interior beam-to-column joints with concrete filled tubes within frame structures.

It consisted of a reaction wall, a hydraulic actua-tor (capacity 1000 kN, stroke ±250 mm), a rein-forced-concrete slab, a lateral frame designed to prevent specimen lateral displacements. Figure 10. Main instrumentation on specimens.

The main instrumentation employed is indicated in Figure 10 and detailed herein: 3.2 Results - 5 inclinometers were utilized in order to measure

the inclinations of the zone adjacent to the joint and of the beams near the connection;

Experimental results are presented on the basis of hysteretic responses. In particular, Force-displacement (F-d) curves for all tested specimens are shown in Figures 11-14, while moment vs. rota-tion (M-φ) relationships are depicted in Figures 15- 18.

- 4 LVDTs detected the interface slip between the steel beam and the concrete slab and between the inferior horizontal plates and the beam flange;

- 2 LVDTs were employed in order to measure the connection deformation; All specimens exhibit a good performance in

terms of resistance, stiffness, energy dissipation and ductility.

- 10 LVDTs were utilized in order to measure con-crete slab deformations;

- 4 Omega strain gauges detected the deformations of the concrete slab;

Both the overall force-displacement relationships and the moment-rotation relationships relevant to plastic hinges formed in the composite beams ex-hibit a hysteretic behaviour with large energy dissi-pation without evident loss of resistance and stiff-ness. In particular we can be observe that the hysteretic loops of moment-rotation relationship are unsymmetrical due to the different flexural resis-tance of the composite beam under hogging and sagging moments, respectively.

- 8 strain gauges monitored axial deformations of the reinforcing bars in order to scrutinise the ef-fective breadth of the reinforcing bars at each loading stage;

- 4 strain gauges monitored deformations of supe-rior and inferior horizontal plates;

- 4 strain gauges recorded flange strains in order to estimate internal forces in steel beams;

- 2 load cell were set on the top of pendula and were utilized in order to measure horizontal and vertical components of axial forces;

For all specimens, the experimental tests show a remarkable and progressive deterioration of strength, stiffness and energy absorption capacity as a conse-quence of the formation of a plastic hinge associated with local buckling of the beam flange. The collapse of all specimens was associated with cracking of beam flanges.

- 1 digital transducer (DT500) was employed in or-der to measure the top column displacement;

- a FieldPoint acquisition system in order to ac-quire experimental data.

Specimen WJ-P1

-1000

-600

-200

200

600

1000

-14 -10 -6 -2 2 6 10 14

Displacement [e/ey]

Forc

e [k

N]

Specimen WJ-PM

0

200

400

600

800

1000

0 6 12 18 24 30Displacement [e/ey]

Forc

e [k

N]

Figure 13. Specimen WJ-PM- Force-displacement relationship. Specimen WJ-P2

-1000

-600

-200

200

600

1000

-14 -10 -6 -2 2 6 10 14

Displacement [e/ey]

Forc

e [k

N]

Specimen WJ-SM

0

200

400

600

800

1000

0 6 12 18 24 30Displacement [e/ey]

Forc

e[kN

] Figure 11. Specimens WJ-P1 and WJ-P2 - Force-displacement

relationships. Figure 14. Specimen WJ-SM- Force-displacement relationship. Specimen WJ-S1

-1000

-600

-200

200

600

1000

-14 -10 -6 -2 2 6 10 14

Displacement [e/ey]

Forc

e [k

N]

Specimen WJ-P1

-1000

-600

-200

200

600

1000

-80 -60 -40 -20 0 20 40 60 80

φ [mrad]

M [k

Nm

]

Right Plactic Hinge

Left Plastic Hinge

Specimen WJ-S2

-1000

-600

-200

200

600

1000

-14 -10 -6 -2 2 6 10 14

Displacement [e/ey]

Forc

e [k

N]

Specimen WJ-P2

-1000

-600

-200

200

600

1000

-80 -60 -40 -20 0 20 40 60 80

φ [mrad]

M [k

Nm

]

Right Plastic Hinge

Left Plastic Hige

Figure 12. Specimens WJ-S1 and WJ-S2 - Force-displacement relationships. Figure 15. Specimens WJ-P1 and WJ-P2 - Moment-rotation re-

lationships of plastic hinges. Table 1 reports some major experimental results for each Test: in detail, the maximum applied Dis-placement (d), the maximum value of the Force (F), maximum values of both Sagging and Hogging Moment (M) and corresponding values of initial stiffness (kθ), total number of Cycles performed dur-ing the tests (Ntot) are collected.

Figure 19 shows moment vs. rotation relation-ships (M-φ) of specimens WJ-PM and WJ-SM, re-spectively. A better behaviour of the specimen WJ-PM fabricated with the electro-welded lattice girders slab is evident in term of strength owing to the better composite action in the plastc hinge section. A nu-merical finite element (FE) model of the slab includ-ing friction was developed in a previous study by Ferrario ed al. (Ferrario et al. 2007) with the objec-tive to understand the activation of the transfer mechanisms between the slab and the beam pro-posed in the Eurocode 8 (UNI EN 1998-1. 2005).

Monotonic tests were interrupted before failure of the specimens in order to reduce the risk of damag-ing the test equipment.

The analyses demonstrated that the distribution of compression forces in the slab depend on the level of

friction existing between the concrete slab and the composite column. Therefore, in order to activate some of the transfer mechanisms proposed in Euro-code 8, (i.e. Mechanism 1 and Mechanism 2), it was necessary to increase the level of friction between the concrete slab and the composite column of specimens.

Specimen WJ-S1

-1000

-600

-200

200

600

1000

-80 -60 -40 -20 0 20 40 60 80

φ [mrad]

M [k

Nm

]

Right Pastic Hinge

Left Plastic Hinge

Specimen WJ-S2

-1000

-600

-200

200

600

1000

-80 -60 -40 -20 0 20 40 60 80

φ [mrad]

M [k

Nm

]

Right Plastic HingeLeft Plastic Hinge

Figure 16. Specimens WJ-S1 and WJ-S2 - Moment-rotation re-lationships of plastic hinges.

As a result, Nelson stud connectors welded around the column were adopted in some specimens along the details indicated in Figure 7. Experimental results showed that the activation of the aforemen-tioned mechanisms was evident and effective in the specimens fabricated with Nelson connectors welded around the column; while in the specimens fabri-cated without stud connectors around the columns Mechanisms 1 and 2 were less effective.

Table 1: major experimental data Specimen d F M kθ Ntot mm kN kNm kNm/mrad

WJ-P1 210 +714.80 -686.30

+962.46 -767.72

+181.27 -141.00

21

WJ-P2 210 +708.90 -695.40

+939.07 -768.48

+183.43 -158.00

21

WJ-S1 210 +661.70 -662.00

+926.51 -693.77

+220.32 -113.50

21

WJ-S2 210 +686.83 -673.46

+977.61 -714.36

+208.25 -117.75

21

WJ-PM 490 +669.85 +874.15 -591.14

+188.95 -105.19

Mon.

WJ-SM 490 +614.08 +814.74 -528.03

+179.67 -81.79

Mon.

Figure 20 shows the total dissipated energy i.e.

the Cumulative Energy for the specimens belonging to the WJ-P and WJ-S series. The dissipated energy in the plastic cycles is practically identical for all the specimens; after the 15th cycle which corresponds to

a displacement equal to 10ey, specimens with elec-tro-welded lattice girders slab dissipate more energy than those with profiled steel sheetings owing to a better influence of the force transfer between col-umns and slabs on plastic hinges.

Specimen WJ-PM

-1000

-600

-200

200

600

1000

-160 -120 -80 -40 0 40 80 120 160

φ [mrad]

M [k

Nm

]

Right Plastic HingeLift Plastic Hinge

Figure 17. Specimen WJ-PM endowed with a slab with electro-welded lattice girders - Moment-rotation relationship.

Specimen WJ-SM

-1000

-600

-200

200

600

1000

-160 -120 -80 -40 0 40 80 120 160

φ [mrad]

M [k

Nm

]

Right Plastic HingeLift Plastic Hinge

(+M)

(-M)

Figure 18. Specimen WJ-SM endowed with a slab with pro-filed steel sheetings - Moment-rotation relationship.

Specimens WJ-PM and WJ-SM

-1000

-600

-200

200

600

1000

-160 -120 -80 -40 0 40 80 120 160

φ [mrad]

M [k

Nm

]

Specimen WJ-PMSpecimen WJ-SM

Figure 19. Comparison of (M-φ) relationships of Specimens WJ-PM and Specimen WJ-SM.

Cumulate Energy

0

100000

200000

300000

400000

500000

600000

700000

0 5 10 15 20N. of Cycles

E [k

J]

WJ-P1 WJ-S2WJ-P2 WJ-S1

Figure 20. Dissipated energy of Specimens WJ-P and WJ-S

Moreover, four fire tests on full-scale substruc-

tures were performed too. In particular, two speci-mens, labelled as T21 and T24, were pre-damaged to simulated damage owing to Type 1 spectrum com-patible accelerograms at 0.4g pga (UNI EN Euro-code 8-1, 2005), and two specimens (T22 and T25)

were not, to clearly appreciate seismic damage ef-fects on fire resistance. The Temperature vs. time curve imposed to the specimens T21-T22 and T24-T25 is shown in Figure 21(a) and (b), respectively.

Moreover, thermal analyses with different fire scenarios were carried out by means of the SAFIR code (Franssen 2000) on a two-dimensional frame model.

Figure 22 shows the considerable reduction of the design capacity moment of the joint as a function of the time of fire exposure.

T21 vs. T22

-900

-450

0

450

900

15 30 45 60

Time [min]

Def

lect

ion

[mm

]

-75

-50

-25

0

25

50

75

T21 TemperatureT22 TemperatureT21 West Deflec.T22 West Deflec.T21 East Deflec.T22 East Deflec.

Tem

pera

ture

[°C

]

Hogging Moment

-1400

-1200

-1000

-800

-600

-400

-200

0-3-2.6-2.2-1.8-1.4-1-0.6-0.2

Rotatio [rad]

M [k

Nm

]

t=20°C t=15 min t=30 min

a)

T24 vs. T25

-900

-450

0

450

900

15 30 45 60 75 90

Time [min]

Def

lect

ion

[mm

]

-75

-50

-25

0

25

50

75

T24 TemperatureT25 TemperatureT24 West Deflec.T25 West Deflec.T24 East Deflec.T25 East Deflec.

Tem

pera

ture

[°C

] Sagging Moment

0

200

400

600

800

1000

1200

1400

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

Rotation [rad]

M [k

Nm

]

t=20°C t=15 min t=30 min

b) Figure 22. Moment-rotation relationships of the joint as a func-tion of the time of fire exposure.

Figure 21. Performance of damaged (T21/T24) and undamaged (T22/T25) specimens.

In detail, the hogging capacity moment becomes approximately the 20% of the initial value after 15 minutes of exposure, while it approaches about 5% of the initial value after 30 minutes.

Specimens T21 and T22 with profiled steel sheet-ing slabs exhibited failure owing to an excessive rate of deflection at approximately 40 minutes. The test on specimen T21 terminated after approximately 34 minutes owing to runaway deflection. Following the fire test, the profiled steel sheeting separated from the slab; then the slab cracked both along the surface and through the depth with extensive buckling at one hour both of the lower flange and the web of the ad-jacent east beam.

Nevertheless, this time is enough to quit the building after a severe earthquake.

The joint endowed with prefabricated slab exhib-its a lower concrete temperature compared to the joint endowed with the steel sheeting owing to the fact that the steel portion of the column heats the concrete slab near the joint but the slab has a uni-form thickness around the column (Fig. 23).

T24 and T25 specimens endowed with prefabri-cated slabs endured one hour of fire; however, in both cases specimens were very close to failure as indicated in Figure 21b, by an increasing rate of de-flections towards the end of the test. However, at this stage, there was no permanent deformation and no sign of any significant damage from fire tests.

a) b)

4 NUMERICAL SIMULATIONS Figure 23. Temperature distribution of joints modelled by the FE Abaqus code: a) Joint with prefabricated slabs; b) Joint with profiled steel sheetings. Different fire scenarios acting in the reference build-

ing were studied with the objective to evaluate the performance of composite beams, composite col-umns and beam-to-column joints under fire load for different times of exposure.

Thermal analysis conducted by means of the Abaqus 6.4.1 code, for different times of fire expo-sure, demonstrated that the joint endowed with pre-fabricated slab exhibits a better fire behaviour com-pared to the joint endowed with steel sheeting as illustrated in Figure 24. Anew, this is due to the uni-form thickness of the lattice girder with respect to the profiled steel sheeting.

The Abaqus code (Hibbitt et al. 2000) was em-ployed to perform thermal analyses of joints for dif-ferent times of fire exposure, i.e. 15 min, 30 min and 60 min, respectively.

In sum we can affirm that both experimental and numerical results demonstrated the adequacy of the proposed joint design for concrete filled tubes to adequately face earthquake loading followed by fire.

6 ACKNOWLEDGMENTS Figure 24. Temperature distribution in the composite beam with prefabricated slab and with steel sheeting slab.

The writers are grateful for the financial support of UE under the project ECSC (RFS-CR-03034). Nonetheless, conclusions of this paper are those of the author and not of the sponsor agency.

The temperature of concrete inside the tube in-creases more quickly close to the joint in compari-son with the portion of column distant from the joint owing to the influence of the vertical through-column web plate (Fig. 25).

REFERENCES

Colombo A., Bursi O.S., 2006, Editors, “Mid-Term Report of Prefabricated Composite Beam-to-Concrete Filled Tube or Partially Reinforced-Concrete-Encased Column Connec-tions for Severe Seismic and Fire Loadings, PRECIOUS Project, University of Trento;

ECCS 1986. Recommended Testing Procedure for Assessing the Behaviour of Structural Steel Elements under Cyclic Loads. ECCS Publication n° 45;

Ferrario F., Pucinotti R., Bursi O. S., Zandonini R. 2007, Steel-Concrete Composite Full Strength Joints with Concrete Filled Tubes. Part II: Experimental and Numerical Results, Proceeding of 21st Conference of Steel Structures C.T.A., Catania, 1-3 October, pp. 397-404, Dario Flaccovio Editor;

Figure 25. Temperature distribution in the composite concrete filled tube.

Franssen J.-M. 2000. “SAFIR; Non linear software for fire de-sign”. Univ. of Liege; 5 CONCLUSIONS

Hibbitt, Karlsson and Sorensen 2000. “ABAQUS User’s manuals”. 1080 Main Street, Pawtucket, R.I. 02860. In this paper, a multi-objective advanced design

methodology was proposed instead of a traditional single-objective design. Experimental cyclic tests and numerical simulations regarding the fire be-havour of joints were carried out. Experimental re-sults show how joint details influence beam-column subassemblage responses both in terms of seismic and fire performance.

SAC 1997, Protocol for Fabrication, Inspection, Testing, and Documentation of Beam-Column Connection Tests and Other Experimental Specimens, report n. SAC /BD-97/02;

UNI EN1991-1-2. 2004. “Eurocode 1: Actions on Structures. Part 1-2 : General Actions – Actions on structures exposed to fire”;

UNI EN 1992-1-2. 2005. “Eurocode 2: Design of concrete structures. Part 1-2: General rules - Structural fire design;

UNI EN 1993-1-1. 2005. “Eurocode 3: Design of steel struc-tures. Part 1: General rules and rules for buildings”;

In detail, test results exhibited rigid behaviour for the designed composite joints; as expected plastic hinges developed in beams owing to the capacity de-sign.

UNI EN 1993-1-8. 2005. “Eurocode 3: Design of steel struc-tures - Part 1-8: Design of joints”;

UNI EN 1993-1-2. 2005. “Eurocode 3: Design of steel struc-tures. Part 1-2: General rules - Structural fire design”; It was evident that owing to local buckling ef-

fects, severe strength degradations appeared to ex-ceed 20 per cent of maximum strength values with rotations of about 40 mrad. As a result, these joints are not suitable for high ductile structures to be used under seismic loading according to Eurocode 8 (UNI EN 1998-1. 2005). Nonetheless, the corresponding minimum plastic limit of 35 mrad remains rather high for typical European earthquakes.

UNI EN 1994-1-1. 2005. “Eurocode 4: Design of composite steel and concrete struc-tures - Part 1.1: General rules and rules for buildings”;

UNI EN 1994-1-2. 2005. “Eurocode 4: Design of composite steel and concrete structures - Part 1.2: General rules – Structural fire design”;

UNI EN 1998-1. 2005. “Eurocode 8: Design of structures for earthquake resistance - Part 1: General rules, seismic ac-tions and rules for buildings”.

Thermal analyses showed that steel elements, di-rectly exposed to fire loading, had the same tempera-ture of air, while for the concrete slab different be-haviuors were observed depending on the adopted typology. In fact, specimens with profiled steel sheetings slab were characterized by temperatures higher than those endowed with electro-welded lat-tice girders. Moreover fire experiments showed the favourable behaviour of joints.