Hysteretic Behaviour of Dissipative Welded Devices for Earthquake Resistant Steel Frames

Transcription

1 Hysteretic Behaviour of Dissipative Welded Devices for Earthquake Resistant Steel Frames Miguel Espinha Department of Civil Engineering of Instituto Superior Técnico, April 211 Abstract: Earthquakes may lead to damage in large extend, which may induce high costs in repair work. In order to achieve more efficient earthquake resistant structures, an innovative fuse device for dissipative beam-to-column connections in moment resistant steel frames was developed. The seismic performance of the device was assessed based on the results provided by ten cyclic and two monotonic tests conducted on a single specimen of a beam-to-column connection with concrete slab, assembled through devices provided with different geometric parameters. Tests showed that the devices were able of concentrating plasticity and dissipating large amounts of energy through non-linear behaviour. Numerical models were developed with ABAQUS and simplified analytical models were proposed. Keywords: welded fuse device, cyclic experimental tests, hysteretic behaviour, numerical modelling 1 Introduction Previous seismic events from Northridge (USA, 1994) and Kobe (Japan, 1995) revealed a widespread failure of a large number of beamto-column connections in moment resisting steel frames (Engelhardt & Sabol, 1997). With the appearance of seismic-resistant technologies, the society in developed countries has become more demanding in terms of the seismic performance of modern structures. In order to deal with these more strict requirements, that go beyond safety, recent advances have been pointing towards achieving high performance standards in terms of economy and sustainability of modern construction. In seismic areas, this aspect results in bigger concerns with the repairability of structures damaged by strong earthquakes. In order to move the plastic hinge and the associated large plastic deformations into the beams and away from the welded areas, avoiding in this way the reduced energy dissipation and the lack of ductility due to local mechanisms and to the potential fragility of the welds, respectively, new design methodologies have been implemented throughout the last years. Among the preferred solutions are the ones that introduce weakened areas near the beam ends, being able to concentrate plasticity and perform large deformations and, thus, force the position of the plastic hinge to occur within certain limits. One of the firsts solutions of this type was the Reduced Beam Section (RBS), introduced by Plumier (199). This solution, despite being able of dissipating energy and concentrating plasticity successfully (Pachoumis et al., 21), enhanced difficulties concerning the repair work after an earthquake. In this way, new solutions point towards devices that can simultaneously dissipate energy through non-linear behaviour of its parts and also be easily repaired. Significant developments have been achieved on the research of dissipate fuses for both eccentrically braced frames (EBF) and concentrically braced frames (CBF). Chan et al. (29), Li & Li (27) and Bruneau et al. (21) conducted experimental tests on EBF with different link geometries. Plumier et al. (24) developed innovative dissipative systems for CBF, comprising pinned and U type connections. Concerning moment resisting frames (MRF), Koetaka et al. (25) proposed alternative weldfree moment connections between composite steelconcrete beams and the column s weak axis. The connection is made through replaceable steel plates bolted to the upper and lower flanges of the beam. Oh et al. (29) developed an innovative fuse based on a damper system assembled at the beam s bottom flange at the beam-to-column connection. Such a reality motivated the FUSEIS project research, aiming at the development of innovative fuse devices for seismic resistant steel frames, with the following functional objectives: easiness in design and fabrication of the devices; high seismic performance; economy in repair work. Facing these objectives, previous numerical investigations were undertaken by Mazza & Pedrazzoli (28), which concluded on the optimal fuse configurations, resulting in a bolted and a welded proposal. 1

2 2 Experimental program 2.1 Fuse specimens The present paper concerns only on the investigations taken on the welded proposal. The fuse is made by plates welded to the beam s webs and bottom flange, as is shown schematically in Figure 1. where M fuse is the maximum moment developed by the fuse device and M pl,beam is the plastic resistant moment of the non-reinforced composite crosssection of the beam. The maximum moments of the M fuse were computed with an analytical model developed for the purpose, which is introduced in Section 5. The corresponding values of the capacity ratio are presented in Table 2 for both sagging (α + ) and hogging (α ) moments. Table 2: Capacity ratio of the fuses. Figure 1: Welded fuse type There is a gap on the slab over the fuse in order to avoid major cracking of the concrete due to its low ductility when compared with steel. The rebars have however continuity, so that the transmission of stresses is assured. In order to avoid yielding of the rebars, since these are irreplaceable, they had to be computed so that the position of the plastic hinge lies near its centre of gravity, reducing this way the strains on the material. The areas shown in grey are reinforced with additional welded plates, to avoid spreading of plasticity to the adjacent irreplaceable steel parts of the beam. In order to assess the performance of the fuse device, a total of ten cyclic and two monotonic experimental tests were conducted on a single subassembly of a beam-to-column connection provided with fuses with different geometric parameters. Each test was conducted until complete failure of the flange plate of the fuse, after which the fuse plates would be replaced by new ones and another test would be performed. The web plates were designed to withstand shear forces and have the same dimensions in every test. In this way, the only dimensions that change between tests are the thickness (t f ) and width (b f ) of the flange plate, as shown in Table 1. Plate A B C D E F α +,45,57,57,47,71,54 α,27,38,39,25,48,3 The testing order is as follows: first plates D, A, B, and C were tested in this order, followed by a repetition of this set of plates, in the same order. Then plates F and E were tested without repetition and finally the sagging and hogging monotonic tests were conducted on plate C, also in this order. 2.2 Experimental set-up and loading history The experimental test set-up is shown schematically in Figure 2. Table 1: Dimensions of the fuse flange plates (in mm). Plate A B C D E F t f b f The dimensions were chosen to provide the fuses with different values of an analysis parameter α, called as capacity ratio and defined by equation 1: α = M fuse M pl,beam (1) Figure 2: Experimental set-up scheme The cyclic displacements are imposed to the specimen by the actuator according to a loading history defined in terms of the interstorey drift, as shown in Table 3. If failure has not been reached after completing the loading history, cycles with an amplitude of 4% drift are performed until complete failure of the flange plate. 2

3 Table 3: Definition of the loading history protocol. Step Interstorey drift (%) Nr. of cycles 1,15 3 2,25 3 3,5 3 4, , 3 6 1, , 3 8 2, , 3 1 3, , 3 be a consequence of the damage accumulation on the parts of the test assembly that are not replaced between tests. The failure modes of all specimens were similar, comprising the development of cracks at the midsection of the flange plate under tension, as shown in Figure 4 3 Analysis of the experimental results 3.1 Global hysteretic behaviour The analysis of the results was based mainly on the moment-rotation (M θ) diagrams of the fuse specimens. As an example, M θ diagrams for both tests on plate D are shown in Figure 3. Plate D: 2nd test Plate D: 1st test M (knm) 3 2 Figure 4: Typical failure mode. Measurements showed that both the column and the composite beam remained under the elastic range, behaving close to rigid bodies. Regarding the composite behaviour, the slippage at the interface slab-beam proved to be relatively small, with values below,2 mm for all specimens. As for the monotonic behaviour, the monotonic M θ diagrams of plate C were compared with the corresponding cyclic diagram, as shown in Figure 5. 1 M (knm) (mrad) 3 2 Cyclic Monotonic Figure 3: M θ diagram - plate D (mrad) The diagrams show that the hysteretic behaviour of the fuses is stable, characterised by a marked pinching phenomenon, due to the buckling of the fuse plates when under hogging rotations, explaining also the asymmetry of the diagram in terms of moments. Concerning the ductility, all specimens were able to perform 35 mrad rotations, which is the minimum value recommended by EN Moreover, with exception of the second tests of specimens A, C and D, all other fuses achieved rotations above 41 mrad. When comparing the diagrams between the first and second tests of the same plate specimen, it becomes clear that there is a slight degradation in terms of strength and energy dissipation, from the first test to the second. This degradation should 4 Figure 5: Comparison between cyclic and monotonic tests on plate C. Both curves are very similar in terms of the initial stiffness and yield moments. The monotonic tests seem to adjust well to the cyclic results within the same rotations, being very approximate to a cyclic envelope curve. In contrast with the cyclic tests, since in the monotonic tests there is no damage accumulation from previous cycles, these were able to withstand larger imposed rotations and, therefore, show higher ductility. The sagging monotonic test showed however a considerable higher hardening. This proves that the 3

4 flange plate could take more advantage of its hardening, increasing the maximum moment. In the cyclic tests however, the flange plate buckles between cycles, introducing a deformation which cannot be completely recovered in the subsequent sagging cycle. Consequently, the plate has always important residual negative strains at the beginning of a sagging cycle, due to the cumulative damage resulting from the cyclic buckling. The performance evaluation of the fuses was performed mainly regarding the stiffness, resistance and energy dissipation. 3.2 Stiffness Figure 6 shows the stiffness at the end of each cycle, in terms of the variation of the ξ ratio, defined by the ratio between the unload stiffness at the end of each of cycle and the initial elastic stiffness of the specimen. The tests presented in the diagram refer to the first test on each of the plates. From the observation of the diagram, it is possible to state that, in general, the curves have a similar evolution for all specimens for both sagging and hogging rotations. As expected, there is overall a stiffness loss when comparing with the initial elastic stiffness, which is evidenced by the values of ξ below 1,. When comparing sagging and hogging behaviour, it becomes clear that a more expressive stiffness degradation occurs at hogging, which could be explained by the degradation of the steel induced by the cyclic buckling. By comparing plates with low and high values of α + (e.g., plates D and E, respectively) the shape of the curves shows that specimens with higher values of α + also have a lower stiffness variation rate, especially at sagging rotations. 1,5 1,,5 Plate A Plate B Plate C Plate D Plate F Plate E ε at the end of each cycle is presented in Figure 8 for the first test of each specimen. This ratio is defined as the bending moment at the end of each cycle divided by the yield moment of the specimen in the considered direction. 2, 1,5 1, 5,5 Plate A Plate B Plate C Plate D Plate E Plate F, Cycle nr ,5 1, 1,5 2, Figure 8: Resistance ratio at the end of each cycle. The evolution of the resistance ratio seems to be very similar between specimens for sagging rotations, presenting overall a considerable hardening, reaching in some cases a values of 1,5 times the yield moment. This phenomenon occurs mainly due to the hardening of the flange plate in tension, explained by the marked hardening present in the steel s stress-strain curves. On the other hand, in the case of hogging rotations, all specimens suffered from a sudden decrease in resistance when achieving the buckling state. The sagging and hogging resistance of the fuses are expected to be directly controlled by the values of the capacity ratios α + and α, respectively. This aspect is evidenced by the diagrams in Figures 9 and 1. M (knm) My Mmax 2, Cycle nr ,5 1 1, 5 1,5,45,45,47,47,54,57,57,57,57,71 + Figure 6: Stiffness ratio at the end of each cycle. Figure 9: Sagging resistance vs. α Resistance In order to make a comparison between tests easier, the evolution of the dimensionless resistance ratio The sagging resistance diagram shows that both yield (M y ) and maximum moments (M max ) increase with α +, showing a good correlation. There are some exceptions though, where the same value of α + corresponds to different values of resistance. 4

5 (a) (b) (c) (d) Figure 7: Degradation of the upper face of the slab after: 2 tests (a), 4 tests (b), 5 tests (c) and 1 tests (d). M (knm) My Mmax Wtotal (x1 3 knm.mrad) ,25,25,27,27,3,38,38,39,39,48 2,25,25,27,27,3,38,38,39,39,48 Figure 1: Hogging resistance vs. α. Figure 11: Total dissipated energy W for different values of α. The specimens showing this are the ones that belong to the last executed tests, exhibiting therefore a decrease in resistance, which is consequence of the damage accumulation between tests. As for the hogging moments, it stands out that these present a slightly higher variation with the capacity ratio, when compared with the sagging ones. This shows that the hogging resistance of the fuse might be more sensitive to a geometry variation of the flange plate and, consequently, of α. 3.4 Energy dissipation The energy dissipation capacity plays one of the most important roles in describing the seismic performance of the fuses. The total amount of dissipated energy W total was computed against the value of α, as shown in Figure 11. The correlation shown by the chart points to the conclusion that this parameter seems to be suitable to interpret the influence of the postbuckling behaviour on the energy dissipation capacity, strengthening the fact that the severity of the buckling has a fundamental influence on the performance of the fuse. This aspect may also be recognised by observing the shapes of the momentrotation diagrams. Figure 12 shows clearly the difference between plates with extreme values of α mainly due to the pinching phenomenon, which affects directly the energy dissipation capacity. The pinching is in practice a constriction of the momentrotation diagram, contributing for a deviation from an elastic-perfectly plastic (EPP) behaviour, resulting in a lower capacity of energy dissipation. The evolution of the degradation between tests may also be interpreted through energetic considerations. To do so, the total amount of dissipated energy between different plates is compared at the 5

6 M (knm) Plate D Plate E is the range of the yield rotations. The correspondent diagram is presented in Figure 14 for the first test of each specimen (mrad) 1, Figure 12: Comparison between M θ diagram of plates D and E. 1,25 1,,75,5,25 Plate A Plate B Plate C Plate D Plate E Plate F Energy failure criteria end of the first and second tests of each plate, as shown in Figure 13., Plastic cycle W (x1 3 knm.mrad) 12 Figure 14: Dissipated energy ratio at the end of each cycle A B C D 1st test 2nd test Plate Figure 13: Energy comparison between first and second test. The chart shows that, with exception of plate D, the first tests were able to reach higher levels of energy dissipation. This aspect reflects that the degradation of the irreplaceable parts, namely the cracking of the concrete slab (Figure 7), has slight influence on the energy dissipation. The evolution of the energy dissipation along the cycles may also provide an ideia of the evolution of the damage accumulation during the test. In order to study this aspect, the dimensionless parameter η/η was computed, where η is the energy ratio at the end of each cycle and η is the same energy ratio at the end of the first plastic cycle. The energy ratio at the end of a cycle i is given by equation 2: η i = W i M y ( θ i θ y ) (2) where W i is the energy dissipated within the cycle, M y is the range of the yield moments, θ i is the range of the imposed rotations at cycle i and θ y The dashed line curve represents a simplified energy failure criteria developed by Calado & Castiglioni (1996). The diagram shows that curves from plates A and D are the ones that cross this limit earlier in the test, with reference to their first plastic cycle. These plates tend to buckle more easily, showing therefore a more marked pinching effect, when compared with the remaining plates. The monotony of the curves also shows the rate of damage accumulation. In this way, the curves with sharper variation tend to accumulate damage more rapidly than the others, starting to present a more damaged configuration earlier in the test. On the other hand, the curves with a smoother variation preserve their undamaged state longer, starting to develop severe plasticity later in their plastic cycles. Moreover, these are able to dissipate larger amounts of energy than the ones dissipated in the first plastic cycle, in relation to an EPP behaviour, which is obviated by the values of η/η above unity. The results showed overall that fuses with higher values of α provide higher performance levels, in terms of stiffness, resistance, dissipated energy and rate of degradation. Nevertheless, fuses with values of α close to unity and, therefore, with a resistance similar to the one of the composite beam, enhance more damage and thus failing to concentrate it within the fuse section, which goes against the main concept of the fuse. Therefore, the value of α should be limited by an upper bound, in order to prevent that plasticity spreads into the irreplaceable parts. 6

7 4 Numerical modelling Aiming at reproducing the experimental results, numerical finite element models were developed in ABAQUS. The models were calibrated based on the experimental results, assuming that the beam and the column had a rigid behaviour and that the composite beam had a full shear connection. Since the behaviour of the fuse is mainly dependent of the yielding and buckling of the steel plates and no major cracking was observed on the firsts tests, the concrete was modelled with an elastic behaviour, reducing considerably the computational cost. The steel stress-strain relationship was obtained based on the results provided by experimental tensile tests performed on steel specimens. The structural steel has a strength class of approximately S275 and the rebar steel of an A5 steel. Firstly, a qualitative comparison is made, by observation of the deformed shape of the sagging and hogging experimental and numerical results for one single specimen, as displayed in Figures 15 (a) and (b). Printed using Abaqus/CAE on: Sun Jan 9 23:51:12 Hora padrão de GMT 211 (a) (b) Figure 15: Comparison of experimental (a) and numerical (b) deformed shape at sagging. (a) Printed using Abaqus/CAE on: Sun Jan 9 23:39:41 Hora padrão de GMT 211 (b) Figure 16: Comparison of experimental (a) and numerical (b) deformed shape at hogging. The diagrams show the plastic deformation of the fuse in terms of the equivalent plastic strain contour plot. The plots show that the fuse was able to concentrate plasticity within the fuse plates. The numerical simulations comprised a monotonic action with displacement control and were compared with the experimental cyclic envelopes. An example is shown in Figure 17. M (knm) Figure 17: plate E Plate E experimental Plate E numerical (mrad) Experimental-numerical comparison for Overall, the models could predict reality with relative accuracy, showing a good fit to the experimental curves, specially under the elastic range. Based on the shape of both curves, it becomes clear that the elastic limit properties of the numerical simulations for both hogging and sagging rotations are in accordance with the experimental results, showing a similar development. On the other hand, in terms of the maximum moment, a higher hardening is observed in the numerical model. Since the finite element model is loaded monotonically, there is absence of degradation from previous cycles, which might have induced such resistance loss in the real tests, in particular, at sagging rotations. As far as the stiffness is concerned, the finite element model is stiffer than the real specimens. This aspect is more expressive for plates C, E and F, which are among the last ones to be tested. This could therefore be a consequence of the elastic stiffness loss felt in those specimens, due to the damage accumulation on the irreplaceable parts. A further aspect observed in the numerical results was that despite the fact that the plastic neutral axis lied close to the centre of gravity of the rebar layers, the sections did not remain plane and, therefore, Bernoulli s hypothesis was not valid, enhancing additional difficulties to the development of analytical models. 5 Development of analytical models and comparison of results In order to allow a simplified calculation of the main properties of the fuse, two analytical models were developed: a resistance and a stiffness model. 7

8 5.1 Resistance model The web plates of the fuse are fundamentally resisting to shear stresses. Bearing in mind that shear enhances brittle failure modes, bending-shear interaction may be very prejudicial to the bending resistance of the fuse, in particular to its ductile behaviour. Therefore, and regarding a plastic analysis, the bending resistance of the web plates may be neglected for simplicity and also for safety against brittle failures. This way, the resistant cross-section of the fuse is the one shown schematically in Figure 18. Rrebar,upper Rrebar,lower Rflange Figure 18: Fuse maximum moment. plastic neutral axis Mmax,fuse Table 4: Comparison between sagging moments (values in knm) Plate M + exp M + ana e (%) M + num e(%) A 25,7 259, 3,3 275,2 9,8 B 294,4 338,5 14,9 37,9 4,6 C 286,5 341,7 19,3 38, 7,5 D 247,6 257,5 4, 266,9 7,8 E 314,5 418,9 33,2 335,4 6,7 F 273,2 39,5 13,3 294,2 7,7 Table 5: Comparison between hogging moments (values in knm) Plate M exp M ana e (%) M num e (%) A -139,2-147,2 5,7-161,9 16,3 B -174,7-161,5 7,6-175,,1 C -2,5-162,1 19,2-196,1 2,2 D -124,6-147,1 18, -145,2 16,5 E -221,8-175,6 2,8-23,7 4, F -13,8-156,2 19,4-162,5 24,2 Despite the fact that cross-sections do not remain plane, the moment may be computed by doing an elastic-plastic analysis with a single curvature, for simplicity purposes. To do so, there is the need of defining the constitutive relationships σ ε of the materials. For the flange plate in compression however, there was the need of modifying the real σ ε curve to account for the buckling phenomenon. In order to do so, a buckling model had to be developed based on the plastic mechanism of the plate at buckling. The model is based on a previously developed model by Gomes (1992) and is defined by equation 3: σ b = 2 2M p AL 1 ε (3) where σ b is the stress with the buckling effect, M p is the plastic moment of the cross-section of the flange plate, A is the cross-sectional area of the flange plate, L is the free-length between the welds where the plate is free to buckle and ε is the strain. The stress σ c at compression is, for each strain, the lower absolute value between the stress at tension σ t and the buckling stress σ b. Both sagging and hogging maximum moments of the fuse were computed with the average constitutive relations of the materials obtained experimentally and were computed with the aid of a MatLab code developed for the purpose. The sagging and hogging moments are shown in Tables 4 and 5, respectively. Only results from the first test of each specimen are compared, since at this stage there is less influence of the cumulative damage. Analytical moment estimations seem to provide relatively adequate results, regarding the expected level of accuracy. The nature of the errors is related mainly to the simplified assumptions that were done, such as the assumption that plane sections remain plane and that the total deformable length is equal to the fuse s free length. Values from plates C, E and F showed less accurate results, overestimating the resistance at sagging, which could be explained again by the degradation of the test assembly, more relevant in these cases since these were the last conducted tests. As for the hogging moments, the errors are not very higher, around 2%, which is enhanced with the higher complexity of the proposed model. Overall, numerical results provided more accurate results than those from the analytical model, with errors below 1% at sagging. At hogging however, results showed higher variation, but are still acceptable. 5.2 Stiffness model The developed model aims at giving an approximate method of computing the initial stiffness of the fuses, K fuse,ini, and is based on the components method of EN The proposed spring model is shown in Figure 19. The four basic components identified are k r,sup, k r,inf, k w and k f, which correspond to the stiffness contribution of the upper rebar layer, lower rebar layer, web and flange plates, respectively. In this case, the web plates shall not be neglected, since at low rotations these are not subject to considerable 8

9 krsup kr,sup krinf kr,inf kw kf Kfuse,ini Figure 19: Spring model of fuse components. Table 6: Comparison between stiffnesses (values in knm/mrad). Plate K ana K num K exp e ana num (%) A 27,5 29,53 23,99 6,9 B 35,91 3,46 25,51 17,9 C 36,35 3,5 25,62 19,9 D 27,43 27,87 29,61 1,6 E 4,64 31,42 2,57 29,3 F 32,83 3,12 23,17 9, shear stresses. Equation 4 was used to model the axial stiffness of the elements, which was also used by Hu et al. (29) to compute the axial stiffness of bolts in end-plate bolted connections. k axial = A (4) L In previous equation A is the total area of the cross section of the component and L is the free deformable length of the component. It was seen in both the experimental and numerical simulations that the fuses may present important shear deformation due to its flexibility. To introduce this effect in the present model, there was the need of considering an additional spring in series with the previous one, using the correspondent coefficient as defined in EN : k shear =, 38A vc βz (5) where A vc is the shear effective area of the crosssection of the component, β is a transformation parameter taken as 1, and z is the lever arm of the shear forces, which was taken equal to the free length L defined in equation 4. Regarding the rebars, an area reduction of 4% was considered to match with the disposals of EN for the shear area of circular sections. The application of the proposed analytical model according to the prescriptions of EN resulted in the initial stiffnesses of the fuse K fuse,ini presented in Table 6. These are hereby compared with the initial stiffnesses provided by the numerical model and by the experimental results from the first test on each specimen for sagging rotations. One should notice that the stiffness experimental values are highly affected from the degradation of the irreplaceable parts. It is therefore not appropriate to compare the experimental and the analytical values, since the second ones are not affected by this effect and comparison effort results meaningless. The error label that appears in Table 6 is thus the relative error between the analytical and the numerical models. Regarding the analytical and numerical values, they seem to be relatively consistent. There is however in the analytical model a tendency of increasing with the capacity ratio α, as may be seen from the considerable difference obtained for plate E. In fact, it was found very difficult to obtain a method capable of simulating the fuse behaviour in terms of stiffness in a simple way, since the proposed model assumes a hypothesis which has proven to be false. This is, namely, the Bernoulli hypothesis that plane sections remain plane, which is inherent to the mechanics of the proposed components method variant, but is not verified in practice. Due to this phenomena, it was impossible to reach compatibility between the proposed components method and the method for the resistance evaluation, which should be applied separately. 6 Conclusions The fuses proved to be easy to replace, showing overall a good performance level in terms of ductility, stiffness, energy dissipation and resistance. The fuses successfully protected the majority of the irreplaceable parts, which generally remained under the elastic domain, achieved by concentrating inelastic behaviour on the fuse plates. The experimental results and observations allowed the calibration of numerical and analytical models. Overall, the numerical models could predict reality with relative accuracy, showing a good fit to the experimental curves, specially under the elastic range. Regarding the proposed analytical models, they proved to provide satisfactory results, regarding the rather simplified assumptions and the desired accuracy. There was however a deviation in some of the values, especially in the stiffness results, so that the model would have to be refined in case more accuracy is needed. 9

10 References Bruneau, M., El-Bahey, S., Fujikura, S. & Keller, D. (21). Structural fuses and concrete-filled steel shapes for seismic- and multihazard resistant design (5). New Zealand Society for Earthquake Engineering. Calado, L. & Castiglioni, C. A. (1996). Steel beam-to-column connections under low-cycle fatigue: Experimental and numerical research. In: In: Proceedings of XI WCEE. Chan, R. W. K., Albermani, F. & Williams, M. S. (29). Evaluation of yielding shear panel device for passive energy dissipation. Journal of Constructional Steel Research (65), EN (24). Eurocode 3: Design of steel structures - Part 1-8: Design of joints. Comité Européen de Normalisation (CEN), Brussels. Pachoumis, D. T., Galoussis, E. G., Kalfas, C. N. & Efthimiou, I. Z. (21). Cyclic performance of steel moment-resisting connections with reduced beam sections - experimental analysis and finite element model simulation. Engineering Structures (32), Plumier, A. (199). New design for safe structures in seismic zones. IABSE Symposium on Mixed Structures Including New Materials, Brussels. Plumier, A., Doneux, C., Castiglioni, C., Brescianini, J., Crespi, A., Dell Anna, S., Lazzarotto, L., Calado, L., Ferreira, J., Feligioni, S., Bursi, O., Ferrario, F., Sommavilla, M., Vayas, I., Thanopoulos, P. & Demarco, T. (24). Two INnovations for Earthquake Resistant Design - The INERD Project, Final Report. Research Programme of the Research Fund for Coal and Steel. EN (24). Eurocode 8: Design of structures for earthquake resistance - Part 1: General rules, seismic actions and rules for buildings. Comité Européen de Normalisation (CEN), Brussels. Engelhardt, M. D. & Sabol, T. (1997). Seismic-resistant steel moment connections: Developments since the 1994 northridge earthquake. In: Progress in Structural Engineering and Materials Vol. 1 (1). John Wiley & Sons, pp Gomes, A. M. (1992). Comportamento e Reforço de Elementos de Betão Armado Sujeitos a Acções Cíclicas. PhD Dissertation, Instituto Superior Técnico, Universidade Técnica de Lisboa. Hu, Y., Davison, B., Burgess, I. & Plank, R. (29). Component modelling of flexible endplate connections in fire. Steel Structures (9), Koetaka, Y., Chusilp, P., Zhang, Z., Ando, M., Suita, K., Inoue, K. & Uno, N. (25). Mechanical property of beam-to-column moment connection with hysteretic dampers for columns weak axis. Engineering Structures (27), Li, H.-N. & Li, G. (27). Experimental study of structure with dual function metallic dampers. Engineering Structures (29), Mazza, I. & Pedrazzoli, F. (28). Numerical Modelling of Innovative Type of Moment Resistant Frames with Dissipative, Easy-Replaceable, Joints. Master s thesis, Politecnico di Milano. Oh, S.-H., Kim, Y.-J. & Ryu, H.-S. (29). Seismic performance of steel structures with slit dampers. Engineering Structures (31),