Elghazouli 2008 Engineering-Structures

Engineering Structures 30 (2008) 1802–1819www.elsevier.com/locate/engstruct

Seismic performance of composite moment-resisting frames

A.Y. Elghazouli∗, J.M. Castro, B.A. Izzuddin

Department of Civil and Environmental Engineering, Imperial College London, UK

Received 29 April 2007; received in revised form 10 September 2007; accepted 10 December 2007Available online 11 January 2008

Abstract

This paper examines the seismic performance of composite steel–concrete moment-resisting frames. A number of studies are carried out inorder to assess the influence of key parameters, related to the structural configuration as well as design and loading considerations, on the inelasticseismic behaviour. Several sensitivity and parametric investigations are undertaken using an advanced analysis program that accounts for materialand geometric nonlinearities. Particular emphasis is given to composite frames designed according to the provisions of the European seismic code,Eurocode 8. The validity of employing simplified nonlinear-static loading approaches is evaluated by comparison against the results of incrementaldynamic time-history analysis. Natural earthquake acceleration records, which are specifically selected and adjusted for compatibility with theadopted design spectrum, are utilised for this purpose. In terms of frame configuration, it is shown that the span of the composite beam, the extentof gravity loading and the number of stories can all have significant implications on the actual inelastic response characteristics of the structure.Moreover, several design assumptions and decisions, such as those related to the behaviour factors, drift considerations, moment redistributionand panel zone contribution, can also have a direct impact on the resulting performance. The studies presented in this paper highlight importantbehavioural observations and trends, some of which point towards the need for further consideration and refinement of current design procedures.c© 2007 Elsevier Ltd. All rights reserved.

Keywords: Seismic response; Moment frames; Composite steel/concrete; Nonlinear analysis; Eurocode 8

1. Introduction

The use of moment frames incorporating compositesteel–concrete floor systems offers several behavioural andpractical advantages over bare steel and other alternatives. Theincrease of stiffness and capacity due to composite actionenables the use of larger beam spans under the same loadingconditions. Accordingly, the demand for larger and moreflexible usable space, coupled with the need for faster optimisedconstruction processes, has led to an increased utilisation ofcomposite frames in recent years.

In terms of seismic performance, composite framesexhibit favourable behaviour due to the enhanced responsecharacteristics including ductility properties. However, thenumber of detailed studies carried out on the seismic responseof composite frames, in comparison with bare steel orreinforced concrete counter-parts, has been relatively limited.

∗ Corresponding author.E-mail address: [email protected] (A.Y. Elghazouli).

0141-0296/$ - see front matter c© 2007 Elsevier Ltd. All rights reserved.doi:10.1016/j.engstruct.2007.12.004

This is attributed, on the one hand, to the assumptionthat the behaviour could be largely inferred from studieson steel or reinforced concrete frames and, on the otherhand, to the complexity of some of the analysis anddesign issues involved, as reported in previous investigations(e.g. [1–3]). Earlier studies have dealt with several modellingand design considerations including connections [4,5], shearinteraction [6], slab effects [7] and behaviour factors [8,9].

Although previous studies have addressed and resolvedimportant behavioural aspects, there is a need for assessingthe key parameters influencing the seismic performance ofcomposite moment frames with typical geometric and loadingconfigurations. In particular, with the recent introduction ofEurocode 8 [10], it is important to examine the performanceof composite frames designed according to the new Europeanseismic code. Similar to other seismic codes, the design ofcomposite frames in Eurocode 8 (EC8) largely stems fromprovisions for bare steel frames with some modifications.Consequently, it is also necessary to explore specific features ofcomposite frames and their implication on the inelastic seismicbehaviour.

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A.Y. Elghazouli et al. / Engineering Structures 30 (2008) 1802–1819 1803

Fig. 1. Structural arrangement for the reference structure: (a) Plan configuration, (b) Elevation of moment-resisting frame.

In this paper, the seismic performance of composite momentframes is examined. Although emphasis is given to framesdesigned according to the provisions of EC8, most of thediscussions are of more general nature irrespective of the codeemployed. After introducing the structural configuration as wellas design procedures and assumptions adopted, the numericalmodels and response criteria utilised in this investigationare described. This is followed by an assessment of thenonlinear-static response of a typical composite moment framewith due account of the sensitivity of the response to theloading pattern and the presence of gravity loading. Thevalidity of using nonlinear-static approaches for this type offrame is evaluated by comparison with dynamic time-historyanalysis, using earthquake records which are carefully selectedand adjusted for compatibility with the design spectrum.A number of numerical studies are then carried out inorder to assess the influence of key loading, geometric anddesign parameters on the response of composite frames. Themain behavioural observations and trends are discussed andtheir implications on seismic performance and design arehighlighted.

2. Structural configuration and design procedures

For the studies described in this paper, a reference structureis selected such that it can be used as a basis for parametricvariations. The structure consists of a five-storey compositesteel–concrete office building as shown in Fig. 1. In plan,columns are spaced at 9 m in both directions. Unlike ‘tube’systems, in which perimeter moment frames may exhibitrelatively short spans, the choice of an internal composite framearrangement would typically be associated with comparativelylarge beam spans. With due account of column inertia, lateralresistance is provided by composite moment frames in onedirection whilst the other plan direction is assumed to have abraced system. The floors consist of 120 mm thick compositeslabs of the re-entrant (dovetail) profile with 1 mm thick steeldeck supported by secondary beams spaced at 3 m.

The internal composite steel/concrete moment-resistingframe, depicted in Fig. 1(b), was adopted in this study as thereference frame (CREF). The frame was designed according tothe provisions of Eurocodes 3, 4 and 8 [10–12]. European steelprofiles were used for the columns (HE) and the beams (IPE).

1804 A.Y. Elghazouli et al. / Engineering Structures 30 (2008) 1802–1819

Full shear interaction was assumed in the composite beams toavoid undesirable behaviour in connectors as indicated in recentstudies [13]. The structure was initially designed according toEurocode 4 for the gravity and wind loading scenarios. Theload values were obtained from Eurocode 1 [14] and comprisethe self-weight incorporating an allowance of 2 kN/m2 forfinishing/partitions, in addition to an imposed load of 3 kN/m2,with appropriate load factors. To satisfy the gravity designsituation, IPE500, HEB300 and HEB450 cross-sections wererequired for the beams, external columns and internal columns,respectively.

The initial frame resulting from the vertical loading scenariowas then checked for seismic conditions according to EC8.The design peak ground acceleration (PGA) was assumed as0.30g, whilst a rock soil (Ground Type A) and Spectrum Type1 were considered. Since the structure satisfies EC8 regularityconditions in plan and elevations, equivalent lateral seismicloading based on an idealised first mode of response wasadopted. The fundamental period of vibration (T1) based on thesimplified expression given by EC8 (T1 = 0.085H3/4, whereH is the overall height in metres) was 0.65 s. However, T1was found to be nearly 1.0 s from modal eigenvalue analysisperformed in a simplified structural model assuming centrelinerepresentation for the members. This value was deemed moreaccurate and hence was used in the evaluation of the design baseshear. On the basis that use is made of Class 1 cross-sections,which satisfy DCH (Ductility Class High) requirements, thereference behaviour factor (q) in this case is 5αu/α1. Therecommended code value for αu/α1 for multi-bay multi-storeyframes is 1.3 which results in a behaviour factor of 6.5. Thetotal mass considered for the seismic design situation wasabout 1440 tons based on a combination of the unfactoreddead load and 30% of the imposed load. The resulting designbase shear was about 700 kN which was distributed linearlyover the height. It should be noted that wind loading was alsoconsidered, based on the assumption of a wind pressure of0.65 kN/m2, but this had no influence on the resulting membersizes.

In addition to satisfaction of the seismic strength demandsin members, other seismic design checks include compliancewith stability and drift criteria as well as capacity designconsiderations [10,15]. Second-order stability effects areconsidered through the sensitivity coefficient (θ) given as:

θ =Ptotdr

Vtoth, (1)

where Ptot and Vtot are the total cumulative gravity load andseismic shear applied at the storey under consideration; h isthe inter-storey height; and dr is the design inter-storey drift(product of elastic inter-storey drift and ‘q’). For θ ≤ 0.1,second-order effects may be ignored. If 0.1 < θ ≤ 0.2, themultiplier 1/(1 − θ) may be used to account for this effect and,in any case, the value of θ should not exceed 0.3.

The other drift check is related to serviceability deforma-tions, by limiting (dr ) in proportion to the storey height (h)such that:

drν ≤ ψh (2)

where ψ depends on non-structural elements, given as0.5%, 0.75% and 1.0% for brittle, ductile or non-interferingcomponents, respectively; and ν is a reduction factor in therange of 0.4–0.5 to represent smaller serviceability events.

Apart from a number of checks to ensure ductile responseof beams, the main capacity design criterion is related to thegeneral concept of weak-beam/strong-column behaviour. Forsteel and composite moment frames, a specific applicationrule is stipulated whereby the design moments (MEd) for thecolumns can be obtained from:

MEd = MEd,G + 1.1γovΩ MEd,E (3)

where MEd,G and MEd,E are the moments due to the gravityloads and lateral seismic forces, respectively; γov is thematerial overstrength factor typically assumed as 1.25; Ω is abeam overstrength factor determined as a minimum of Ωi =

Mpl,Rd,i/MEd,i of beams, where MEd,i is the design momentin beam ‘i’ and Mpl,Rd,i is the corresponding plastic moment.

For the reference composite frame under consideration(CREF), the member sizes selected on the basis of the gravitycombination were found to satisfy all the seismic design checksas stipulated in EC8. However, the external columns had to beincreased to HEB 340 in order to limit (θ) to a value of 0.2 toavoid nonlinear analysis at the design stage. Six reinforcementbars of 20 mm diameter were also provided in the slab atthe beam-to-column joints, as recommended by the detailingrules given in Annex C of EC8. The steel part of the beam-to-column connections were considered to be fully welded inorder to achieve an idealised rigid response. Additionally, tocomply with EC8 provisions, the column panel zones had tobe strengthened through the addition of sets of two doublerplates with thickness 21 mm and 29 mm to external and internalcolumns respectively.

3. Analytical approaches

3.1. Numerical modelling

The nonlinear finite element program ADAPTIC [16] whichaccounts for material and geometric nonlinearities is used forthe analysis of the composite frames. The accuracy of theprogram in modelling the response of steel and compositestructures has been extensively validated in previous studies(e.g. [17–22]).

For the steel columns, four cubic elasto-plastic elements(cbp2) [17] are utilised. The cbp2 elements account for thespread of plasticity across the section, using a fibre approach, aswell as along the length using Gaussian sections in conjunctionwith a cubic shape function. The same elements are used forthe composite beams where two parallel lines of 18 cbp2elements are considered at the centroids of the steel andconcrete constituent parts, as illustrated in Fig. 2.

Composite action is developed by introducing link elements(lnk2) with rigid properties connecting the steel and concreteline elements. A rigid connection is considered on theassumption that a sufficiently high degree of shear connectionis present such that the influence of slip between the steel


Fig. 2. Modelling of composite beams.

beam and the slab on the response is not significant. It shouldbe noted however that the degree of shear connection mayneed to be notably higher than that implied by a theoreticalfull-connection (i.e. 100% shear connection) in order toachieve these idealised conditions. In a recent study [23] inwhich composite beams were modelled with nonlinear springsrepresenting a range between 50% and 150% shear connection,it was observed that only with 150% or more can the inelasticproperties be reasonably represented by a rigid connectionidealisation. Analytical simulations of experimental studiescarried out on a full-scale composite frame [24] designedto Eurocode 8 also indicated a discrepancy of less than 3%on the response when the composite beams were modelledusing the actual shear connection or with a rigid idealisation.Whilst low and moderate levels of shear connection are notrecommended for seismic-resisting composite systems [13],due to the susceptibility of the connectors to low-cycle fatigue,it is important to emphasise that the slip between the slab andthe steel beam may need to be incorporated in numerical modelsunless a very high degree of shear connection is adopted indesign.

For simplicity, the effective slab width was assumed as1.25 m based on an estimate of the contra-flexure location andthe provisions of EC4 for the sagging moment region. Althoughthe provisions vary in EC4 and EC8 for sagging/hoggingregions and for design/analysis, a sensitivity study indicatedthat these differences do not have a notable influence on theoverall frame response [23]. Further discussion dealing withthe issue of effective width in composite beams can be foundelsewhere [25].

Special attention is given to the modelling of the panel zoneregions, for which a modified frame approach [22] is adoptedby assembling link (lnk2) and joint (jel2) elements. Fig. 3shows the model for a typical external connection. Distinctionis made between the main panel zone and the top part whichis in contact with the slab. The contact behaviour betweenthe composite slab and the column is also taken into accountthrough the introduction of joint elements. Asymmetric curvesare employed for panel zones located at external joints, whilstsymmetric relationships are considered for internal panels. Forthe top panel, a linear elastic response is considered. On theother hand, a plastic curve is considered for the slab-contactjoint. Under negative bending, the capacity at the contact regionis based on the maximum yield force that can develop in thereinforcement bars whilst, under positive moment, the jointcapacity is based on the maximum axial force that can bemobilised in the steel profile of the composite beam. Theinclusion of a joint element to model slab interaction also

Fig. 3. Numerical models for composite joints [22].

permits the consideration of partial-strength connection effectsif necessary.

For steel materials, a bilinear elasto-plastic cyclic modelwith strain-hardening (stl1) is adopted for both structural andreinforcing steel (Fig. 4(a)). On the other hand, concretenonlinear behaviour is accounted for through a uniaxial cyclicconstitutive model (con1) featuring both compressive andtensile softening (Fig. 4(b)). In all the analyses performedin this study, the steel elastic modulus (E) and the strain-hardening coefficient (µ) are assumed as 210 × 103 N/mm2

and 0.5%. The yield strength values ( fy) for structural steeland reinforcement bars are assumed as 275 and 500 N/mm2,respectively. For concrete, the values of Ec1, Ec2, fc1, andfc2 are assumed as 15 × 103, 8 × 103, 30 and 18 N/mm2,respectively. It is also assumed that the residual strength incompression is reached at a concrete strain of 0.35%, whilstconcrete behaviour in tension is ignored.

For eigenvalue as well as time-history analysis, appropriaterepresentation of the mass and damping characteristics of thestructure is required. In the case of eigenvalue analysis, onlyconcentrated mass elements (cnm2) are incorporated in themodel and located at the thirds of the beam spans. For dynamicanalysis, viscous damping, assumed as 5% of the critical, isalso considered by inclusion of dashpot elements (cnd2) atthe same locations of mass elements. It is worth noting atthis point that discrete modelling of the panel zone has aninfluence on the stiffness, and in turn on the natural periodsof vibration. As noted before, for the reference frame, thefundamental period was found to be around 1.0 s, using aconventional centreline modelling approach. When the panelzone is realistically incorporated, with due account of the


Fig. 4. Constitutive models adopted for: (a) Steel and (b) Concrete.

Fig. 5. M–Φ curves for composite beam: (a) Sagging and (b) Hogging moment.

additional stiffness of the doubler plates, the period reduces to0.88 s. On the other hand, if all the panel zone componentswithin the discrete model are assumed to be infinitely rigid, theperiod reduces further to about 0.8 s.

3.2. Response criteria

In order to assess the inelastic seismic performance ofcomposite frames, it is necessary to define specific criteriathat can be associated with the yield and ultimate responselevels. Various approaches and assumptions can be adopted toselect comparative measures for this purpose. The notion of‘significant yield’, which is typically associated with plastichinge formation in framed structures, is employed in thisinvestigation. Due to the more detailed strategy adopted in therepresentation of the composite members, it was decided toestablish the local yield criteria based on fundamental cross-section data that ADAPTIC is able to provide. Instead ofrelating the plastic hinge formation to the achievement ofa theoretical plastic moment, plastic hinges are assumed toform in the composite beam under negative moment whenboth extreme fibres of the steel profile reach the yield strain.On the other hand, a plastic hinge is assumed to form underpositive bending when the axial capacity of the steel beam is

reached; this criterion is valid for ductile sections in whichthe plastic neutral axis is located in the slab. The two criteriadescribed above are highlighted in Fig. 5 which depicts themoment-curvature relationship for the composite beam undersagging (positive) and hogging (negative) bending. The twocurves clearly indicate that the selected criteria are closelyrelated to the plastic moment (Mpl) of the composite cross-section, and hence can be used in identifying the formationof a plastic hinge. For the steel columns, similar to beams innegative bending, a plastic hinge is assumed to form when bothextreme cross-section fibres attain the yield strain as proposedin previous studies [26].

In the case of ultimate response, global criteria related tointer-storey drift are adopted, as in comparison with localfailure considerations (such as local buckling of lower flangeor concrete crushing in compression), they are found to governthe behaviour of the frames dealt with in this investigation.The study frames are designed and detailed such that all steelmembers satisfy cross-section requirements for the highestductility class. Additionally, the amount of the additional slabreinforcement over supports reduces the extent of possiblestrain localisation in hogging moment. On the other hand, thecomposite beams in sagging moment satisfy limitations on thedepth of the plastic neutral axis for achieving high ductility. It is


also important to note that the idealised uniaxial representationof concrete behaviour makes the definition of a strain-based failure criterion in the slab unrealistic. Moreover, thisinvestigation is concerned more with assessing the response interms of global capacity considerations and plastic mechanismsrather than examining local failure conditions. It was thereforedecided to assess ultimate conditions, where necessary, basedon a limiting inter-storey drift which, in any case, is directlyrelated to the ductility demand imposed on structural members.A limiting value in the range of 2.0%–4.0% [27,28] is normallyadopted for framed structures, with 3.0% being commonly used(e.g. [9,29]).

The numerical model described above is utilised in thefollowing section to assess the nonlinear-static as well asdynamic response of the reference composite frame presentedearlier. Subsequently, a number of parametric variations areconsidered to examine key behavioural and design issues.

4. Lateral seismic loading

4.1. Nonlinear-static response

The use of nonlinear-static, or pushover, analysis for seismicassessment and design has increased significantly in recentyears. It can be employed to assess the overall capacity andstability, and to identify the likely plastic mechanisms andassociated dissipative regions. The attractiveness of pushoveranalysis stems mainly from its relative simplicity, in terms ofmodelling and computational demands as well as interpretationof results, in comparison with nonlinear dynamic analysis [30].The main purpose of this section is to examine the responseof the reference frame using ‘conventional’ forms of pushoveranalysis, commonly referred to as the ‘triangular’ and ‘uniform’distributions. The adequacy of the pushover idealisations inrepresenting the inelastic response of the frame is then assessedin subsequent sections by comparison against the results ofincremental dynamic analysis.

The nonlinear-static response of the reference frame ispresented in Fig. 6. The analysis is conducted by increasingthe displacement at the top of the frame incrementally up to aglobal lateral drift of about 3% of the overall height. The resultsare shown for both the linear (triangular) and uniform patternsof the lateral load. Due to the regularity and simplicity of thestructure, the triangular pattern represents closely a first mode-dominated response. On the other hand, the uniform pattern isexpected to be important if higher modes contribute notably tothe response or when significant inelastic concentrations occur.

As shown in Fig. 6(a), the uniform loading pattern results inhigher stiffness and capacity in comparison with the triangulardistribution, as expected. The inter-storey drifts in the frameat yield (formation of first plastic hinge) as well as at ultimate(achievement of 3% inter-storey drift) are depicted in Fig. 6(b).The inter-storey drift distribution is almost uniform at theyield stage since the behaviour is largely elastic. At ultimate,however, lower stories exhibit significant inter-storey driftscompared to upper levels. The inelastic concentration is morepronounced for the uniform load pattern which forces a

more distinct soft-storey behaviour at the first storey. Theinelastic distributions shown in Fig. 6(b) are closely related toFig. 6(c) which illustrates the sequence and location of plastichinges occurring in the beam and column members up to theattainment of the defined ultimate limit.

Apart from the difference in response for the linear anduniform patterns, a number of important observations arenoteworthy with reference to Fig. 6. Firstly, the frame is shownin Fig. 6(a) to possess an actual capacity which is considerablyhigher than that assumed in design. A degree of overstrengthis generally expected for framed structures [15], particularlywhen a relatively high behaviour factor is employed, sincethe member sizes are likely to be governed by the gravityloading scenario or by deformation-related limits (from inter-storey or stability checks). This overstrength effect becomeseven more pronounced in the case of composite frames. Oneof the reasons for this is the relatively large spans that may bepresent in composite frames, hence increasing the dependenceof beam sizes on the gravity loading design scenario. Moreover,composite frames exhibit significant levels of redistribution,in terms of the ratio between the ultimate base shear andthat at yield (referred to as αu/α1 in EC8), which contributessignificantly to the global overstrength of the frame.

Another important observation is related to the formation ofplastic hinges predominately on one side of the beam spans,as illustrated in Fig. 6(c). This is caused by a combination ofthe influence of gravity moments (which can be significant inrelatively large spans enabled by composite action), coupledwith the difference between the plastic moment capacity of thecross-section under positive and negative moments which ischaracteristic of composite beams. Unless considerable levelsof overstrength are provided to the columns to preclude theformation of column hinges at significant inelastic drifts, beamhinges may not form in the other side of the spans at ultimate,which is the case in the reference frame. This characteristicbehaviour occurs even when the gravity loading is disregarded(as shown in Fig. 6(d)), thus indicating the significance of theasymmetry in the moment capacity of the beams.

In relation to the above discussion, it is also important tonote that whilst codes aim to achieve a ‘weak-beam/strong-column’ behaviour, a degree of column hinging is unavoidablein many cases unless large overstrength factors, well beyondthose implied in codes, are adopted for column design. This isillustrated by the hinge patterns shown in Fig. 6(c), and wasalso reported in previous studies on steel frames (e.g. [31]).In the inelastic range, the points of contra-flexure in memberschange and consequently the distribution of moments varyconsiderably from the idealised conditions assumed in design.Therefore, in terms of actual response, the purpose of meetingcode requirements can perhaps be viewed as aiming to achieverelatively strong columns such that beam rather than columnyielding predominates in several stories. Further discussion ofthese capacity design issues is presented later on in this paper.

In order to examine the general validity of nonlinear-static procedures for assessing the overall seismic response ofcomposite moment frames in further parametric and sensitivitystudies, the pushover results of the reference frame are


Fig. 6. Reference structure: (a) Pushover curve, (b) Inter-storey drifts, (c) Hinge formation for different load patterns and (d) Hinge formation for case withoutgravity loads.

compared with those from dynamic analysis in subsequentsections. To achieve a realistic basis for this comparison, carefulattention is given to the selection and scaling of the time-historyrecords in relation to the code-spectrum assumed in design.

4.2. Input for time-history analysis

A number of difficulties arise when choosing specificrecords for dynamic analysis, particularly in cases whereseismic hazard studies are not available. In general, two

different criteria for record selection can be employed. Oneconsists of choosing the records according to strong-motionparameters and the other is based on geophysical criteria.The search based on strong-motion parameters consists offinding records that have similar shape to the responsespectrum provided by the code, obtained for a site withsimilar characteristics. An efficient procedure may consist,for example, of calculating the average root-mean-squaredeviation (Drms) of the spectrum of the tentative record


Table 1Selected records and scaling factors

EQ ID Earthquake Date Country Magnitude (Mw) Distance (km) Soil type PGA (g) Scaling factor

CLUC Campano Lucano 23/11/1980 Italy 6.9 10 Rock 0.0599 6.0MHILL Morgan Hill 24/04/1984 US 6.5 28 Stiff 0.0340 10.0LDRS Landers 28/06/1992 US 7.3 30 Stiff 0.1351 2.3NRDG Northridge 17/01/1994 US 6.7 31 Stiff 0.0732 5.0CHCH Chi-Chi 20/09/1999 Taiwan 7.6 39 Stiff 0.4145 1.1TMRN Taumaranui 05/01/1973 New Zealand 6.6 65 Rock 0.0358 10.0HMINE Hector Mine 16/10/1999 US 7.1 61 Rock 0.0568 5.3

from the target design spectrum. Alternatively, spectrum-compatible artificial records may be generated from whitenoise. This latter technique however tends to create unrealisticrecords both in terms of frequency, phase content, number ofcycles and duration of motion [32]. On the other hand, forsituations where the site is well characterised in seismologicalterms by either a deterministic or probabilistic seismic hazardassessment, then one or more earthquake scenarios can beestablished. Accordingly, the fault mechanisms and the rangesof magnitudes and fault distances of the earthquakes affectingthe site are known [33]. In this case, the selection can beconducted by searching a strong-motion database for recordsthat match those geophysical parameters.

The selection of records is carried out herein by combiningthe two criteria discussed above. The records are selectedsuch that they correspond to the shape of the code-spectrumemployed (Type 1) as well satisfy the seismological criteriainferred by this choice. Seven records, with the lowest Drmsfrom the target code-spectrum, were sought for this study fromthe strong-motion database available at Imperial College, byimposing the following conditions: (i) moment magnitudes(Mw) larger than 6 since Spectrum Type 1 corresponds tohigh magnitude events [34]; (ii) records involving near-fault orforward directivity effects are avoided; (iii) rock or stiff soilsites, for consistency with the soil type assumed in design; (iii)PGA larger than 0.03g, to avoid applying unrealistically highscaling factors. The duration was not specifically consideredsince the model structures do not exhibit significant degradationeffects. Records with forward directivity effects were notconsidered although it should be noted that they can pose a highdamage potential to flexible structures [35]. However, thesetypes of records have been rarely observed in Europe with theonly relevant case recorded during the 1995 Aegion earthquakein Greece [36]. The seven records selected are listed in Table 1together with the scaling factors adopted in order to have agood fit with the code-spectrum for the period range of interest,which in this case is considered to be between 0.8 and 2.4 s.The response spectra, for 5% damping, for all seven records aswell as the target design spectrum are shown in Fig. 7(a).

The selection was then followed by another adjustmentprocess for matching the records to the target design spectrum,based on the introduction of wavelets to the accelerationtime series [37]. For this purpose, a modified version [38] ofthe program RSPMATCH [39] was employed. The modifiedprogram incorporates new wavelet types that avoid the needfor baseline corrections and permit adjustment for multiple

Fig. 7. Response spectra of the selected records (all normalised to PGA) andtarget design spectrum (EC8-Type 1): (a) Initial records, (b) Wavelet-adjusted.

damping levels. The correction with RSPMATCH2005 isperformed in two steps. In the first step, each record is modifiedin order to match the target spectrum within the period rangebetween 0 and 1 s. In the second step, wavelets are introducedin the time series in order to match the target spectrum forthe whole period range, i.e. between 0 and 4 s. A maximumtolerance of 5% for the mismatch between the record spectrumand the target spectrum is defined.

The response spectra for the final adjusted records as wellas the target spectrum are shown in Fig. 7(b). On the other


(a) CLUC (before adjustment). (b) CLUC (after adjustment).

(c) MHILL (after adjustment). (d) LDRS (after adjustment).

(e) NRDG (after adjustment). (f) CHCH (after adjustment).

(g) TMRN (after adjustment). (h) HMINE (after adjustment).

Fig. 8. Earthquake records adopted in the analysis.

hand, the acceleration time-histories for the final adjustedrecords are shown in Fig. 8. It is worth mentioning that theground motion characteristics, as well as the acceleration andvelocity time series, are not significantly affected by the waveletadjustment. For example, this is illustrated in Fig. 8(a) and (b)which depict the acceleration record of CLUC before and afteradjustment, respectively. However, some differences can occurin the displacement series in some cases, where a reduction inthe number of ground displacement cycles is observed. This

situation develops in cases where the wavelet modificationalgorithm requires several iterations to adjust the record.

4.3. Dynamic response

Using the seven modified records described in the previoussection, incremental dynamic analysis [40] was performed forthe reference frame. Each of the seven records was appliedwith increasing ground motion intensity, resulting in over 100


Fig. 9. Time-history results vs pushover curves.

analyses. In this study, the intensity measure was consideredas the spectral acceleration at the fundamental period and, foreach analysis, the base shears as well as global and inter-storeydrifts were examined. A large amount of data was obtainedfrom the dynamic analysis hence, for brevity, only results whichare considered to be of direct relevance to the issues addressedin this paper are discussed below.

The overall dynamic response is compared in Fig. 9 to thatobtained from pushover analysis, with both linear and uniformpatterns. The response is presented in terms of base shear versusmaximum drift at the top of the frame. A special procedure isrequired to extract these results from the output of the dynamicanalysis since the maximum displacement does not necessarilycoincide with the time corresponding to the peak base shear. Assuggested in previous studies [41], the maximum displacementwas matched with the peak base shear within a specified timeinterval, for which a value of ±0.5 s was found to produceconsistent results.

Fig. 9 indicates relatively low dispersion between the resultsfrom the seven earthquakes, which is clearly aided by therecord-selection and spectrum-matching procedures describedbefore. More importantly, the plot reveals relatively goodcorrelation with the pushover results, particularly for the linearpattern case. As shown in the figure, the average of the time-history results is more closely related to the linear pattern formost of the response range, except perhaps at significant driftlevels when inelastic deformations may tend to concentrate inlower storeys. In contrast, the uniform pattern provides moreof an envelope of the global behaviour rather than a realisticrepresentation of the mean level of dynamic response.

Pushover analysis using a linear triangular pattern of loadingalso provides better correlation with time-history results interms of inter-storey drifts. For example, Fig. 10 depicts thedistribution of the maximum inter-storey drifts correspondingto a top displacement of about 0.15 m, as this correspondsapproximately to average inter-storey drifts in-between yieldand ultimate. The same procedure adopted for extracting thebase shear versus top drift results was also used in this caseto obtain the inter-storey drifts corresponding to the specifiedlevel of global deformation. The plot shows that the pushover

Fig. 10. Inter-storey drifts for 0.15 m global drift (time-history, linear anduniform pushover).

Fig. 11. IDA curves of maximum base shear for control structure.

analysis with triangular loading pattern is more closely relatedto the mean of the time-history analyses. On the other hand,the uniform pattern idealisation tends to overestimate the inter-storey drifts at lower stories and to underestimate the demandat upper levels.

The maximum base shears and top frame drift obtainedfrom the incremental dynamic analysis are depicted in Fig. 11and Fig. 12, respectively. These are shown in relation to thescaled spectral acceleration corresponding to the fundamentalperiod of vibration and 5% damping, i.e. Sa (T1, 5%).Whilst the maximum base shears do not show considerablevariation for the various records, the dispersion of globaldrifts increases with higher seismic intensity as the levelof inelastic deformations becomes more significant. Moreimportantly, as illustrated by Fig. 12, the peak top displacementis largely proportional to the seismic intensity. This suggests thegeneral validity of the ‘equal displacement approximation’ rulecommonly adopted in several seismic codes, including EC8, forthe structure under consideration. It should be noted howeverthat if the response is assessed on the basis of individual


Table 2Summary of parameters considered and section sizes adopted

Study Designation Remarks Gravity loading design Seismic design

Primarybeams

Secondarybeams

Externalcolumns

Internalcolumns

Externalcolumns

Internalcolumns

GC*

Referenceframe

CREF – IPE 500 IPE 330 HEB 300 HEB 450 HEB 340 HEB 450 c

Designconsiderations

CR4 q = 4 IPE 500 IPE 330 HEB 300 HEB 450 HEB 500 HEB 500 aCRED30 Redistribution IPE 400 IPE 330 HEB 300 HEB 450 HEB 300 HEB 450 cCPZ No doubler

platesIPE 500 IPE 330 HEB 300 HEB 450 HEB 340 HEB 450 c

CSI015 PGA = 0.15g IPE 500 IPE 330 HEB 300 HEB 450 HEB 340 HEB 450 cCSI050 PGA = 0.50g IPE 500 IPE 330 HEB 300 HEB 450 HEB 550 HEB 600 d

Spacing andgeometry

CTB3 3 m spacing IPE 300 – HEA 220 HEA 240 HEA 320 HEA 360 cCTB12 12 m spacing IPE 550 IPE 400 HEB 400 HEB 650 HEB 400 HEB 650 bCPB6 6 m span IPE 330 IPE 330 HEB 260 HEB 320 HEB 450 HEB 500 cCSTOR3 3-storey frame IPE 500 IPE 330 HEB 260 HEB 300 HEB 260 HEB 320 a

∗ Governing Criterion: a. seismic strength demand, b. gravity loading scenario, c. stability coefficient, d. serviceability drift.

Fig. 12. IDA curves of maximum top displacement for control structure.

records, rather than the mean or median, then significantdeparture from the equal-displacement approximation would beobserved. This is illustrated in Fig. 12 by noting, for example,the drift response of the structure under the TMRN record withincreasing seismic intensity.

After examining the various approaches for applyingthe lateral seismic loading, subsequent parts of this paperinvestigate the influence of other salient geometrical and designparameters on the response. As discussed above, pushoveranalysis based on the linear pattern was shown to be in goodagreement with the average results obtained from the dynamicanalysis. Accordingly, and given the relative simplicity ofperforming and interpreting pushover analysis coupled withthe comparative nature of parametric and sensitivity studies, itwas decided to limit the discussions presented in subsequentsections to results from pushover analysis with triangular loadpatterns.

5. Design parameters and considerations

The design process of a composite moment frame involvesseveral assumptions and choices that can have a direct influence

on the performance. Using the reference frame describedbefore as a basis, a number of variations to several designparameters and assumptions are considered in this sectionin order to highlight important behavioural implications. Themodifications carried out to the reference frame are summarisedin Table 2, together with the resulting member sizes andgoverning criteria. In all cases, only one design parameter orassumption is varied within each specific frame. It should benoted that, where possible, it was decided in most cases to retainthe beams sizes based on gravity design and adjust column sizesto satisfy the various design criteria.

A significant number of parameters were varied but, forbrevity, only those which are thought to provide some insightinto key aspects of the behaviour are described herein. Thissection deals with parameters related to design, for a structureof the same geometry. In particular, focus is given to thebehaviour factor, moment redistribution, panel zone design andseismic intensity. On the other hand, Section 6 is concernedwith variations to the structural configuration.

5.1. Behaviour factor

The recommended values for the behaviour factor ‘q’ inEC8 are, in principle, intended as upper limits. However, inpractice, the recommended behaviour factor is often used asa default with a view to exploiting available ductility andenergy dissipation capabilities. It is important to note, however,that lower behaviour factors may be adopted and such achoice could often be more rational [15] particularly in low-to-moderate seismicity regions.

In order to examine some of the issues involved in usinga lower behaviour factor, the design of the reference frame(CREF) is modified (CR4) in accordance with ‘q’ of 4 (DCM)instead of 6.5 (DCH). This enables the use of Class 2 ratherthan Class 1 cross-sections if necessary but, more importantly,the change in ‘q’ results in the need for larger column sizes,as indicated in Table 2, in order to satisfy the higher lateralstrength requirement. As indicated in Fig. 13(a), this leads toan increase in the lateral capacity of the frame in comparison


Fig. 13. Influence of design considerations on lateral response: (a) Pushovercurves, (b) Inter-storey drifts.

with the reference structure. Also, as depicted in Fig. 13(b),due to the increased column-to-beam strength in the modifiedframe, an improved distribution of inter-storey drift is obtained,which is directly related to enhanced hinge formation andreduced second-order effects. This is also supported by thehinge pattern shown in Fig. 15(b), which shows more extensivehinge formation in upper stories at ultimate in comparison withthe reference frame.

In general, it should be noted that the influence of the choiceof behaviour factor on the behaviour is inter-related to severaldesign parameters and criteria. A change in behaviour factorcan clearly lead to a modification of member sizes, and hencedirectly influence the stiffness and capacity as well as theoverstrength and inelastic performance of the frame. However,for a particular frame configuration, these effects are largelydependent on a number of other parameters including the levelof seismicity, and the second-order design criteria such as thestability coefficient in EC8.

5.2. Moment redistribution

Due to the relatively large spans often associated withcomposite floors, the beam sizes are normally governed by

the gravity design rather than the seismic loading combination.Consequently, it may be beneficial to make use of theredistribution ability of composite beams in the gravity designsituation to reduce the required beam size. The reductionin beam size can be significant in composite beams due tothe characteristic difference between the positive and negativemoment capacity of the cross-section.

To examine the effect of gravity moment redistribution, thedesign of the reference frame (CREF) is modified (CRED30).In this case, the gravity design is based on a 30% redistributionin accordance with the allowance in Eurocode 4 [12]. Thisresults in a reduction in beam size from IPE 500 to IPE 400.However, this is also accompanied by a reduction in stiffnessand capacity which, depending on the seismic intensity andbehaviour factor as well as other design parameters, may stillbe adequate. For the frame under consideration (CRED30),although the reduced beam size enables a reduction in columnsizes from a capacity-design viewpoint, satisfaction of thestability criterion (θ) necessitates the use of at least the samecolumn sections.

Fig. 13(a) includes the pushover response of CRED30in comparison with CREF, whilst the inter-storey driftdistributions at yield and ultimate are shown in Fig. 13(b).On the other hand, the hinge pattern at ultimate is depicted inFig. 15(c). Whereas the stiffness and capacity of CRED30 areclearly lower than the reference structure, the overall seismicresponse is more favourable due to the more dominant beamyielding as demonstrated in Figs. 13(b) and 15(c). This isfacilitated by the lower beam-to-column capacity ratio attainedin CRED30. In general, the benefits of employing gravitymoment redistribution depend on the specific case but it is anissue that merits consideration within the design process.

5.3. Panel zone design

As noted previously, the column panel zone in the referenceframe was designed according to the provisions of EC8 whichdo not imply significant yielding within this component. Tosatisfy code requirements, each web panel was strengthenedwith two doubler plates, as described before. Other seismiccodes (e.g. [42]) explicitly allow partial yielding in panel zonesin recognition of their inherent ductility and in order to alleviateexcessive demands on the plastic hinges in the beams. It wasdeemed of interest therefore to examine the response of amodified frame (CPZ) which has the same characteristics asthe reference case (CREF) but without strengthening the webpanels.

The lateral response of frame CPZ is included in Fig. 13(a),whilst the drift distributions over the height are given inFig. 13(b). On the other hand, the hinge pattern and sequencefor this case are shown in Fig. 15(d). In comparison withCREF, it is evident that the flexibility and yielding of the panelzone leads to a significant reduction in the overall stiffnessand capacity of the frame. The panel zones in CPZ havean average yield capacity of about 35% of the beam whichresults in a reduction of about 30% in the overall framecapacity. Importantly, this also results in a highly stable plastic


mechanism involving only the panel zones and the base of thecolumns, as shown in Fig. 15(d), and hence more uniform inter-storey distribution over the height (Fig. 13(b)).

Despite the stable inelastic behaviour of panel zones,the rationale behind limiting energy dissipation within thiscomponent in EC8 is based on merited concerns. The localbending of column flanges associated with inelastic sheardeformations of the panel may cause excessive local strainsin the welded connection components, which can precipitateearly failure as suggested by several researchers [43,44]. Onthe other hand, allowing yielding in panel zones can alleviateinelastic demands on the beams and lead to favourable overallinelastic frame response. Accordingly, there are also benefitsfrom allowing a degree of yielding of panel zones, as impliedin recent design documents [45]. More detailed examinationof this issue is beyond the scope of this investigation but it isclearly an area that deserves further assessment.

5.4. Seismic intensity

In Eurocode 8, the seismic intensity is represented primarilythrough the design peak ground acceleration (PGA). In orderto assess the influence of seismic intensity on the design ofthe reference frame (CREF), for which a PGA of 0.30g wasassumed, two variations are considered (CSI015 and CSI050)corresponding to PGA of 0.15g and 0.50g, respectively. Otherdesign parameters and criteria used for the reference frame areretained.

In the case of CSI015, the design process results in identicalmember sizes to those used in CREF. This is expected sincethe criterion governing the design in both cases is the stabilitycoefficient (θ) which, as evident by close observation of Eq.(1), is independent of the seismic intensity. Since the designbase shear for CSI015 is half of that in CREF, whilst themember sizes are identical, the direct implication is that theframe overstrength is doubled. Clearly, this is a consequence ofemploying a high ‘q’ factor in a case of relatively low seismicdemand, leading to undue levels of overstrength as reported inprevious studies on framed structures [15].

For CSI050, the increase in PGA changes the governingcriterion in design. Similar effects can also be observed whenthe soil classification is changed. In CSI050, the sizes of thecolumns are governed by the inter-storey serviceability driftlimit assumed as 0.75% of the storey height. The inter-storeydrift increases with PGA, unlike the stability coefficient, andbecomes more critical. If the inter-storey drift limit of 0.75% isrelaxed to 1.0%, the seismic demand on the columns becomesthe governing factor. In comparison with CREF, the largercolumn sizes lead to relatively higher stiffness and capacity(Fig. 14(a)), as well as improved distribution of drift and hingeformation over height (Fig. 14(b) and Fig. 15(f), respectively).Moreover, as expected, the level of global frame overstrength issignificantly lower than that in CREF.

6. Frame spacing and geometry

Apart from the design parameters and considerationsdiscussed in the previous section, the frame configuration can

Fig. 14. Influence of seismic intensity on lateral response: (a) Pushover curves,(b) Inter-storey drifts.

also have a direct influence on the seismic response. In thissection, a number of geometric parameters related to the frameconfiguration are considered, as summarised in Table 2. Theseinclude the frame spacing, primary beam span and the numberof stories.

6.1. Transverse spacing

Depending on the structural plan arrangement of thebuilding, the frame spacing may be considerably differentfrom that assumed in the reference structure selected. This,in turn, would have a direct effect on the level of gravityloading imposed on the primary beam which forms part of thecomposite moment frame. In order to examine the influence ofthese possible variations, two additional cases are considered,in which the frame spacing is changed from the 9.0 m assumedin the reference frame (CREF) to 3.0 m (CTB3) and 12.0 m(CTB12). In the case of CTB3, and with reference to Fig. 1(a),the slab is assumed to span directly onto the moment frames.

As shown in Table 2, for CTB12 the sizes of beamsand columns are larger than those for CREF, as these aregoverned by the gravity loading conditions. On the other


Fig. 15. Sequence of hinge formation for: (a) CREF, (b) CR4, (c) CRED30, (d) CPZ, (e) CSI015, (f) CSI050.

hand, for CTB3 smaller member sizes are required due to theconsiderably lower gravity loading present, but the columnsizes are ultimately governed by the stability criterion. Asshown in Fig. 16(a), the resulting response, in terms of stiffnessand capacity, reflects the difference in member sizes betweenthe three frames. Due to the high column-to-beam capacityratio in CTB3, it provides a favourable distribution of plasticityover the height as shown in Fig. 16(b), as well dominant beamhinging as indicated in Fig. 18(a). However, the significantincrease in column sizes from that required for gravity loading,in order to satisfy the stability criterion rather than seismicdemands, results in relatively high frame overstrength in thecase of CTB3.

6.2. Primary beam span

In order to examine the influence of the span of the beamon the seismic response of the composite frame, the beamspan is modified from 9.0 m in the reference frame (CREF)to 6.0 m (CPB6). Due to the shorter length and reduced gravityloading, the member sizes required to resist the gravity situationare smaller than in CREF as listed in Table 2. However,the column sizes are ultimately governed by the second-orderstability criterion. The adopted design procedure therefore leadsto larger column sizes in CPB6 compared to CREF.

The response of CPB6 in comparison with CREF is shownin Fig. 17(a). The reduction in beam size leads to lower stiffness


Fig. 16. Influence of frame spacing on lateral response: (a) Pushover curves,(b) Inter-storey drifts.

and capacity in CPB6. However, as expected, the highercolumn-to-beam capacity ratio results in a more favourableinelastic distribution over the height as shown in Fig. 17(b).This is also illustrated in Fig. 18(c), indicating an ‘ideal’hinge formation in the beams. It is interesting to note that,in this case, the smaller beam spans and the prevention ofcolumn hinging permit the formation of hinges at both sides ofmost members. Clearly, the lower beam span in CPB6 enablesthe use of smaller beam cross-section hence facilitating aweak-beam/strong-column design within practical size ranges.However, the observations made above in relation to inelasticresponse characteristics could obviously change if the designerdecides to increase beam sizes, rather than column sizes only,in order to satisfy the stability criterion.

6.3. Number of stories

Another variation of the reference frame (CREF) isconsidered whereby the number of storeys is reduced to three(CSTOR3). This case is selected to highlight an issue whichbecomes particularly important in the case of frames with largebeam spans and/or low number of stories [15,46]. For CSTOR3,the relatively low overall gravity loading applied to the ground

Fig. 17. Influence of beam span and number of stories on lateral response: (a)Pushover curves, (b) Inter-storey drifts.

floor columns means that the stability coefficient does notgovern the size of the column and it is instead determinedby seismic strength demands. In principle, the capacity designrules for columns should ensure that relatively strong columnsare provided such that more dominant beam hinging occurs.However, direct implementation of the application rule of Eq.(3) in isolation can lead to unsatisfactory performance.

Figs. 17(a), (b) and 18(d) depict the pushover response,inter-storey drift and plastic hinge pattern, respectively,obtained for CSTOR3. Clearly, the column sizes are inadequatein this case to prevent the formation of an undesirablecolumn mechanism. This is largely an implication of Eq.(3) which in its suggested form ignores the influence ofgravity loading in calculating the overstrength parameter Ω .As a result, Ω is underestimated considerably for cases inwhich the gravity moment in the beams (MEd,G) constitutesa significant proportion of the total moment (MEd). This isoften the case in composite configurations as these typicallyincorporate relatively large beam spans. Although this issueis not necessarily related to the number of stories, the limitsimposed on the stability coefficient lessen its effect on therequired column sizes except in low-rise frames. In fact, a more


Fig. 18. Sequence of hinge formation for (a) CTB3, (b) CTB12, (c) CPB6, (d) CSTOR3.

realistic representation of the overstrength parameter Ω wouldnecessitate a modification of the code-specified relationshipto (Mpl,Rd − MEd,G)/MEd,E . Such modification results ina significantly higher value of Ω for CSTOR3, leading tocolumn sizes which are at least as large as those in CREF,hence preventing the formation of a storey mechanism. Theincrease in column sizes would evidently improve the seismicperformance of the structure and, importantly, it would havean insignificant influence on the final cost of the structure.Alternatively, instead of the suggested modification, the coderules could be replaced by (or used in conjunction with) othercapacity design measures.

7. Concluding remarks

This paper assesses the inelastic seismic performance ofcomposite steel/concrete moment-resisting frames designedaccording to the provisions of Eurocode 8. After discussingthe design procedures and assumptions, the numerical modelsand response criteria adopted in this investigation are described.The analysis accounts for material and geometric nonlinearitiesand the models incorporate detailed representation of steel andconcrete elements as well as the panel zone components. Anumber of studies are then carried out in order to examinethe influence of several key loading, geometric and designparameters on the performance of composite moment frames.

The nonlinear-static behaviour of a typical compositeframe is firstly assessed by comparing the response obtainedfrom triangular and uniform pushover loading patterns. Thevalidity of adopting nonlinear-static approaches for this typeof structure is evaluated by comparison against the resultsof incremental dynamic analysis. For this purpose, sevennatural earthquake acceleration records, which are selectedand adjusted for compatibility with the design spectrum, areutilised. It is shown that the results of the dynamic time-history analysis are closely related to the nonlinear-staticresponse adopting triangular loading pattern, following thetrends observed in other frame types of regular configurations.The suitability of pushover analysis with triangular loading,in comparison with the uniform pattern, is supported by theoverall structural response as well as local inter-storey driftdistributions.

A degree of global overstrength is normally expected inframed structures, particularly when relatively high behaviourfactors are adopted, since member sizes are often governedby gravity considerations or by deformation-related criteria.This overstrength becomes even more pronounced in compositeframes, due in-part to the typically large spans which increasethe dependence of beam sizes on gravity loading conditions.The frames examined in this study exhibited strength levelsconsistently exceeding four-fold the assumed design base shear.The global overstrength is also contributed to by the significant


redistribution in terms of the ultimate-to-yield resistance ratio,referred to as αu/α1 in EC8. This ratio was higher than 1.5 inall the study frames (which is well above the recommendedcode value of 1.3) provided that local storey mechanisms areprevented and a realistic level of gravity loading is considered.

It is shown in this paper that several design parametersand assumptions have direct implications on the inelasticbehaviour of composite frames, as assessed through the overalllateral response, inter-storey drift distribution and plastic hingepatterns. In particular, the choice of behaviour factor inrelation to the seismic intensity, the incorporation of momentredistribution under gravity conditions, and the panel zoneeffects are all issues that merit careful consideration in design. Itis also shown that a number of geometric parameters, related tothe structural configuration, including frame spacing and beamspan, have a significant influence on the behaviour.

The plastic hinge patterns observed in the study framesprovide useful insight into the behaviour. In most cases, exceptin frames with relatively short beam spans or when significantpanel zone yielding is permitted, hinges form primarilyon one side of the beam spans under pushover loading.This effect is particularly significant in composite framesdue to the characteristic asymmetry in beam cross-sectioncapacity, coupled with the influence of gravity moments. Moreimportantly, local storey mechanisms did not occur in thestudy frames except in the case of the three-storey structure.This undesirable performance, to which low-rise and/or largespan configurations are particularly vulnerable, highlights theneed for careful interpretation or modification of the suggestedcapacity design provisions.

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Elghazouli 2008 Engineering-Structures