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Effect of earthquake damage on the behaviour ofcomposite steel frames in fireDOI:10.1177/1369433218761138

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Citation for published version (APA):Suwondo, R., Gillie, M., Cunningham, L., & Bailey, C. (2018). Effect of earthquake damage on the behaviour ofcomposite steel frames in fire. Advances in Structural Engineering. https://doi.org/10.1177/1369433218761138

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Effect of earthquake damage onthe behaviour of composite steelframes in fire

Riza Suwondo1 , Martin Gillie2, Lee Cunningham1 and Colin Bailey3

AbstractFire loading following earthquake loading is possible in any building in a seismic-prone area. However, most design approaches do notconsider fire following earthquake as a specific loading case. Moreover, seismic design philosophies allow a certain degree of damagein structural elements which make structures more vulnerable when subjected to post-earthquake fire. This study uses three-dimensional numerical models to investigate the effect of earthquake damage on the fire resistance of composite steel-frame officebuildings. A total of two types of earthquake damage, fire insulation delamination and residual lateral frame deformation, are investi-gated. It is concluded that earthquake damage can significantly reduce the fire resistance of composite buildings, with delamination offire protection having the greatest effect. The results of this study can be used by designers to improve the post-earthquake fire resis-tance of composite buildings.

Keywordscomposite construction, earthquake engineering, fire engineering, fire following earthquake, fire insulation

Introduction

Fire occurring after earthquake has been demonstratedto be a major threat for buildings in seismic-proneregions. Damage caused by fire following earthquake(FFE) can be worse than that caused by the earth-quake itself (Scawthorn et al., 2005). Recorded experi-ences indicate that FFE has caused numerous deathsand extensive losses. For example, it was estimated that80% of total damage and fatalities in the 1906 SanFrancisco seismic event were caused by fires followingthe earthquake (Scawthorn, 2008). In the 1923 Tokyoearthquake, the losses due to fire following the earth-quake accounted for over 70% of total building lossesand resulted in 140,000 deaths (Scawthorn, 2008).More recently, in the 1995 Kobe earthquake, therewere approximately 7000 buildings destroyed by FFE(Faggiano, 2007).

Although the damage caused by post-earthquakefire can be very significant, it is not considered as aspecific loading case in seismic design. Moreover, thecurrent philosophy of seismic design permits a certaindegree of damage to the structural elements, connec-tions and fire proofing during seismic action, whichpotentially makes structures more vulnerable whensubjected to post-earthquake fire. With the substantialdamages and losses recorded in past events, it seems

necessary to understand and quantify the behaviour ofstructures under multi-hazard events such as FFE.

Several previous studies have been conducted on thebehaviour of earthquake-damaged steel frames in fire.Della Corte et al. (2003) developed numerical modelsto investigate the response of steel moment resistingframes (MRFs) under post-earthquake fire scenarios.They showed that the drift ratio is an important para-meter that affects the fire resistance. In addition, it wasobserved that the seismic design philosophy adoptedcan significantly affect the performance of steel struc-tures under post-earthquake fire. Faggiano et al.(2008) applied pushover and coupled thermal–mechanical analysis to evaluate two-dimensional (2D)steel MRFs subject to FFE. The results showed thatthere was a small reduction in fire resistance when thestate of a frame after earthquake reached near collapselevel. However, the fire resistance of the frames was

1School of Mechanical, Aerospace and Civil Engineering, University of

Manchester, Manchester, UK2School of Engineering, University of Warwick, Coventry, UK3Queen Mary University of London, London, UK

Corresponding author:

Riza Suwondo, School of Mechanical, Aerospace and Civil Engineering,

University of Manchester, Manchester, M13 9PL, UK.

Email: riza.hakim@postgrad.manchester.ac.uk

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still the same when the frame did not exceed the opera-tional performance level limit.

Zaharia and Pintea (2009) presented post-earthquake fire resistance of two different seismicallydesigned frames (moderate and severe seismic regions).The results showed that the frame designed for stron-ger seismic effects had a higher fire resistance rating inthe case of post-earthquake fire. Memari et al. (2014)presented post-earthquake fire performance of MRFswith reduced beam section (RBS) connections. Post-earthquake fire exposure was only at the location ofthe RBS connections based on the assumption that fireinsulation was damaged at the RBS connection afterearthquake. The results showed that a post-earthquakefire produces smaller inter-storey drift ratios (IDRs)than the earthquake itself. No structural collapseoccurred as a result of applied post-earthquake fires.Behnam and Ronagh (2015) performed pushover andthermal analyses to investigate the behaviour of 2Dsteel frames. Fire analysis only was also performed onan undamaged frame, as a benchmark. The resultsrevealed that the damaged frames have lower fire resis-tance than the corresponding undamaged frame.

All of the aforementioned studies have focused onnon-composite steel frames. Most of these structureshave been analysed assuming 2D plane-frame beha-viour, without considering the effects of out-of-planebehaviour. By contrast, this study focuses on compo-site steel-frame, concrete floor construction as is com-monly used in mid-rise commercial frame structures.Although some essential aspects of the behaviour ofthese structures in fire can be captured in two dimen-sions, such an approach misses load redistribution inthe out-of-plane direction which is a key load-carryingmechanism available at large deflections, as identifiedby the Cardington fire tests (Gillie et al., 2001; Kirby,1998). Thus, when a composite slab is present, three-dimensional (3D) behaviour must be considered in ananalysis to obtain realistic behaviour on a steel framein fire. Additionally, previous research on post-earthquake fire behaviour has focused on unprotectedsteel frames. There is as yet a lack of detailed researchinto the influence of fire insulation on steel frames dur-ing FFE.

To address these knowledge gaps, this study extendsSuwondo et al.’s (2017) previous work on post-earthquake fire behaviour of composite steel-framedstructures. The 3D numerical models are employed tosimulate post-earthquake fires on earthquake-damagedstructures. A total of two types of damage, fire insula-tion delamination and residual lateral frame deforma-tion, are investigated to represent earthquake damage.The consideration of the two types of damage will beexplained in the section titled ‘Earthquake damage’.

The main objectives of this study are to understand theinfluence of the following:

1. Delamination of the fire insulation on primarybeams due to earthquake damage;

2. Delamination of the fire insulation on columnsdue to earthquake damage;

3. Residual deformation on the post-earthquakefire behaviour of composite steel frames.

The remainder of this article is structured as fol-lows: a description of the generic building studied, thetypes of damage considered, the modelling approachadopted, the results of the study, and conclusions.

Generic building

A generic five-storey composite steel frame is analysedin this study. Figure 1 shows the structure’s plan andelevation. The building is designed for high seismicityusing a MRF with medium ductility according toEurocode EN 1993-1-1 (CEN, 2005a), EN 1994-1-1(CEN, 2004a) and EN 1998-1-1 (CEN, 2004b). In thisstudy, steel behaviour was assumed to be elastic-perfectly plastic for both columns and beams with ayield stress of 355 MPa and Young’s modulus of 210GPa at ambient temperature. The compressivestrength of concrete in the slabs was taken as 30 MPa,and the yield strength of the rebar is 450 MPa.Thermal expansions of steel and concrete are taken as1.35 3 105�C21 and 9 3 105�C21, respectively. Theconcrete and steel properties at elevated temperaturesfollow the recommendations in Eurocode EN 1992-1-2(CEN, 2004c) and EN 1993-1-2 (CEN, 2005b). Theconcrete slabs have a thickness of 130 mm and therebar mesh consists of 6-mm diameter bars at 200-mmcentres each way as shown in Figure 1(c).

Primary beams and columns are assumed to be pro-tected with a 10-mm thickness of lightweight insulatingmaterial which has thermal conductivity of 0.2 W/mK,specific heat of 1100 J/kgK and density of 300 kg/m3.In order to utilise tensile membrane action in the slabpanels, which are surrounded by the primary beams,the secondary beams are left unprotected, as is com-mon in a performance-based fire design. The totaldesign load at the fire limit state (1 3 dead + 0.5 3live) is taken as 5.5 kN/m2. This load level results in aload ratio for the internal columns at bottom level andprimary beams of 0.3 and 0.5, respectively. The loadratio is defined as the ratio of the applied load to thebearing capacity of a member. Although permitted bythe Eurocode EN 1991-1-1 (CEN, 2002a), no furtherload reduction is considered in the structure in orderto be conservative. For simplicity, the columns are

2 Advances in Structural Engineering 00(0)

assumed to have the same section size at all heights.Only the ground floor is exposed to fire. The groundfloor is assumed to have the worst fire scenario sincethe columns have the largest load ratio compared tothe upper floor. Moreover, the plastic hinges will onlyform during earthquake in the columns at groundfloor so they are most likely to be damaged at groundlevel.

Earthquake damage

A key aspect of post-earthquake fire analysis is deter-mining a structure’s physical state after an earthquakeas this represents the initial condition for the subse-quent fire event. It is well known that providingdetailed information regarding earthquake damage isdifficult due to randomness and uncertainties of bothstructural properties and earthquake vibration.Accordingly, in this study, two generic types of earth-quake damage are considered: fire insulation delamina-tion and residual lateral frame deflections as a result ofplastic deformations developed during the earthquake.

Fire insulation delamination

The philosophy of seismic design permits a certaindegree of damage to structural elements. Past eventshave demonstrated that earthquakes can cause damageto active fire protection systems such as sprinklers and

reduce the effectiveness of firefighting capability. Thus,passive fire protection systems, such as sprayed fire-resistive material (SFRM), play a critical role in miti-gating the effect of the post-earthquake fire on thestructural system in a building. However, SFRM mayalso be damaged by becoming detached. Both experi-mental and field observations have indicated thatSFRM can delaminate under static and dynamic loadsituations (Arablouei and Kodur, 2015; Braxtan andPessiki, 2011; Kodur and Arablouei, 2015; Wang et al.,2013). SFRM delamination is most likely to occur inlocations of high strain, such as where plastic hingesform.

The effect of fire insulation delamination on thesteel can jeopardise the structural stability of the build-ing. There are several studies focused on studying theeffect of fire insulation delamination on the fire resis-tance. Tomecek and Milke (1993) investigated thebehaviour of steel columns with loss of protectionmaterial based on predicted thermal response and ther-mal criteria. The results showed that there is significantreduction in the fire resistance of steel columns. Milkeet al. (2002) also investigated the effect of the loss offire protection on the thermal response of the steel col-umn. They found that the temperature rise in the col-umn is mainly caused by the area of missing fireprotection regardless of the protection thickness. Theyconcluded that the column size has a relatively smalleffect on the reduction in fire resistance. Wang and Li

Figure 1. Generic frame: (a) plan view of building, (b) elevation view of structural frame and (c) composite floor slabs.

Suwondo et al. 3

(2009) conducted an experimental study to investigatethe behaviour of steel columns with partial loss of fireinsulation. They found that the failure mode of speci-mens can be yielding when the fire protection damageis short or buckling when the fire protection damage islong.

The above studies indicate that passive fire protec-tion system damage has a significant role in determin-ing the vulnerability of steel structures to FFE.Therefore, it is important to quantify the delaminationin the assessment of a building subjected to FFE.Accordingly, in this study, it is assumed that fire insu-lation delamination may occur at both ends of theprimary beams and at the bottom of the columnswhere the plastic hinges may occur during an earth-quake, as shown in Figure 2. In these delaminationregions, the steel is assumed unprotected. The length

of delamination is assumed to be 5% or 10% of themember length.

Residual deformation

Residual frame deformation is the irreversible lateraldeformation that remains after an earthquake. Thisdeformation can be dangerous since it enhances P–Deffects that can increase stress on the structure underloads acting after the earthquake as illustrated inFigure 3. The repetitions of plastic deformations dur-ing the earthquake also cause some reduction ofmechanical properties. However, for MRFs designedaccording to Eurocode 8, the plastic deformation isrelatively small. In this case, the reduction of mechani-cal properties is negligible (Della Corte et al., 2003).Therefore, an ideal elastic-perfectly plastic structurewith a non-degrading component is considered in thisstudy.

As discussed in Bruneau et al. (1997) and DellaCorte et al. (2002), the structural damage is mainlycaused by IDR. Thus, there is good agreement that theIDR can be used for measuring damage of a frameunder an earthquake. The IDR is the ratio betweenstorey displacement relative to the adjacent storey andthe storey height. IDR may also be associated withlocal plastic deformation (Gupta and Krawinkler,2000).

The level of residual deformation after an earth-quake was here estimated by performing a ‘pushoveranalysis’. In a pushover analysis, the building is pushedincrementally using a specific lateral load to arrive at atarget displacement, and the load is then reduced tozero again. Lateral storey forces on the structure areapplied in proportion to the product of storey massand fundamental mode shape (ATC, 1996). The targetdisplacement (represented by the roof displacement) isthe expected displacement of the building when sub-jected to a design earthquake. The residual deforma-tion pushover analysis results can be used as an initialstep for the subsequent fire event. Pushover analysis isan approximate method, and dynamic effects are notdirectly considered.

According to ASCE standard 41-06 (ASCE, 2007),the level of earthquake damage can be categorised intothree different performance levels: immediate occu-pancy (IO), life safety (LS) and collapse prevention(CP) that represent minor to major damage. These per-formance levels are tied to the IDR as an indication ofglobal stability of the structure. The IDR value is lessthan 0.7%, 0.7%–2.5% and 2.5%–5% for perfor-mance level of IO, LS and CP, respectively.

This study considers two different levels of damages,damage 1 and damage 2. Damage 1 is LS level withIDR 2%, and damage 2 is CP level with IDR 4.5%.

Figure 2. Fire insulation delamination on the beam andcolumn.

Figure 3. Residual deformation of structure after earthquake.

4 Advances in Structural Engineering 00(0)

Local buckling is not a possible failure mechanism inthe composite structural form analysed and so was notconsidered here. This assumption is in line with thephilosophy of seismic design that performance levelsof structures do not exceed the intended level whensubjected to design earthquake. Material deteriora-tion other than plasticity was also not considered.Steel in the case of damage states considered here isunlikely to suffer permanent changes to its key para-meters. Concrete may suffer from cracking duringearthquake, which is crudely captured in the modeland possibly spalling during a fire. However, experi-mental and real-fire experience suggests that theseeffects are at worst secondary matters when consider-ing structural behaviour during a fire. Spalling in thecase of steel-deck composite construction is largelycontained, for example. According to previous stud-ies, spalling becomes important in elements withmore than 4- to 5-cm cover (Majorana et al., 2010) ormade of high-strength concrete (Kodur, 2005).

The two damage types discussed above are consid-ered to represent earthquake damage in this study. Itis apparent that when fire insulation delaminationoccurs, residual deformation always exists as a resultof plastic deformation. Thus, the combined effect offire insulation delamination and residual deformationcould be considered. However, the aim of this study isto examine and quantify the effect of the two damagetypes separately in order to investigate which damagemost significantly affects fire behaviour. For this rea-son, the two types of damage have been analysed sepa-rately in this study.

Post-earthquake fire loading and failure

A total of two different fire scenarios, a ‘standard fire’(ISO, 1975) and parametric fire EN 1991-1-2 (CEN,2002b) as shown in Figure 4, were used to simulatepost-earthquake fires in this research. The fire load of511 MJ/m2 taken in the parametric fire is calculatedaccording to Annex E of EN 1991-1-2 (CEN, 2002b),which is for office buildings. It should be noted thatthe parametric fire curve depends on several factors,such as ventilation, fire load, fire protection, and soon. When an earthquake occurs, these factors maychange, which makes the appropriate parametric firecurve for use before an earthquake different from thatfor use after an earthquake. Opening factors of 0.06and 0.08 are taken for the parametric fire before andafter earthquake, respectively. This assumption isconsidered as there is a possibility of window break-age or damage to the building envelope after anearthquake that can increase the opening factor ofthe compartment. Moreover, firefighting measuresmay be different between fire in normal conditions

and fire after earthquake. Using appropriate para-meters from Table E.2 in Eurocode EN 1991-1-2(2002b), the assumed parametric fire after an earth-quake (PEF) results in higher temperatures and alonger period than those before an earthquake (F) asshown in Figure 4.

Since in modern buildings, open-plan offices are themost common, full-floor compartments are consideredand the hot gases are assumed uniform in the fire com-partment. With these gas temperature–time curves, theprocedure in Eurocode EN 1993-1-2 (CEN, 2005b)was used here to calculate the steel temperatures and anumerical heat transfer analysis to obtain concrete slabtemperatures. Temperature–time curves for each firescenario and for the steel and concrete are plotted inFigure 5.

In this study, two approaches are applied to deter-mine the capacity of the frame structure exposed topost-earthquake fire. First, failure is defined when theprimary beams are not able to transfer loads effectivelyor to reliably provide compartmentation. This stagecan be crudely characterised when the mid-span deflec-tion of a beam exceeds L/20, where L is beam span(Dwaikat and Kodur, 2011). Second, failure is definedby buckling of columns which is identified when verti-cal displacement of the top of a column rapidlyreduces as temperature increases (Rezvani andRonagh, 2015). It should be acknowledged that inboth cases, the notional threshold of failure is some-what arbitrary. However, for the purpose of this study,these serve as a reasonable and practical measure ofstructural performance.

Numerical model and validation

The general finite element software ABAQUS wasused to analyse the generic building. Steel beams andcolumns were discretized using two-node linear beamelements (B31). The concrete slabs were modelledusing four-node shell elements with reduced

Figure 4. Gas temperature–time curves.

Suwondo et al. 5

integration (S4R). To represent rebar in the slab,layers with appropriate steel material properties werespecified in the shell elements. Tie constraints betweensteel beams and the concrete slabs were applied to rep-resent full composite action. For simplicity, it wasassumed that the beam-to-column and secondarybeam-to-primary beam connections behaved as rigidand pinned connections, respectively, which approxi-mates common practice for steel frames designed forseismic regions. Connection failures are not consideredhere.

Structural performance of elements under fire canbe affected by surrounding structure, due to the effect

of thermal restraint (Ali and O’Connor, 2001; Laı́mand Rodrigues, 2016; Rodrigues and Laı́m, 2017;Valente and Neves, 1999). The axial restraint can gen-erate extensive unforeseen forces in the element duringfire that may cause unpredictable structural failure, incontrast to the rotational restraint, which can avoid asudden failure. Thus, 3D frames as shown in Figure 6are developed to consider thermal restraint affected bysurrounding structure. A mesh size of 0.5 m 3 0.5 mis used for the first floor slab, and a mesh size of 1.0 m3 1.0 m is used for the upper slab to save the comput-ing cost. A total of 8 elements are meshed for all col-umns and 18 elements for the beams.

Figure 5. Structural temperature: (a) standard fire, (b) parametric fire before earthquake (F) and (c) parametric fire afterearthquake (PEF).

6 Advances in Structural Engineering 00(0)

A non-linear static analysis procedure was usedbecause it was less time-consuming and easier to inter-pret than a dynamic procedure. Numerical conver-gence problems were prevented by use of artificialviscous damping in the analyses. An appropriate dissi-pated energy fraction should be specified by trial anderror (Abaqus, 2014). However, the results’ accuracycan be affected significantly if the fraction is too high.In this research, the default energy dissipation factorof 0.2 3 103 is used, and the ratio of the energy dissi-pated by viscous damping to the total strain energy islimited by an accuracy tolerance of 5%.

Using existing experimental and analytical workavailable in the literature, a series of validations werecarried out to confirm that the results from the analysisgive an acceptable level of accuracy. To validate thepushover analysis, the ‘European calibration frame’previously analysed by Vogel (1985) was chosen. Theframe was subjected to proportionally distributed grav-ity loads and concentrated lateral loads as shown inFigure 7. The yield stress of all members was 235 MPa,and Young’s modulus was 205,000 MPa. The result inFigure 8 shows good agreement between this study andthe previous analysis and confirms that the analysisconsiders P–D effects correctly.

To validate the fire analysis, a steel–concrete com-posite frame was chosen which is a simplified versionof the Cardington British Steel fire test (Kirby, 1998)and which was previously analysed by Gillie (2009).Figure 9 shows the geometry of the composite frame.A load of 5.48 kN/m2 is applied as a gravity load overthe slab. Temperature loadings are applied on theshaded area. The secondary beam is heated to 800�C.The lower surface of the slab is heated to 600�C with alinear gradient of 4.6�C/mm. The structure was then

cooled to ambient temperature. The comparison inFigure 10 shows that the general pattern and magni-tude of the deflections are in good agreement in allcases and also confirm that the model can be used forthermal–mechanical analysis of steel–concrete compo-site frames.

Figure 6. Finite element mesh of the generic frame.

Figure 7. Dimension and loading condition of the Europeancalibration frame.

Figure 8. Load-deflection behaviour of the frame.

Suwondo et al. 7

Results and discussion

Impact of delamination of fire insulation on theprimary beams

First, frames with and without delamination of the fireprotection on the beams were analysed. The unpro-tected secondary beams, the primary beams with andwithout delamination and the slab were heated. Inorder to clearly investigate the effect of delaminationon the primary beams, the column remained at ambi-ent temperature. This is because heating columnswould conceal the influence of delamination on theprimary beams.

Figure 11 shows mid-span deflections of primarybeam B12 (see Figure 1). It can be seen that the beamdeflection exceeds the limit for all fire scenarios exceptunder a parametric fire before earthquake. In general,the delamination reduces failure time of the beams, butthe effect of the length of the delamination on the beamdeflections is small.

It can be observed that the influence of fire insula-tion delamination is negligible, up to 130�C. Butbeyond this temperature, there is a sudden increase indeflection. This is expected because the temperature ofthe delamination region reaches 400�C, the point atwhich the yield stress of steel starts decreasing. The dif-ferences become more obvious when temperaturesincrease further. However, as can be seen in Figure 11,the differences become insignificant again when thetemperature is above 800�C using the standard fire. Atthis point, the temperature difference between pro-tected and unprotected beams becomes small (due toprolonged high temperatures overwhelming the fireprotection) so that the delamination has a small effect.Besides, at high temperature, the role of the concreteslab becomes dominant in supporting the loads sincethe beams have lost almost all strength and stiffness.

Figure 12 shows a comparison of the vertical deflec-tion contours of concrete slabs without delaminationand with 10% delamination. It shows that the

Figure 9. Geometry of simplified version of the first Cardington test (Gillie, 2009).

Figure 10. Comparison of mid-span beam deflections.

8 Advances in Structural Engineering 00(0)

delamination influences tensile membrane action in theconcrete slab. When no delamination occurs, tensilemembrane action can be fully mobilised and the slabbends into a bowl-like shape. However, the delamina-tion on the beam reduces the vertical support to theslab. The tensile membrane action is reduced, and thedeflected shape is more like that of a one-way slab orlike catenary action in a beam. This behaviour willhave a negative effect on structural strength sincecatenary action provides a weaker load-carryingmechanism than tensile membrane action. In this case,the benefit of tensile membrane action cannot beassumed in design.

Impact of fire insulation delamination on thecolumns

In this case, the delamination occurs only in the columns.The whole ground floor including beams, columns andslab are exposed to fire. As discussed earlier, the columnfailure is defined when the vertical displacement of the col-umn drops suddenly. Figure 13 shows vertical displace-ment of the top of column C2 (see Figure 1) for all cases.The fire resistance time of the column very significantlyreduces (by around 70%) when the delamination occurs.

It can be observed that the column vertical deflec-tions are initially positive due to thermal expansion.

Figure 11. Mid-span beam deflection of primary beam B12 against beam temperature: (a) ISO 834 standard fire, (b) EC parametricfire before earthquake and (c) EC parametric fire after earthquake.

Suwondo et al. 9

Figure 12. Concrete slab vertical deflection contours after 90-min standard fire exposure (unit in m).

Figure 13. Vertical displacement of C2: (a) ISO 834 standard fire, (b) EC parametric fire before earthquake and (c) EC parametricfire after earthquake.

10 Advances in Structural Engineering 00(0)

After a smooth and low rate of generation, the verticaldeflection reaches a maximum value and then startsdecreasing as the mechanical properties of the steelcolumn reduce with time and temperature. At failure,the column starts to buckle, and the deflections reverse,when the axial force reaches a peak as shown in Figure14.

Figure 15 shows horizontal displacements againstthe height of column at certain gas temperatures. Itcan be seen that the effect of delamination is still negli-gible at gas temperatures up to 500�C. After that, thedelamination increases the horizontal deflection of thecolumn. The deflection of the column with 10% dela-mination is much higher than that of the column with5% delamination at high temperature. This deflectionincreases P–D effects on the columns, but it is

essentially insignificant on the fire resistance of the col-umns. This is in line with the previous study by Wangand Li (2009) that showed that the length of fire insu-lation delamination on the column has small effect onthe failure time.

Impact of residual deformation

As explained in the section titled ‘Earthquake dam-age’, the residual deflection damage is represented bypushing the building horizontally prior to fire. Thiscan be achieved by performing a three-step procedure.First, the building is subjected to gravity load. Second,a non-linear pushover analysis is performed to simu-late the earthquake. Figure 16 shows lateral load ver-sus top storey displacement. It can be seen that the

Figure 14. Axial force of C2: (a) ISO 834 standard fire, (b) EC parametric fire before earthquake and (c) EC parametric fire afterearthquake.

Suwondo et al. 11

pushover analysis results in residual deformation thatcan be used as the initial step for the subsequent fireevent. Damage 1 and damage 2 have residual deforma-tions at the first floor level of 0.08 m (IDR 2%) and0.18 m (IDR 4.5%), respectively. Finally, a thermal–mechanical analysis of the damaged frame is con-ducted to simulate a fire.

Figure 17 shows mid-span deflection of primarybeam B12 with and without damage. In general, thefire resistance on the beam slightly reduces due to theinitial deflection after damage. It can be seen that thereis an initial deflection in the primary beam for the dam-aged frame. However, the patterns of deflection in firein the undamaged and damaged frames are very simi-lar. For the case of standard fire and parametric fireafter earthquake, the analysis stops before the beam

deflections exceed the limit since failures in the heatedcolumn due to buckling occur earlier than failure in theheated beams. Hence, fire resistance of the compositeframe is determined by failure in the column.

Figure 18 shows vertical deflection of the columnC2. As can be seen, there is no collapse in the buildingwhen subjected to parametric fire before earthquake,but in other cases, runaway failure occurs. Overall, thedamaged structure is more vulnerable than the unda-maged structure. This is because the residual deforma-tion increases moments in the column after earthquakedue to P–D effects. The residual deformations reducefailure time from 98 to 92 min and from 78 to 76 minwhen exposed to standard fire and parametric fire afterearthquake, respectively.

Summary

This study has presented a numerical investigation of asteel-framed composite floor system under FFE. Atotal of two types of earthquake damage, fire insula-tion delamination and residual deformation, were con-sidered. Failure of the structure was defined using twomeasures, a beam deflection exceeding span/20 andcolumn buckling.

For clarity, Table 1 shows failure times for all casesexamined. It can be seen that the earthquake damagecan significantly reduce the fire resistance of the com-posite building. The reduction in fire resistance timesresult mainly from fire insulation delamination,

Figure 15. Horizontal displacement of C2 exposed to ISO 834 standard fire: (a) gas temperature at 500�C, (b) gas temperature at700�C and (c) gas temperature at 825�C.

Figure 16. Pushover curves for damage 1 and damage 2.

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particularly in the columns, rather than residual defor-mation. Moreover, the results show that there is nofailure for the structure exposed to FFE using a para-metric fire suitable for the pre-earthquake conditionexcept when the delamination occurs in the columns.Therefore, the building can resist the applied load dur-ing earthquake or fire separately. However, the earth-quake event can result in damage to the structures thatmake the structure more vulnerable to post-earthquakefire than the corresponding undamaged structure.Besides, a fire after earthquake can be more onerousthan that in normal conditions since an earthquakemay increase the compartment opening factor bybreaking the windows and also by affecting the effi-ciency of firefighting.

Conclusion

Based on this study, the following conclusions can bemade for the structural configuration and damagetypes investigated:

1. Delamination of fire protection at the ends ofprimary beams, as is likely with seismic eventinduced damage to ‘strong-column, weak-beam’ structures, means that tensile membraneaction may not be fully available as a reserveload-carrying mechanism in FFE (Figure 11).This has significant consequences for designbecause the benefits of tensile membrane actionare often used for performance-based firedesign of composite structures. Either designers

Figure 17. Mid-span deflection of primary beam B12: (a) ISO 834 standard fire, (b) EC parametric fire before earthquake and(c) EC parametric after earthquake.

Suwondo et al. 13

Figure 18. Vertical displacement of C2: (a) ISO 834 standard fire, (b) EC parametric fire before earthquake and (c) EC parametricfire after earthquake.

Table 1. Failure time (min).

Delamination Standard fire Parametric fire (F) Parametric fire (PEF)Delamination on beam

0% 130 – 1005% 121 – 9210% 120 – 90

Delamination on column0% 100 – 775% 20 23 2110% 20 20 20

Residual deformationUndamaged 98 – 77Damage 1 98 – 77Damage 2 92 – 76

14 Advances in Structural Engineering 00(0)

should specify fire protection systems in theseregions that are resistant to the deformationsand strains that are likely to develop or thebenefits of tensile membrane action should notbe relied upon. The results are only weaklydependent on the length of the assumed delami-nation, so consideration of beam length compa-rable with a plastic hinge is sufficient.

2. Delamination of fire protection at the bottomof columns will cause a large reduction in fireresistance time (up to 70%) because columnswill buckle with no alternative load paths avail-able. This finding is in line with earlier experi-mental work (Kirby, 1998) and consequentrecommendations that columns are always fullyfire protected. For design purposes, either col-umns with sufficient bending strength to avoiddelamination of fire protection should be used(with due consideration of the implications of thisfor seismic design) or, again, deformation resis-tant fire protection systems should be specified.

3. Residual lateral deflections of earthquake-damaged structures have a significant but lessereffect on their fire resistance. The increase in P–Dmoments in the columns may lead to prematurefailure earlier than in an undamaged structure,and this should be accounted for in design, butno particular changes to fire protection or struc-tural systems appear to be required.

4. Fire protection systems that are tolerant of sub-stantial strains should be investigated for appli-cation to building structures. Such protectionsystems do exist for application where blastresistance is needed, but they are expensive. Itmay be that local application of such systems inregions where plastic hinges are likely to formdue to earthquake damage would be a practicaland efficient solution.

5. The choice of design fire for use in an FFEanalysis is not obvious. This study shows signif-icant differences in predicted response betweenstandard and parametric fires and between dif-ferent parametric fires. The best choice of para-metric fire is probably structure dependent andmay not be apparent at the design stage. Arange of likely fires should be considered ifperformance-based fire design is used for FFEloading.

Declaration of Conflicting Interests

The author(s) declared no potential conflicts of interest withrespect to the research, authorship and/or publication of thisarticle.

Funding

The author(s) disclosed receipt of the following financial sup-port for the research, authorship and/or publication of thisarticle: This PhD research was funded by the IndonesiaEndowment Fund for Education (LPDP).

ORCID iD

Riza Suwondo https://orcid.org/0000-0002-4917-2968

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