Composite Floors Part 1

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P A P E R : B a i l e v / M o o r e

Paper

C. G.Bailey,BEng, hD,

Building Research

Establishment

D. B. Moore,

BTech, PhD, CEng,MIStructE

Building Research

Establishment

Keywords:

steel frames,

floors, composite

construction,

fire resistance

behaviour,

fire tests, reinforced

concrete slabs

The structural behaviour of steel

frames with composite floorslabssubject t o fire: Par t 7: Theory

Synopsis

This paper presents the development o f a new

design method for calculati nghe performance of

steel framed buildings, wi th composite flooring

systems, subject to fire. This design ethod is based

on the results from a series of full-s caleire tests on

an eight-storey steel framed building, together with

associated theoretical and further xperimentalinvestigations. The results from thiswork show

that the performance fcomposite steel deck

flooring systems in fire is under-utilised n current

design procedures. This iswing mainly to the

ability ofightly reinforced omposite slabs to

bridge over the supporting fire-damagedsteel

beams and t ransfer load, using membrane action,

to the undamaged art s ofhe steel structure. From

these observations a simple designmethod is

developed that isbased on a holistic, rather than

an elemental, approach. This allows thearious

interactions between the components ofa

composite slab, supported by a grillage of teel

beams, to be taken into ccount, producing cheaper

and more innovative, site-specific, fire-engineeringsolutions. A companion paper shows how this

design method can be applied to practical

buildings.

Notation

Volume 78/No 7 7 6 un e 2000

is theaspect ratio(LIZ)

is theparameter-defining the magnitude of

membrane force

is theeffective depth of reinforcement

is theenhancement of yieldline load due to

membrane action

is theenhancement of element 1due to

membrane forces

is the enhancement of element 2 duet o

membrane forces

is theenhancement of element 1due to bending

action

is the enhancement of element 2 due t o bending

action

is theyield stress

is theparameter-fixing depthof the compressive

stress block when no membrane force is present

is theparameter-defining the magnitude of

membrane force

is the shortest spanf the rectangular slab

is the largest spanf the rectangular slab

is themoment capacityof the composite slab at

the fire limit sta te

is the esistance momentlunit width t ayieldline when membraneorces are present

is themoment about he support due o

membrane forces for element 1is themoment about the support due o

membrane forces for element 2

is the moment capacity f the composite beamat

the fire limit sta te

MO is the moment capacity of the composite slab

n is he parameter-defining the yieldline pattern

p is he loadcarryingcapacity of the composite slab

U is he maximum vertical displacementof a

W is he deflectionof

yieldlinetup0 is the loadcarrying capacity of the composite floor

and grillage of composite beamsat a temperature8

wbeame is the loadcarrying capacity f the grillage of

composite beams a t a given tempera ture 8

x is he parameter-defining the intersection of

yieldlines

E is the train in the reinforcement

y is the parameter-defining the loaded area

when no membrane force is present

at the ire limit stat e

parabolic curve

supported by a composite beam

is the oad levelo r load ratio

Introduction

The traditional method of ensuring tha t a steel frame

building, with a composite flooring system, satisfies the

regulatory requirementsor fire resistance iso protect all

the columns and all the exposed surfaces of the downstand

supporting steel beams with a proprietary nsulating

material. Although this approach has proved to be satis-

factory, the current method1 for specifying the required

thickness of the insulating material s extremely conser-

vative, since t ignores the inherent fire resistance of the

structure. Furthermore, the prescriptive nature of this

approach can mpose unnecessary costs and stifle innova-

tion. The developmentof structural fire design Codes, (e.g.

BS 5950:Part 8 2 and the European prestandardsurocode

3:Part 1.23 and Eurocode4: art 1.24)provide a more solid

scientific foundation for the provision of fire resistance.

However, these design Codes were developed from s tan-

dard fire tests5 on isolated beams and columns and typi-cally ignore significant struc tural behaviour by disre -

garding the interaction between members.

It has become increasingly clear, in recent years, that

continued research into the performanceof individualcom-

ponents should not e a highpriority. Better understand-

ing of the interactions etween different components, ead-

ing t o an appreciation of the way in which complete

struc tural systems function, when subjecto fire, is likely

to prove more cost-effective. urthermore, there is gener-

al agreement that the struc tural ontribution of compos-

ite flooring systems, comprising teel deckkoncrete com-

posite floorslabs supported y a grillage of steel beams, is

under-utilised in current fire design procedures. This,

togetherwith nvestigations rom eal fires, such as

occurred a t Broadgates, suggests hat theperformance ofcomplete structures is s ignificantly better than thatf the

single element from which fire resistance is currently

assessed.

The developmentof BRE's Large Building Tests Facility

(LBTF)at Cardington provided the construction industry

with a unique opportunityo carry out ull-scale fire tests

on a complete building designed and built t o current prac-

19



P A P E R : Ba i l e v /Moo r e

tice. Consequently, in the 1990s a series of six compart-

ment fire tests were conducted ona full-scale steel framed

building a t Cardington. The test building was eight storeys

high and covered a plan areaof 21m x 45m. The compos-

ite flooring system comprised steel downstand beams act-

ing compositely with a floorslab, constructed using a trape-

zoidal steel deck, lightweight concrete, and an anticrack

A142 steel mesh. The overall depth of the slabwas 130mm

thick, with the mesh placed 15mm above the steel deck.

Bailey e t aZ7 and Martin e t U P give a detailed account of

these t ests, which confirmed tha t th e performance of a

steel frame, with a composite flooring system, is signifi-

cantly bet ter than that suggested by current fire design

methods.

Observations from these ests uggested hat he

improved performance of composite flooring systems is

due to th e ability of lightly reinforced composite slabs to

bridge over the supporting ire-damaged steel beams and

transfer load, using membrane action, to theundamaged

parts of the steel structure. The enhanced loadcarrying

capacity of lightly reinforced concrete labs has been inves-

tigated by a number of authorsg-13 but , because of the very

large displacementsassociated with this mode of behav-iour, it was previously considered o have no application in

the design of buildings. Although this is true for normal

temperature design, at elevated temperatures large dis-

placements are acceptable provided that structural col-

lapse or breach of compartmentation is avoided. This

prompted Bailey14 to inves tigate theoadcarrying capaci-

ty of lightly reinforced concrete slabs a t large displace-

ments. Based on observations from the Cardington fire

tests, together with a full-scale composite floor t est a t

BREI5 and a number of small-scale tests9,10,13,16,17,18,

Bailey developed simple method for alculating the load-

carrying capacity of lightly reinforced concrete slabs a t

largedisplacements. However, thi s method ha s been

applied only to square or rectangular concrete slabs at

ambient temperature.Logically, th e nex t step is to extend this work t o com-

posite flooring systems, subject to fire, that incorporate

both the composite slab and suppor ting grillage of steel

beams, with the im of developing a design method based

on a holistic, rather than an eleme ntal, approach. This

would permit the various intera ctions etween the com-

posite slab and its supportingmembers t o be taken into

account, allowing cheaper and more innovative, site-spe-

cific, fire engineer ing olutions to be adopted. This paper

develops a simple but conservative design method for

steeVconcrete composite flooring systems based on the

work on membrane action put forward by Bailey14 and

the experimental ork undertaken as partf the Carding-

ton fire tes t programme. A companion paper19 shows how

this method can be applied to practical buildings.

The behaviour of composi te loo ring systems in

fire

At normal temperature theoadcarrying capacity f a com-

posite flooring system is estimatedby considering, n iso-

lation, the lexural strength of the composite slab and the

supporting composite beams.The strength of the compos-

ite slab isypically based on the steel deck and concrete,

with any anticrack mesh ignored. Current fire design

methods follow a simi lar approach, except tha t the heo-

retical design method for composite slabs in fire ignores

any contribution from the steel eck. This was considered

necessary following observations fromctual fires, such s

the Broadgate6 and Basingstokezo fires, which showedha t

the steel eck had debonded, owing tohe release of steam

from the concrete, during he fire. Since th e deck is

ignored, the theoretical flexural strength of the slab in

fire is based on the mesh reinforcement and the concrete.

Therefore, the struc tural omponents tha t are ssumed to

contribute to theoadcarrying capacity of the flooring sys-

tem in a fire a re the concrete, mesh reinforcement, and

supporting composite beams. n the esign methods, hese

components are reduced in strengthor the fire resistance

period considered.Observations from full-scale ire tests7 have shown tha t

the current fire design methods for composite floor sys-

tems ar e very conservative. This is due to currentdesign

being based on the load path mechanisms that are

assumed a t th e ultimate and serviceability limit states .

These assumptions are adequate for the small, vertical

displacements experienced by the flooring system at nor-

mal temperature.However, they do not consider the true

load-path mechanism experienced by the flooring system

at large, vertical displacements, which are typically expe-

rienced during a fire.

Baileyl4J5 has shown tha t lightly reinforced, square or

rectangular concrete slabs, which are vertically supported

around the perimeter nd are subjected to significant ver-

tical displacements, have the ability t o support loads inexcess of the estimatesobtained considering basic flexur-

al behaviour. This is due to th e development of membrane

forces within the slab. It is thisoadcarrying mechanism,

which is different t o that assumed at ambient tempera-

ture, th at enhances the oadcarrying capacity of the com-

posite flooring ystem above that calculated using current

fire design methods. To utilise the enhanced loadcarrying

capacity identified by Bailey, the entire floor plate of the

building must be divided into square or rectangular slab

panels, with each panel incorporating a number of unpro-

tected beams. The edges of each panel must be vertically

supported, for the dura tion f the fire, using either:

-protected beams

-unprotected beams or slabs that are outside the firecompartment area

Floor plate divided into Denotes protectedslab panels that incorporate beams (or beams

unprotected beams designed to support

the applied loadduring the fire)

Denotes unprotected beams

I

Fig '. possible slab panels that incorporateFloorplate divided into

options for dividing unprotected beams

LDenotes protectedbeams (or beamsdesigned to support

the floor plate into the applied load

slab panels Denotes unprotected beamsduring the fire)

20The Structural Engineer



P A P E R : B a i l e y I M o o r eI

-beams, within the fire compartment a rea, tha t are

designed to support he applied load for the dura tion

of the fire

Two examples are shown in Fig 1. n each example, the

fire compartment is assumed to be the whole floor plate

and the edges of each slab panelare vertically supported

using protected composite beams. Each slab panel ncor-

porates a number of unprotected composite beams. The

size of the panel depends n the stren gth f the compositeslab (which ncludes membrane action), the strengthf the

composite beams, the fire resistanceeriod, and the llow-

able vertical displacement,hich is based on the ultimate

structural collapse of the system.

The question now arises as t o whether these heated

slab panels can be considered unrestrained or restrained

against horizontalmovement. If the edge of the slab coin-

cides with the dge of a building ora service duct, it is obvi-

ously unrestrained along th at edge. However, f the slab is

in the centre of the floor plate, the reinforcement will be

continuous over it s boundaries. However, the test s at

Cardington7.8 on the two larger and more realistic sized .compartments showed tha t a large crack occurred in the

heated slab aroundhe perimeter of the fire compartment.

This resulted inhe reinforcement fracturing in this area

(Fig2). I t is believed that the rack occurred during the fire

because of large hogging moments over the partially pro-

tected beams which, together withhe membrane forces in

the heated slab, ed t o fracture of the reinforcement. This

suggests that the eated slab panelsre unrestrained dur-

ing the ire.

Simplified fire design method, including

membrane action

Considering a simply supported rectangular or square

floor slab, supportedn a grillage of steel composite beams,

the loadcarrying capacityt a particular emperature can

be calculated using he following energy equation:

Internal work doneby the

composite slab in bending

External work done by the[pe=e

floor systemfunit load l+ ....1)Internal work done by the

beam(s) in bending

External work done by the

floor systemfunit load

where e is the enhancement due to membranection in th e

composite slab. To simplify the above design equation,

catenaryaction of thesteel beams is conservatively

ignored.

For a composite floor ystem subjectedo fire, the shape

of the yieldline pattern will be dependent on the behaviour

of the supporting steel composite beams, which are con-

tinually reducing in strength. This is best explained by

considering a simple example. Consider the BRE corner

fire test7, which consisted of heating a 9.0m x 6.0m area

of slab with one secondary beam. The mode of behaviour

of the system, which is continually changing ith increas-

ing temperature, is shown in Fig 3. The change rom one

mode to the next is dependent on the stren gth f the steel

composite beam, which is continually reducing. The final

mode of behaviour, as the apacity of the composite beam

tends towardszero, is due to the .0m x 6.0m simply sup-

ported two-way spanningslabsupporting heentire

applied load.For the yieldline patterns shown in Fig 3, membrane

action can occur. It has been shown from ambient tem-

perature tests n unrestrained slabs that,nce membrane

action occurs, the shapeof the yieldline pat tern does not

change with increasing vertical displacement. However,

this is not true in a fire situation, since the supporting

steel beam,which controls the shapeof the yieldline pat-

Volume 78/No l l 6 une 2000

Fig 2. Crack through

the concrete slab

around the heated

perimeter (arrows,

from to p to bottom,

show shear stud nd

fracture of mesh

reinforcement)

Fig 3. Slab and beam

behaviour with

increase in

temperature

tern, is continually reducing in strength. Thisill result

in a change inhe mode ofbehaviour with increasing tem-

perature (Fig 3). With continual changes in the mode of

behaviour and continual changes in membrane action

(which is dependent on the mode of behaviour), it can be

seen that applying eqn (l), even to the simple example

described above, can be complicated and time-consuming.

However, by assuming tha t the dominant loadcarrying

capacity of the system is due to the composite slab, the fol-lowing assumptions can be applied t o obtain a conserva-

tive estimate.

(1)The load carried by the flexural behaviourof the gril-lage of composite beams, within the fire compartment, s

based on the lower-bound mechanism for the beam with

the highest load ratio2 (i.e. the beam th at will ‘fail’ irst in

the fire). The beams are assumed to be simply supported

Behaviour mode i)Composite slab is one-way

spanning onto he beam

Behaviour mode ii)

centre of steel compositePlastic hinge forms at

beam and a fan yieldpattern forms n the slab

Behaviour mode (iii)

composite beam reduces n

Plastic hinge nsteel

strength with increase ntemperature resulting n the

yield pattern shown

Behaviour mode iv)With increasing os s ofstrength for he compositebeam, the slab yieldline

pattern tends towardshe

slab acting without he beamlower bound patternor a

=nL forlower-bound pattern

21




TABLE 1 -Critical temperatures for unprotected steel composite beams

supporting a floor slab

LoadLevel 0.01.02.04.06.08.1.23.4.5.6 0.7

526 558 590 629 1150 11000000060 8202571Temp. (“C)

Note: Interpolation can be used

Fig 4. Assumedloaded area

supported by beams

and support a loaded area calculated assuming tha t the

slab is simply supported (i.e. typical assumptions taken for

normal design).

(2)The load supported owing t o the flexural behaviourof

the composite slab is calculated based n th e lower-bound

yieldline mechanism, assuminghat the eams have ero

resistance.

(3)The enhancement due o membrane action in the om-

posite slab (e) s based on th e lower-bound yieldline mech-

anism of the slab.

(4) The loadcarrying capacity f the composite beams an d

slab (enhanced owing to membrane action) are added

together, as shown in eqn( l ) .

By assuming a ower-bound mechanism, togetherwithmembrane action based on this mechanism, conservative

estimates will always be obtained for the loadcarrying

capacity of the composite floorslab. The amount of load

tha t the composite beams support is calculated using a

simple area model. Since this can lead to n error in the

external work done by the loads, further investigation is

needed t o determine the accuracy of the method.

Consider the simple example hown in Fig 4, consisting

of a rectangular slab and two composite beams. Using

assumption (1) the load carried by the beams (Wbeame) is

based on the lower-bound mechanism of one beam given

by:

....2)

where y defines the width of slab supportedby one beam,

and using assumption (1) y =113. The accuracy of thi s

assumption is considered by comparing eqn (1)with the

proper mechanism fo r the complete beam nd slab system,

assuming tha t a diagonalyieldline forms (modes (iii) and

(iv), Fig3).Using eqn (1) the load supported by the two

composite beams, shown in Fig4, is given by:

Internal work done

by beams -Wbeame =External work done by th e

- .... 3)x

floor s ys ted un it load

wherex defines the position of the intersection f the yield-lines, as shown in Fig 3. The value f x varies between L/2

and nL, as the beams continueo lose strength. From eqn

(3) the lowest value of Wb em e is obtained when x =L/2.

Therefore, taking this value, equating eqns (2) and (3) and

solving fo r y results in y =113, which i s the exact value

taken in the simplified approach (assumption (l)) . y

using the above procedure it can be shown, for various

2

Fig 5. Failure of BRE

membrane test

structural beam layouts, that the simplificationof assump-

tion (1) s always conservativerovided that there isore

than one beamwithin the system.For the case where there

is only one beam within the heated area (i.e. the BRE cor-

ner fire est7) the error, when calculatinghe load carried

by the beam, can be as high as 33% at the oint at which

x =L/2 (mode iii),Fig 3). However,as the eam continues

to rise in temperature the imension x reduces and the

error tends towardszero. It should also be noted that , at

the point a t which the maximum error occurs, the com-

posite beam has nominal strength, with the slab (inem-

brane action) supportingmost of the load. Therefore, the

actual error inerms of loadcarrying capacityf the whole

flooring system, which the designer is concerned with, issmall. It can also be shown tha t, if catenary action of the

steel beam s considered, the unconservatism for the case

where only one beam is within the heated system is

removed. It can therefore be argued that the ssumptions

will result in conservative estimatesor all structural sys-

tems. It is worth mentioning th at including catenary

action of the beam will influence the calculation of th e

membrane action in the composite slab. This results in a

complicated design method tha t can be used practically

only as a pecification for computer software.or this rea-son th e conservative assumption of ignoring catenary

action in the steel beams was adopted in this implified

design approach.

To calculate the loadcarrying capacity f the composite

flooring system, using eqn (1) and the assumptions dis-

cussed above, the designer needs o calculate the flexural

stren gth of the supporting beams and the membrane

strength of the composite slab. The procedure is discussed

below.

Calculation of moment capacityof composite beams n

fireTo determine the moment capacity of composite beams

during a fire, amodified form of the critical temperaturemodel given n EC 4:Part 1.24 is used. he equation inEC

4 uses an adaptation factorof 0.9 to take intoccount the

top flangeypically being at lower temperature compared

with the rest of the steel beam. However, the Code state s

tha t the quation can e used provided tha t theibs of the

composite slab are insulated above the beam. Previous

research21, albeit on protected sections,has shown that, for

composite beams with full shear onnection and ‘unfilled

voids’, the adaptation factor shoulde increased t o 1.0 fo r6Omin fire resistance.Therefore, using the same symbols

given in EC 4:Part 1.2 :

....4)

Using eqn (4), together with able 3.2 n EC 4: Part 1.2,he

critical temperatures (termed ‘limiting temperatures’ in

BS 5950Part 8) an be calculated for a given load level

(term ed ‘load ratio’ in BS 5950: Part 8), s shown in

Table 1.

The Structural Engineer



P A P E R : B a i l e y l M o o r e

k b K T

DA

lbKT0 i'

t

ICalculation of the membrane strength of a lightly

reinforced composite lab

Bailey14 has developed a theoretical design method th at

estimates theoadcarrying capacity of a lightly reinforcedconcrete slab, for a given vertical displacement, based on

the in-plane stresses (membrane ction) in the lab. These

in-plane stresses ar e governed by the mode of failure

shown in Fig 5. (A full deviation of the method is given in

refs 14 and 22.) The basic equations required to use he

design methodand develop design har ts are hown below.

The in-plane stress distribution is hown in Fig 6. The

values ofk and b t ha t define the magnitude of the in-plane

stress can be calculated by considering equilibrium and

taking moments about E , assuming fracture of th e rein-

forcement along the lineEF. These values are iven by:

4na2(1- 2n)+k =

4n2a2+1

and

1.112b =

8(A+B+C-D)

where

B=1(""I$-+ k -&[ +

. . (5)

.... 6)

Fig 6. In-plane stress

distribution

Fig 7. Enhancement

factors for composite

slab used on the

Cardington frame

0 1 2 3 4 5 6 7 8 9 1012 13 145 167 18

Displacement/effectivedepth

The loadcarrying apacity of the slab can ow be esti-

mated by considering the membrane forces and the ffect

of these forces on the bending resistanceof the slab. By

relating theffect of the membrane forces to the ieldline

load, a simple enhancement actor can be obtained. Each

of the elements1and 2 (Fig 6) re considered separate-

ly, with the contr ibution of in-plane or vertical shear,

along the yieldlines, initially ignored. The loads dete r-

mined by considering elements 1and 2 will generally beunequal, and an average value,considering the contri-

bution of she ar forces (as suggested by Hayesls), can be

calculated.

For element 1 he enhancement factor for membrane

action is given by:

Similarly, for element 2 the enhancement factor for

membrane action is given by:

2+3k2m= =L()[--1 k3 ....(8)

Mol 3+ g o dl 6(l+ 6(1+ k)2

The effect of the membrane forceon the bending

moment will result in annhancement factor forelement

1 f:

and similarly for element 2:

e2b=-=M l+-(k-1)--(K2-k+1)b p b 2 ....10)MO1 2 3

The net enhancement is obtained by combining these

effects foreach element, i.e.

el =elm +elb

e2 =e2m +e2b

Typically el and 2 are not qual. Hayes13 suggests that,

if the difference can be explained by the effects of vertical

shear or th e in-plane shear, the verall enhancement may

be shown o be given by:

The above method as been shown o give excellent sti-mates of the loadcarrying capacity of the slab, for a given

displacement, for the large-scale test conducted by BREI5

and a large number of small-scale tests conducted by other

authorsl4. The method has also been shown to give accu-

ra te resu lts, irrespective of the aspect rat io (a=L/Z) of the

slab.

Calculation of flexural s trength of composite slab

Wood9 has shown tha t, for rectangular or square simply

supported slabs,

Using eqn (12) the value of p12/m can be calculated for

various aspect ratios, leading to an assessment of the load-

caring capacity<p).The moment capacity of the slab (m)s

dependent on the position of the reinforcement, the

streng th of the materials used, and the temperature

through the cross-section. The calculation is simple and i s

shown in curren t esign Codes24 and guides23.

Volume 78 /No 7 7 6 l une 200023



P A P E R : B a i l e y / M o o r e

A B C D E F

9000000 -L 9000 -- 9000 9000W -

Fire I! compartment

v -

. I

0

0

3

2 #

1

Assumed yieldline pattern

Fig 8. Assumed simplified yieldline pattern for Cardington ests 1, 2, and 4

A B C D E F

9000 9000000000 9000

FireI compartment I 8I

4

0

3- I1

Assumed yieldline

Load ratioof beams were low,

such that they provide vertical

support during the fire test

Column ignored

for simplification

Fig9.Assumed simplified yieldline pattern for Cardington ests 3, 5, and 6

Calculation of the enhancement factor (e) owing to tensile

membrane action

The enhancement factor due to membrane action can be

calculated from first principles using theequations sum-

marised earlier. However, these equations can be repre-

sented as simple design charts, based on th e effective

depth of the reinforcement (represented by the value g,,).

For the composite floorslab used at Cardington, Fig 7

shows the enhancement due to membrane action.

Validation of the simple design method

The simple design method was compared agains t the six

fire tests onducted on the Card ington rame. The follow-

ing assumptions were taken:

(1)A yield strength of 600N/mm2was used for the A142

mesh reinforcement. In all the tests the reinforcement

remained below 4OO0C, and therefore did not significant-

ly reduce in strength.

(2) An averagemeasured concrete cube streng th of

47N/mm2 was used.

(3)The measured yield stres s of 308N/mm2 forgrade 43

steel and390N/mm2 for rade 50 steel wassed.

(4) The moment capacity of the slab, ignoring the stee ldeck, was 4.2kNm7based on a n effective depth of 51mm.

(5) The total applied load (including the self-weight) was

4.9kN/m2.

(6) Since no struc tural collapse occurred in the tests, no

limits were imposed to define fracture of the reinforce-

ment.

(7)All beams were assumed toe pin-ended.

24

Fig 10. Comparisonbetween the simple

design method and

test 1

To allow the simple design method to be compared

aga ins t th e test results the ollowing procedure was fol-

lowed. For ach tes t the oadcarrying capacity of the gril-

lage of composite beams was calculated, for a given tem-

perature.This was based on the composite beamwith the

highest load ratio2, within the heated area i.e. the beam

in thegrillage tha t fails first). If the loadcarrying capaci-

ty of the beams was lower than he applied load of

4.9kN/m2, the load that must be carried by the slab was

calculated. This allowed the enhancement factor to be cal-

culated and thus theequired displacement (Fig 7) to be

defined. This displacement for a given steel temperature

could then be compared against the test results. In ach

test a lower-bound yieldline pat tern was considered, as

shown in Figs 8 and 9, and th e comparisons between the

simplified method and the test resu lts arehown in Figs

10-15. These comparisons indicate th at th e simplified

method providesan accurate estimate f the loadcarrying

capacity of the struc ture infire. For full calculations,he

reader is irected towards ref. 22. However, there are ome

interesting test observations and comparisons with the

design method that are worth mentioning.

In the British Steel corner test (test 3, Fig 9) straingauges placed, at the centreof slab, on the surface of the

concrete in the direction of the long span recorded high val-

ues of tens ile strain. Thisorresponds to he mode of fail-

ure of a crack forming across the shor ter spant the cen-

tre of the slab, as shown in the BRE membrane test (Fig

5).Following he fire, it was found th at a crack did formn

the slab, as shown in Fig 16. This provides confidence ha t

the correct mode of behaviour is considered in the impli-

fied design method.

In the British Steel emonstration test (test 6, Fig 91,

the test results howed that thecomposite floor was up-

ported, during the ire test, by the loadbearing blockwork,

which formed the compartment wall. Unfortunately, the

extent of the support rovided by the wall cannot be deter-

mined from the tes t results. Fig 9 shows a ‘guess’ at the

behaviour and load paths of the structure, assuming that

it is supported by a large extentof the wall. However, in

the absence of tes t data, o firm conclusions can be drawn

from this comparison.

Failure criterion

The previous simple design method, and its validation

against the Cardington fire tests , did not includea check

on the ultimate failuref the system.As shown from est s

at ambient temperature, themode of failure of a simply

supported concrete slab is dueo a full depth crack form-

ing across the shorter span (Fig 5). To predict failure at

normal temperature the mechanical stra ins in the rein-

forcement need to be considered. For concrete slabs sub-jected to fire, the prediction of ultimate failure becomes

more complicated, sinceoth the mechanical and thermal

effects of the slab eed to be considered.

Test 1 (British Steel Restrained Beam)

1000 Prediction using simple

design method

aTest results

(Maximum recorded displacementand temperature)

0.00 50.00 100.00 150.0000.0050.00 300.00~ 350.00

Displacement (mm)

The Structural Engineer




Test 2 (British Steel 2-D Test)

" " V I

800-

700-h

600-

Prediction using simple

design method2

2500-

a, Test results

400-(Maximum recorded displacement

and temperature)-

300-

200

100-

ti

O ' " " , " " , " " , " " ~ " " , " "

0.00 100.0000.0000.0000.0000.00

Displacement (mm)

Fig 11. Comparison between the simple design method and test 2

Test3 British Steel Corner Test)

1200 IPrediction using simple

1000 design method

W22 600-

2

ec, 400-

ti

a,

200

0' I I I i I I I I0.0 50.0 100.050.000.0 250.0 300.050.000.0

Prediction using simpledesign method

I I I i I I I I0.0 50.0 100.050.000.0 250.0 300.050.000.0 450.0

Displacement mm)

Fig 12. Comparison between he simple design method and test 3

Test4 (BRE Corner Fire Test)

1200

Prediction using simple

design method

/ Maximum recorded displacement

Test iesults

# and steel temperature)

1

0 50 100 150 2005050

Vertical displacement(m)


Mechanical strains

If we consider th e longer span of the concrete slab and

assume that the slabdeflects in the form of a paraboliccurve, the str ain in the einforcement can be calculated

(approximately) by:

8v2&=-

3L2-413)

Eqn (13)assumes that the strain in the einforcement

is the same value long the length of the slab, whereas in

reality the strains concentrated at crack locations. Forhe

failure mode considered (where a full-depth crack forms

across the shorter span), the strain in theeinforcement

will increase significantly once the crack forms, resu lting

in fracture f the reinforcement. Predicting the strain ev-

els at which the crack forms is complex. Therefore, prag-

matic approach is proposed where a limit is defined for he

average strain in theeinforcement, based on maximum

stres s for the reinforcement of 0.5fy.This leads to a maxi-

mum allowable displacement of

..(14)

In addition, a geometrical limit of 1/30 is applied to the

above equation.

Thermal effects

Thermal effects can be beneficial o the membrane action

of concrete slabs, since the vertical displacements are

increased without an increase in mechanical strains. This

increase in vertical displacement is due partly o thermal

curva ture hroug h he beams and slab. In addition,depending on the restraint to therm alxpansion, the gril-

lage of composite beams can be in a post-buckled state24

during the fire, which will also increase the vertical dis-

placement of the structure,without significant increase in

mechanical strains.

Because of the difficulty of defining the rest raint to ther-

mal expansion, the effect of the grillage of beams being in

a post-buckled s tate is conservatively ignored. To include

the effects of thermal curvature, the temperaturedistri-

bution through the slab isssumed to be linear, allowing

the displacements to be estimated using:

a(T2-T1)Z2Vtherm

@hwhere

v is the vertical displacement

a is the coefficient of thermal expansion

T2 is the bottom temperature

T I is the op temperature

h is the depth of slab

1 is the length f shorter span of the slab

(conservative)

is the actor of safety

.... 15)

was considered necessary owing to he temperatureyp-

ically varying throughout thecompartment in a fire. The

required value isdifficult to define. However, if we consider

both the mechanical and thermal effects and carry out

comparison with hedisplacementsobtained n he

Cardington tests, a conservative factor of 2.4 is obtained,

as shown below. This is based on a value of 770°C for T2-

TI or all the tests.

Combining mechanical and thermal effects

Considering mechanical and therma l effects, the maxi-

mum deflection is given by:

a(T2- +/ O.Lfy) -L2V =

19.2hReinf't,,.,

8

but,

a(T2-Tl)lV <

19.2h +z30..(l61

The comparison between eqn (16)and themaximum ver-

tical displacements recorded in the Cardington tests is

shown in Table 2. The factor of safety in eqn (16)has been

calculated such that the estimated aximum vertical dis-

placement corresponds to four of the tests. In the BRE

corner test an d British Steel restrained beam test, eqn

Volume 78/No 7 7 6 l une 200025




Test 5 (BRE Large Compartment Test)

""" I I700-

IF

Prediction using simpledesign method

LTest results

100- (Maximum slab displacement andsteel temperatureof secondary beam)

0 ' I l I I I I I I I0 100 200 300 400 500 600 700 800 900 loo0

Displacement (mm)


Test6 (British Steel DemoTest)

Prediction using simple Idesign method

b 400 7 0

Displacement (mm)

Fig 15. Comparisonbetween the simple design method and test 6

Fi q 16. Crack at

(16)gives displacements slightly higherhan those record-

ed in the ests. However, this seems reasonable,since no

failures occurred. In addition, the above equation will

always give conservative results since

(1) he thermal curvatures calculated based n the short-

er span f the slab;

(2)any additional vertical displacements dueo the system

being in a post-buckled sta te are gnored;

(3) any contribution from the steel deck is ignored;(4)any contribution from the resistance of the steel beams

to the ultimate failure mode is ignored;

(5) the increase in ductilityf the mesh reinforcement,as

it increases in temperature,s ignored.

Furthe r research is needed t o obtain a better estimate

of the vertical displacement at which ultimate failure

occurs. Meanwhile, eqn (16)can be conservatively adopt-

ed for design purposes.

Conclusions

During the 1990sa considerable amount of workwas under-

taken to investigate the behaviour of steel framed buildings,

with composite flooringystems,subject to fi re.Observations

from this work supported the general view tha t the per-formanceof composite-steel-decking flooringystems subject

to f i e s under-utilisedin current design procedures.This is ,

due mainly to the ability of lightly reinforced concrete labs

to bridge over he fire-damaged supporting steel beams and

transfer load, using membrane action, to the undamaged

parts of the steeluilding. On the basis of these observations,

the authorshave developed a simple design method foral-

culating the erformanceof composite flooring ystems sub-

ject to fire.This method assumes that the tructural compo-

nents that contribute to the loadcafiying capacity of the

flooring system in a fire are the concrete, the mesh rein-

forcement, and the supporting steeleams.

The method uses a simple energy approacho calculate

the loadcarrying capacity of a composite flooring system.The energy of the lightly reinforced composite lab isbased

on th e yieldline approach modified to account for the

enhancement due to in-plane forces. This is added o th e

energy of the supporting composite beams, which are cal-

culated usinga modified form of the critical temperature

model given in EC 4 Part: 1.2.This approach s validated

against the resultsf six full-scale fire tests carried outn

the eight-storey steel framed uilding at Cardington. In all

but one of these comparisons the method gives accurate

predictions. In the one case where the method gives less

accurate predictions, the perimeter compartment walls

supported the composite floor during the fire. The extent

of this support to the floor was not identified inhe test,

and an estimate had toe made.

A companion paper19 shows how th e method can be

applied'to practica l composite flooring systems and com-

pares the method with the traditional approachesof pro-

viding fire resistance. n all the cases examined, the pro-

posed method. predicts behaviourhat is superior t o that

obtained using he existing design approaches.

centreGf slab acrossth e span TABLE 2 -Comparisonbetweenallowable (eqnl6)and recorded maximumdisplacement for theCardington fire tests

Test

A due to

mechanical

Maximum

test AllowableA

due to

(mm)mm)strains

AllowableAaximum

Eqn(l6) test A(mm)

L (m) 1 (m)

I 2)Britishteel 2-D test I 14.0 I 9.0 I O* I 300 I 300 I 293 I 1.02

(3) ritish Steel corner es t

0.9957 5520052.01.05)BRE €mge comDartment test

1.16 269 3112.0.04) BRE comer test

1.028 430 237 193 7.87

(6) British Steel demo. test 1.0 641311 I 333 I 644

0.0 14.6

Note: Owingto the small areafheated slab the displacement due to thermal curvature was taken asero.

The Struct ural Engineer




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2. BS 5950Struc tural use of steelwork n buildings: Part8: Code ofpract ice for fire resistant design,London, British Standards

3. Eurocode 3 Design of steel structures: Part 7.2: General rules, Struc tural fire design,ENV 1993-1 2, Brussels, European

4. Eurocode 4 Design of com posite steel and concrete struct ures: Part 7.2 General rules, Struc tural fire design,ENV 1994-1

5. BS 476 Meth od for determin ation of th e fire resistance of elements of construc tion: Part 20 ,London, British Standards

6. Structural f ire engineering investigation of Broadgate phase8 ire, Ascot, Steel Construction Institute, 1991

7. Bailey, C. C., Lennon, T., and Moore, D. B.: ‘The behaviour of full-scale steel framed buildings subjected to com-

8. Martin, D. M., Moore, D. B.: ’Introduction and background to the research programme and major fire tests at BRE

9. Wood, R.H.: Plastic and elastic design of slabs nd plates, wi th partic ular reference to reinforced con crete floor slabs,London,

10. Sawczuk, A., Winnicki, L.: ‘Plastic behaviour of simply supported reinforced concrete plates at moderately arge

11. Taylor,R.: ‘A

note on possible basis ora

new method of ultimate loaddesign of reinforced concrete slabs’,Magazine

12. Kemp, K.O.: ‘Yield of a square reinforced concrete slab on simple supports allowing for membrane forces’,he Structural

13. Hayes, B.: ’Allowing for membrane action in the plastic analysis of rectangular reinforced concrete slabs‘, Magazine

14. Bailey, C.C.: ’Membrane action of unrestrained lightly reinforced concrete slabs t large displacements’, Engineering

15. Bailey, C.C., White, D.S., and Moore, D.B.: ‘The tensile membrane action of unrestrained composite slabs simulated

16. Hayes, B., Taylor, R.: ’Load-testingRC slabs‘, The Consulting Engineer, November 1969, pp46-47

17. Taylor, R., Maher, D.R.H., Hayes, B.: ‘Effect of the arrangement of reinforcement on he behaviour of reinforced con-

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24, No. 81, December 1972, pp209-218

19. Bailey, C.C., Moore, D.M.: ‘The structural behaviour of steel frameswith composite floor labs subject to fire: Part 2:Design‘, The Struc tural Engineer, 78, No. 11, 6 J une 2000

20. ‘Fire damage structural survey report to ChurchillPlaza, Churchill Way, Basingstoke’, Amos Broome Associates plc,

confidential report, J une 1991

21. Newman, C.M., Lawson, R.M.: ‘Fire resistance of composite beams‘,echnical Report 7 09 , Ascot, The SteelConstruction

Institute, 1991

22. Bailey, C.C.: Design of steel structures with composite slabs a t the fire limit state’, final report prepared for the

Department of the Environment, Transport& the Regions and the Steel Construction Institute, Report No. 87475,

Carston, The Building Research Establishment, 2000

23. Newman, C.M.:. ‘The fire resistance of composite floorswith steel decking’ (2nded.), SCl Publication 056, scot, The

Steel Construction Institute, 1991

24. Rotter, J .M.: ‘Behaviour of highly redundant multistorey buildings under compartment fires’,dvances in SteelStructures,

ICASS’99, 7 , Proc. 2nd Int. Con f. on Advances n Steel Structures, Hong Kong, 15-17 December 1999 ,Elsevier Science

The Steel Construction Institute, 1992

Institution, 1990

Committee for Standardisation, 1995

2, Brussels, European Committee for Standardisation, 1994

Institution, 1990

partment fires’, The Structu ral Engineer, 77, No. 8, April 1999, pp1 5-21

Cardington‘, Natio nal Steel Constr uctio n Conference,London, 13-14 May 1997, pp37-64

Thames & Hudson, 1961

deflections‘, l n t . 1.Solids Structu res, 1, 1965 pp97-111

of Concrete Research, 17, No. 53, December 1965, pp1 83-1 86

Engineer, 45, No.7, J uly 1967, pp235-240

of Concrete Research, 20, No. 65, December 1968, pp205-21 2

Structures (to be published).

under fire conditions’, Engineering Structures (to be published)

New I S t r u c t E reports available now!

Structural use of glass in buildingsPrice 255 plus post & packing*, ISBN l 874266 51 4,230mm x 297mm, 168 pages, 72 colour plates,

105 line illustrations. PublishedDecember 1999.

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Volume 78/No 7 1 6 l une 2000

Composite Floors Part 1

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Transcript of Composite Floors Part 1