The axial load behaviour of composite walling

J. Construct. Steel Res.Vol. 45, No. 3, pp. 353–375, 1998 1998 Elsevier Science Ltd. All rights reserved

Printed in Great Britain0143-974X/98 $19.00+ 0.00PII: S0143-974X(97)00030-8

The Axial Load Behaviour of Composite Walling

Howard Wright

Department of Civil Engineering, University of Strathclyde, 107 Rottenrow,Glasgow G4 0NG, UK

(Received 5 February 1996; revised version received 18 March 1997;accepted 25 March 1997)

ABSTRACT

Composite walling is a novel system comprising two skins of profiled steelsheeting with an infill of concrete. The axial load capacity of the walls isdependent upon the capacity of the concrete and steel and the transfer ofload between the two. Full scale tests reported in the paper show that thefull yield stress in steel and concrete is not achieved. This is thought to bedue to buckling of the thin steel component plates of the sheeting and theinability of the cross section profiled shape of the concrete to carry eccentricmoments caused by imperfections in the specimen and line of loading. Modi-fications based upon plate buckling theory for the steel and an empiricalmodel for the concrete are proposed that provide closer agreement betweenexperiment and the commonly used squash load design formula. A device tomore equally divide load between steel and concrete is also described alongwith a discussion on the efficiency of embossments in the profile. 1998Elsevier Science Ltd. All rights reserved

NOTATION

N or Npl. Rd* Axial load capacity of the wall in kNfcu or fck* Characteristic strength of concrete in N/mm2

Ac Cross-section area of concrete in the wall in mm2

As or Aa* Cross-section area of steel in the wall in mm2

py or fy* Yield stress of steel in N/mm2

a Reduction factor to account for profiling of concreteD Overall thickness of the wall in mm.p Pitch of profiles in wall in mmAcp Cross-section area of concrete in one pitch of the wall in

mm2

353

354 Howard Wright

ga* Partial safety factor for material steelgc* Partial safety factor for material concreteNcr Euler critical load of wall in kN(EI)c Longitudinal elastic bending stiffness of the wall in Nmm2

l Effective height of the wall in mmEs Young’s modulus of the steel in N/mm2

r Radius of gyration of wall section in mmIa Second moment of area of the steel profiles in mm4

Ecm Short term Young’s modulus of the concrete in N/mm2

Ic Second moment of area of the concrete in mm4

l Non-dimensional slenderness parameterf Slenderness parametera1 Buckling reduction factorx Buckling reduction factorNSd Design axial load capacity

*Refers to EC4 nomenclature

1 INTRODUCTION

Composite walling refers to a new building system consisting of two outerskins of profiled steel sheeting with an infill of concrete. Its development hascome about as an extension of the now well known composite flooring systemused world-wide. Similar systems have been used as missile- and blast-resist-ant walls, although these have normally relied upon a loose sand or stonerather than concrete infill. The composite walling described herein was orig-inally conceived for use as a core wall to stabilize steel frame building struc-tures, although it also has potential in concrete buildings, basements and blast-resistant structures.

Fig. 1 shows a schematic diagram of the wall. It can be noted that the steelsheeting will act to stabilize the building frame as soon as it is fixed, andprovides permanent formwork for the infill concrete, albeit with the assistanceof temporary whaling supports. Once the concrete has hardened axial load,lateral load and in-plane loads will be carried through both the steel andthe concrete.

As with composite flooring, the advantages of composite walling lie in thespeed and convenience of construction [1]. However, it is still necessary toestablish the structural behaviour of the system when resisting constructionand service loading.

Pilot studies on four samples of the walling system [2] have shown thatthe steel sheeting provides similar characteristics to conventional formwork for

355The axial load behaviour of composite walling

Fig. 1. Schematic diagram of composite walling.

concrete casting. It was also shown that the axial load capacity is influenced bythe bond between the steel and concrete, especially in the area close to thepoint of application of load.

This paper describes a development of the pilot study and collates infor-mation from a further nine full-scale tests on walls formed with proprietaryprofiled steel sheets supplied by various decking manufacturers. The paperalso describes the need for, and the development of, a load transfer device,and proposes necessary modifications to code-based methods of design andprovides a worked design example.

2 EXPERIMENTAL STUDY

In all, 13 full-scale walls were tested under a predominantly axial load regime.Table 1 shows the range of test specimens and the parameters in their construc-tion that were being studied. The first four tests formed a pilot study, whichhas already been described in Wright and Gallocher [2].

A purpose-made static loading frame was developed for the testing of thefour pilot walls. Capable of delivering 300 kN, this self-straining device allowsaxial load tests on walls up to 6 m high. However, the pilot tests showed thatlocal crushing failure occurred even on walls that had a nominal slenderness

356 Howard Wright

TABLE 1Test details and main parameters studied

Test no. Profiled steel sheeting Width (mm) Test feature

1 Super Holorib 150 Pilot test on untiedplain wall

2 Super Holorib 150 Test with constructionties in place

3 Super Holorib 150 Slender wall, weldedmesh connection

4 Super Holorib 150 Slender wall, meshand bar reinforcement

5 Ward Multideck 80 220 Load applied to bothsteel and concrete

6 Ward Multideck 80 220 Load applied to steelalone

7 Ward Multideck 80 220 Load applied toconcrete alone

8 Alphalok 180 Hook connectors used9 Ward Multideck 60 180 Tied hook connectors

used10 Alphalok 180 Lightweight concrete11 Alphalok 180 Grade 40 concrete12 Alphalok 180 0.05 eccentricity ratio13 Ribdeck 60 180 Concrete strain

measurements

(height to width) ratio of over 22. This indicated that local crushing wouldoccur prior to global buckling, and consequently all the subsequent tests havebeen carried out on relatively short walls (1.8 m high) with nominal slender-ness ratios of just less than 13.

2.1 Tests 5, 6 and 7

In the pilot tests the ultimate load and steel strains recorded demonstrated thatneither the steel nor the concrete was taking its full squash load capacity. Thesteel tended to debond from the concrete in the first two pilot tests, which ledto the introduction of wire mesh fabric welded to the steel sheeting at the topand bottom of the specimen, to improve the load transfer from the concreteto the steel. This had tended to shift the final failure mode from the wall headto a position just below where the wire mesh had been terminated. The de-bonding of the thin steel from the concrete was clearly a major factor in theload capacity.

The identification of load transfer between embossed thin gauge steel and


concrete in composite slabs has been the subject of much study [3,4], but thisis mainly associated with shear transfer when the two media are held togetherwith the normal reaction forces at the beam ends. In the wall there are nonormal forces and the nature of the shear bond is likely to be different.

To identify the nature of shear bond in the composite walling, three testswere devised whereby the application of the load could be restricted to oneor the other, or both materials. These tests are noted as 5, 6 and 7 in Table1, and were carried out on nominally identical specimens.

In order to ensure that the load conditions were met a special load transferdevice was developed. The loading beam at the head and the reaction beamat the base of the loading rig were drilled to accept a series of 12 threadedrods. These were welded to steel angles along the lower thirds of their length,thus providing a flat surface for spot welding. For test 5 the angles were spot-welded to the inner face of the steel sheeting, thus ensuring that load wastransferred to the steel and, by casting the concrete around the angles androads, to the concrete. In test 6 the rods and angles were spot-welded to theouter face of the steel sheeting, thus ensuring that load transferred only to thesteel. In test 7 the rods were simply cast into the concrete ensuring load trans-fer to the concrete only. The rod/angles were bolted to the spreader beamsand adjusted to ensure uniform transfer of load to each rod. Figure 2 showsthe loading rig and the arrangement of threaded bars for test 5.

2.2 Tests 8 and 9

The results of tests 5, 6 and 7, described later, were successful in showingthat load transfer between steel and concrete was critical. The detail used fortest 5 was found to provide a relatively good transfer of load from the loadingplate into both the steel and the concrete. However, the complexity of thisdetail was such that it was unlikely to be adopted in practice. A less complexdetail, shown in Fig. 3, was then developed. This comprised a hook of mildsteel that was spot-welded to the sheeting. Load from the jacks could then beapplied directly to the concrete and this would then transfer into the steel. Allremaining tests used this detail. However, it is interesting to note that for test8 a lower than expected load and a shearing failure immediately below theload head showed the need for tying the hooks on opposing sides of the walltogether. This was achieved using 5 mm mild steel wire looped around thehook of each connector to form a horizontal link. All subsequent tests weretied in this manner.

2.3 Tests 10 and 11

Specimens 10 and 11 compared lightweight aggregate concrete with relativelyhigh strength normal weight concrete.

358 Howard Wright

Fig. 2. Schematic diagram of composite wall tests, 5, 6 and 7.

Fig. 3. Detail of final load transfer device.


2.4 Test 12

This was followed by test 12 in which the axial load was applied at a nominaleccentricity of 0.05 times the width of the wall. The loading was appliedthrough stiff bearing plates onto a horizontal bar lying along the wall. Thiswas, in turn, bearing onto stiff capping plates. This is shown in Fig. 4.

2.5 Test 13

The final test, 13, was carried out on a type of profiled steel sheeting not usedin the previous tests. There were also more strain gauges attached to both thesteel and embedded in the concrete.

Various manufacturers decking types were used. This allowed several geo-metries of decking to be tested, which allowed some refinement of the designmodel, described later, to be made. Figure 5 gives the geometry of each ofthe steel sheet types used. These were cut longitudinally to provide an integernumber of pitches of the profile as shown in the figure.

Table 2 gives the steel and concrete properties associated with the tests.Cube results relate to the average of three cubes cured with the main testspecimen and crushed at the time of testing the wall. This was generally 7days after casting, with the longest period being 28 days.

Fig. 4. Detail of eccentric load application in test 12.

360 Howard Wright

Fig. 5. Details of profiled steel sheeting used in tests.

3 EXPERIMENTAL OBSERVATIONS

The pilot tests had shown that the failure was initiated by steel buckling fol-lowed by concrete crushing, at a load considerably less than the combinedsquash load of the steel and concrete. In the final pilot test (denoted test 4here) the welded mesh reinforcement had moved the failure from the head ofthe wall to a point just below the extent of the mesh. The initial series oftests 5, 6 and 7 were devised to investigate the transfer of load from concreteto steel and vice versa.

3.1 Test 5

This test was loaded to ensure identical strain in the steel and concrete for adepth of 380 mm (the height of the welded bars). Failure was initiated by the


TABLE 2Steel and concrete properties (all in N/mm2)

Test no. Profiled steel sheeting Concrete

0.2% yield Ultimate E value Cube result

1 280 350 182,000 25.92 280 350 182,000 25.73 280 350 182,000 23.74 280 350 182,000 285 361 469 190,000 30.86 361 469 190,000 32.27 361 469 190,000 318 352 443 182,000 27.79 329 411 186,000 29.210 352 443 182,000 25.411 352 443 182,000 37.812 352 443 182,000 34.3313 350 450 183,000 28.7

outer flange steel plates buckling just below the extent of the bars. The narrowouter flange plates on either edge of the wall were unsupported and formeda free edge and demonstrated a consequential low buckling load. Test speci-mens after number 6 were fabricated with cross-wall ties in the outer flanges,thus preventing outer flange buckling.

The steel buckling was followed by concrete crushing at the same level.Figure 6 shows the strains recorded down the wall to the mid height positionat the failing load. Between the top of the wall and 380 mm down the wall

Fig. 6. Strains recorded in steel and concrete down the specimen in test 5.

362 Howard Wright

the load is being carried by the transfer device which is shedding load intothe concrete. At the point where this stops (380 mm) the concrete is carryinga maximum strain of 2000 micro-strain. The steel sheeting carries very littleload in this region due to the stiffness of the transfer device. Towards thebottom of the transfer device the load transfers into the sheeting and strainslikely to cause buckling are present almost immediately. The concrete andsteel strains converge and are relatively similar by a point 900 mm downthe wall.

This indicates that the load transfer device had worked and enabled themaximum load in the steel (at its buckling stress) and loads approaching thecrushing load in the concrete to be achieved.

3.2 Test 6

In wall 6, the steel alone was loaded and failure occurred as a result of outerflange buckling. (Test 6 had been fabricated prior to testing specimen 5 andthe cross-wall ties included in subsequent tests were not present). This wasaccompanied by the steel separating from the concrete. The strains recordedin the steel suggest that the two profiles sustained a load of 581 kN. Takingthis from the total load sustained would indicate that only 330 kN was beingcarried by the concrete. The concrete strain gauges suggest a stress level of3.15 N/mm2, which would result in a concrete load capacity of 386 kN. Thetransfer of load between the steel and the concrete via the embossments wouldat first sight appear to be low.

Between the end of the transfer device and the mid-height of the wall thesteel sheeting will deform. Assuming elastic behaviour, this will be a linearreduction in deformation to zero at the mid-height. The load applied to theconcrete via the embossments could be assumed to be in proportion to theaxial deformation exhibited by the steel sheeting.

Taking an individual embossment, the maximum load that it is capable ofwithstanding may be considered as its attack area multiplied by the concretestrength. The attack area is defined as the projected area of embossment lengthacross the web times the embossment height. Between the end of the loadtransfer device and the mid-height position the deformation is reducing andthe load could be considered to reduce proportionally. For the Ward 80 profile,used in test 6, this approximate calculation of load capacity gives a total loadtransfer of 368 kN to the concrete via the embossments, which is close to thatderived from the steel strain measurements.

The strains recorded during the tests on wall 6 show that the steel wascarrying 250 N/mm2 on the outer flanges, up to 100 N/mm2 on the innerflanges and negligible stresses in the webs. The outer flange stress level islarger than the 190 N/mm2 buckling capacity and shows that the central rib


stiffener had increased the buckling capacity of the outer flange. The variationof stress across the section was similar to that found in test 13, and the patternis assumed to be similar for all the tests. A possible reason for this variationis described later in the paper.

3.3 Wall 7

For wall 7 the load was applied to the concrete. The failure was initiated byseparation of the steel from the concrete in a peeling action starting just underthe load head. This occurred well before the strain levels in the steel hadreached that likely to cause buckling. If the failure strain of concrete isassumed to be approximately 2000 micro-strain this would give rise to a defor-mation between the wall head and mid-height position of about 1.8 mm. Withthe lack of normal force holding the steel sheeting to the concrete, it is unlikelythat the embossments in the sheeting could withstand this much deformation,and separation is likely. Even if separation occurs only over a short length,the concrete would carry all the load over this length, and crushing is likelyat this position. Therefore, for the case of load being applied to the concreteonly, the steel is unlikely to contribute to the load capacity of the wall.

In practice the strains recorded in the test of wall 7 show that steel strainswere very low. It is only in the re-entrant portions of the inner flanges thatdiscernible strains could be detected. These are most probably due to theincreased chemical bond that develops around the folds in the re-entrant por-tion. The maximum of these give rise to a stress level of 100 N/mm2 acrossthe 50 mm wide strip of steel leading to the possibility of the steel sheetingbeing able to carry approximately 30 kN. This is only 1.6% of the concretecapacity and may be considered negligible.

3.4 Tests 8 and 9

The head details used in the tests 5, 6 and 7 were designed specifically toisolate the load application to either steel or concrete. The load transfer deviceused in these tests is not a practical detail, and, as already described, a reviseddetail consisting of a hook flat spot-welded to the sheeting has been used inall subsequent tests. Tests 8 and 9 were, therefore, used to establish the behav-iour of this detail. Test 8 used the hook detail alone and test 9 used the hookdetail with linking reinforcement.

For test 8, which used Alphalok profiled steel sheeting, the rib stiffener inthe outer flange is substantial, and it is possible that the full yield of the steelcould be achieved under ideal situations. However, the failure of this wallwas initiated by splitting and shearing of the concrete around the hook. It wasclear from the failure mode that benefit could be obtained by preventing the

364 Howard Wright

splitting failure by holding the two steel sheets together with wire reinforce-ment. Test 9 was therefore prepared with the wire reinforcement, and sustaineda load much closer to its theoretical load calculated later in the paper.

3.5 Tests 10 and 11

The failure mode in virtually all of the remaining tests was one of local buck-ling in the steel followed by concrete crushing close to the head of the wall.Figure 7 shows the sequence of failure that was typical of most of the speci-mens and Fig. 8 shows a photograph of a failed specimen at the stage of steelbuckling. There was no discernible difference in behaviour between test 10with the lightweight aggregate concrete and other tests. However, for test 11there was a much higher failing load, and failure occurred more suddenly andin a brittle fashion.

3.6 Test 12

Walls are often considered to be axially loaded even when there is a nominaleccentricity of load. Test 12 was devised to impose a 0.05 eccentricity to theload acting on the wall. An eccentricity of loading of 0.05 times the wallthickness is an eccentricity ratio that is often assumed to occur in practicalsituations when the load is nominally axial. The failure was similar to thatoccurring in tests 10 and 11, and little variation in strain readings betweenthe two profiled steel sheets was noted.

Fig. 7. Typical failure sequence of wall tests.


Fig. 8. Photograph of failed specimen 10.

3.7 Test 13

This test included a more detailed array of strain gauges in both the steel andthe concrete, and provided more detail on the transfer of load between theouter and inner flanges of the steel sheeting and between the concrete and thesteel. Figure 9 shows the strains recorded in the web of the profiled steelsheeting just below the transfer device and at a load stage close to failure.

Fig. 9. Strain profile in web of test 13.

366 Howard Wright

The outer flange carries a substantial load and indicates that, whereas stressesof 164 N/mm2 occur in the outer flanges, very low stresses are present in theembossed web of the profile and lower stresses occur on the inner flange. Thepresence of the embossments in the web constitutes a discontinuity in the steelplate, through which the stress cannot transfer. This is similar to the situationof a steel plate punctured with holes. The load transfer between outer flangesand unloaded flanges must occur through the embossments that span betweenthe two fold lines.

4 LOAD TRANSFER IN COMPOSITE WALLING

During the pilot study it had been found that the load capacity of the fourwalls tested was less than the theoretical capacity evaluated using the BS 8110[5] simplified short-wall method. This method assumes that the axial capacityis the sum of the steel and concrete capacity with a reduction of 10% foreccentric loading and imperfections. Without material safety factors clause3.8.4.3. of B.S.8110 can be quoted as:

N = 0.6fcuAc + 0.87Aspy (1)

The actual test capacities for all the walls are shown in Table 3 and canbe seen to be very much lower than predicted by the code. The discrepancywas originally attributed to the difficulty in maintaining bond between thesteel and concrete, and this was the stimulus for tests 5, 6 and 7. It can beseen in Table 3 that wall capacities for these three tests were again lower thanpredicted by B.S.8110 [5].

In test 6, where only the steel was loaded, the reduced capacity was thoughtto be mainly attributable to the early local buckling of the widest componentplates of the steel sheeting. It was assumed that the transfer of load from thesteel to the concrete would be very small. The profiled steel sheeting usedwas primarily intended for composite slab construction and had therefore beendesigned mainly for bending behaviour. The width of individual plates in theupper compression flanges of these profiles are kept low and therefore buck-ling is less likely. By contrast, the tension flanges and webs are generally widerand less effectively stiffened, although negative moments over the supports ofcontinuous decking do give rise to compression in these lower flanges. Evenif buckling begins in the compression flange, post-buckling reserves associatedwith membrane and some stress redistribution allows higher moments ofresistance to be achieved. In the wall the whole section is subjected to com-pression and the largest width or least stiffened plate is the weakest link. No


TABLE 3B.S.8110 and test axial load capacities

Test no. BS8110 capacity (kN) Test capacity Failure mode

Conc. Steel Total

1 1879 813 2692 2150 Concretecrushing steel

separation2 1864 813 2677 2000 Concrete

crushing steelseparation


buckling4 2031 813 2844 2100 Concrete

crushing steelbuckling


buckling6 0 992 992 581 local steel



8 2113 661 2774 1827 Concreteshear steelseparation



crushing steelbuckling

11 3089 661 3750 3000 Concretecrushing

(explosivefailure)




368 Howard Wright

redistribution between the two flanges of the same profiled sheet is possible,and failure follows almost immediately after the buckle first forms.

The stress at which elastic buckling first occurs in the widest plate may becomputed using the method proposed by Wright [6] for plates buckling againsta rigid medium. It was assumed that the longitudinal edges of each plateprovided a degree of rotational support half-way between the pinned and fixedconditions. For conventional cold-formed section design a pinned conditionis assumed. Clearly the concrete in-fill will provide significant restraint andan average condition between pinned and fixed is assumed to be conservative.It was also conservatively assumed that the longitudinal stiffeners rolled intothe plates were ineffective. For tests 5 and 6 the buckling load was evaluatedfor the free longitudinal edges of the plates of the profile that occur at theextreme edges of the wall. These plates were observed to buckle before theothers, and this was thought to precipitate collapse. In subsequent tests theedge plates of the profile were tied together with closer fixings to preventpremature buckling in this way. It can be seen in Table 4 that when the steelresistance is computed for these and the subsequent tests using this methodsubstantial reductions in the capacity ensue.

In test 6 it had been assumed that no load transfer would occur betweenthe steel and the concrete and that the steel would carry all of the load. How-ever, as noted earlier, the strains recorded in the concrete suggest that over300 kN was being transferred. The embossments in the steel were, therefore,

TABLE 4Axial load capacities for walls taking into account steel buckling and profiled concrete effects

Test Capacity (kN) Ratiotest/theory

Steel Concrete Combined Test

1 222 1785 2007 2150 1.072 222 1771 1993 2000 1.003 222 1633 1855 1950 1.054 222 1929 2151 2100 0.985 253 1432 1685 2151 1.286 253 297 550 581 1.067 0 1441 1441 1868 1.308 470 1627 2097 1827 0.879 221 1848 2069 2007 0.97

10 470 1491 1961 2100 1.0711 470 2379 2849 3000 1.0512 470 2016 2486 2802 1.1313 137 1614 1751 1700 0.97


able to transfer some load from the steel to the concrete. The theoretical valuefor the load carried by the concrete, quoted for wall 6 in Table 4, has assumeda transfer of 300 kN from the steel to the concrete. However, for wall 7 theload transfer from concrete to steel has already been shown to be negligibleand Table 4 assumes that there was no load transfer for test 6.

The overall capacity of walls 5 and 7 were still less than predicted byBS8110. This was also found to be true of many of the walls tested sub-sequently. The discrepancy between theory and test for walls 5 and 7 could,therefore, only be attributed to a reduced concrete capacity.

The reduction in theoretical capacity of 10% assumed in eqn (1) is intro-duced to account for the possible additional compressive bending stresses cre-ated during slight eccentric loading, or as a result of imperfections in the wall.These additional stresses are greatest at the extremities of the section. In thecomposite wall tests it was noted that on dismantling the failed specimenslocal crushing had occurred at the extreme edges of one side of the walls,generally just below the load transfer device.

Further consideration of this action is necessary in the case of the profiledcomposite wall. The extreme edges of the wall do not present a solid massof concrete, and the extra bending stresses must be carried on only that con-crete in the ribs of the profile. Consequently, a profiled wall of the sameoverall width as a solid wall will be less able to carry bending moments causedby eccentric load application and imperfections in the wall.

An empirical correction to the assumed concrete capacity may then bederived. The reduction in load capacity may be assumed to be directly pro-portional to the extent of void created by the profiling on the compressed edgeof the wall. A reduction factora, which may be applied to the strength ofthe concrete in the wall may then be expressed as:

a = profile voids on one face/Ac (2)

Applying this reduction to walls 5 and 7 and then to the preceding andfollowing tests showed that much closer agreement between test and theorycould be achieved. Table 4 shows the resulting comparisons. The only notableexception is for test 8, in which, as described earlier, the cross-profile tieswere omitted on the load transfer hooks. It should be noted that no differencein behaviour was noted for tests 10 and 11, thus suggesting that lightweightor higher strength concrete does not need special consideration. It is alsointeresting to note that the theory predicts well the load capacity of test 12,despite an enforced load eccentricity of 0.05 times the wall thickness.

The load capacities presented in Table 4 have been calculated taking intoaccount the detailed geometry of each profile type. All of the profiles wereproprietary and were designed for use in composite slabs. In all cases, the

370 Howard Wright

breadth of the component plates were such that significant reductions in steelstrength were assumed to occur due to buckling. There is no doubt that pur-pose-designed profile geometries with larger stiffeners and smaller platescould avoid this problem. This is also true when considering the reductionsdue to profiling of the concrete cross-section. In the case of the specimensconstructed using the Holorib profile the wall acts almost as a solid sectionwith only a small (10%) reduction for profiling. For the Ward 80 mm profilethe reduction in capacity is over 30%. The remaining profiles have a reductionvalue of about 20%.

5 DESIGN OF THE AXIAL LOAD CAPACITY OF COMPOSITEWALLS

Three methods may be adopted for the design of composite walling. The firstfollows the BS8110 procedure, whereby for stocky walls subject to mainlyaxial loading may be designed assuming a nominal reduction on the squashload. The second involves the calculation of additional moments based uponthe height to thickness ratio of the wall. This is again presented in BS8110.Finally the Eurocode for Composite Construction, EC4, may be used [7]. Inthis last method a reduction factor is evaluated based upon the slenderness ofthe wall as defined in the European strut buckling curves.

All the walls tested in this programme, apart from walls 2 and 3, were, bythe BS8110 definition, stocky, and legitimately fall within the criteria associa-ted with the simplified BS8110 method. Walls 2 and 3 had a height to thick-ness ratio of 22 and would suggest that, according to clause 3.8.3. of BS8110,an additional moment of 0.244 times the applied load should be included inthe design. Given the fact that these walls failed by concrete crushing closeto the bottom of the reinforced head section and not by lateral bending, itwould suggest that the bending stiffness is generally high, thus improving theglobal buckling resistance. The load capacity achieved was comparable withthe more stocky walls and it may be deduced that slender walls may be con-servatively designed using the appropriate method of BS8110. This aspect isbeing further explored in an experimental programme, in which both axialand bending loads are applied to composite walls. This investigation will bereported along with a comparison of stiffness between conventional reinforcedconcrete and composite walls in Wright [8].

The analysis of axial load capacity for composite walls in EC4 is morecomplex, requiring the determination of a Euler buckling load, which is usedto establish a slenderness factor, which in turn provides a capacity reductionfactor. In the determination of the Euler buckling load it is necessary to evalu-ate the stiffness of the wall in longitudinal bending. The fully composite stiff-


ness of the concrete and steel would appear most suitable, although the stiff-ness of the steel or concrete alone could be appropriate for situations in whichearly separation of the two materials occurs. The capacity factor has beencalculated for each wall assuming fully composite behaviour, and is given inTable 5. This reduction factor is less than the 10% used in the BS8110 simpli-fied method for all the samples, except the more slender walls 3 and 4. Thisshows that direct application of the EC4 method to the test samples wouldlead to overestimation of the wall capacities. It is, therefore, suggested thatthe factors proposed in the last section of the paper need to be applied whencalculating the plastic resistance of the section to EC4.

It must, however, be noted that nearly all of the samples failed by concretecrushing close to the head of the wall. Such a failure is not indicative of aglobal buckling failure upon which the EC4 method is based. It is the author’sview that for stocky composite walls the added computational effort of theEC4 method is unnecessary and not well founded. For such walls a concretecrushing failure should be assumed.

The calculation of buckling stress using Reference [4] is quite complex andfor satisfactory design may be unnecessary. Figure 10 shows a plot of stressagainst b/t ratio for the elastic buckling of plates that are in contact withconcrete. This has been derived using Reference [4], and assumes that thelongitudinal edges of the plate are half-way between being simply supportedand fixed. The vertical axis on the right of the graph shows the reduction instressb due to buckling over the yield stress of the plate (assumed in thisfigure as a nominal 280 N/mm2). Consequently, the reduction in strength of the

TABLE 5Capacity reduction factors calculated using EC4

Test no. EC4 reduction

1 0.942 0.913 0.764 0.745 0.956 0.957 0.958 0.929 0.91

10 0.9311 0.9112 0.9213 0.93

372 Howard Wright

Fig. 10. Buckling stress for plates of various b/t ratios.

steel in a profile wall may be easily read from this graph once the maximum b/tratio for the plates is calculated.

The reduction in concrete capacity calculated in the previous sectionrequires some detailed calculation of the geometry of the section. This calcu-lation will, however, be necessary in evaluating the cross-section of the con-crete, and therefore no simplification of this process is offered here. It shouldbe noted, however, that the section area need not be calculated with extremeprecision. The stiffening ribs and other details rolled into the profile generallyaccounts for less than 1% of the concrete area and may be ignored. Thereduction factor,a, that must be applied to the concrete may be expressed as:

a = 1 2 ((D*p) 2 Acp)/2Acp (3)

The factorsa and b may be applied to the calculation for ultimate axialload resistance in both the BS8110 and the EC4 design method.

6 DESIGN EXAMPLE

A composite wall is to be designed to carry an axial load of 300 kN dead and400 kN live load per metre run. The wall spans vertically between steel floorbeams at 3.0 m centres which effectively offer sway restraint. A Ward 60 mmprofiled steel sheet profile of 0.9 mm steel thickness and 280 N/mm2 yield


strength has been chosen for a 200 mm wide wall that is constructed withconcrete having a characteristic strength of 25 N/mm2.

6.1 Design according to B.S. 8110

The height to thickness ratio of the wall is 3000/200= 15This is defined as short in BS8110 clause 3.8.1.3The wall is defined as braced in BS8110 clause 3.8.1.5The area of steel per metre is established from manufacturers tables as

2276 mm2

The maximum b/t ratio is for the outer flange plate and is 132 (the inter-mediate stiffeners have been ignored)

From Fig. 7 this gives ab reduction factor of 0.22The area of concrete per metre of wall may be calculated as 135,728 mm2

From eqn (2) the profilinga reduction factor is 0.76Using clause 3.8.4.3 of BS8110 and includinga and b

N = 0.4fcuAca + 0.75fyAscb = 1136 kN

Actual factored loading is 1.4× 300 + 1.6 × 400 = 1060 kNTherefore the wall is adequate to carry this load

6.2 Design according to EC4

The area of steel per metre is established from manufacturers tables as2276 mm2

The maximum b/t ratio is for the outer flange plate and is 132 (the inter-mediate stiffeners have been ignored)

From Fig. 7 this gives ab reduction factor of 0.22The area of concrete per metre of wall may be calculated as 135,728 mm2

From eqn (2) the profilinga reduction factor is 0.76Using clause 4.8.3.4 of EC4

Npl.Rd = Aafyb/ga + 0.85Acfcka/gc = 1595kN

Using clause 4.8.3.7 of EC4

Ncr = p2(EI)e/l2 = 8762kN

where

(EI)e = Es(Aar2 + Ia) + 0.8EcmIc/gc = 799× 1010 Nmm2

374 Howard Wright


l = √(Npl.Rd/Ncr) = 0.42


f = 0.5[1 + a1(l 2 0.2) + l2] = 0.63

wherea1 is the imperfection factor from Table 5.5.1 of EC3= 0.34Using clause 5.5.1.2 of EC3

x = 1/(f + √(f 2 2 l2)) = 0.91

RecalculatingNpl.Rd to includex using clause 4.8.3.8 gives:

NSd = 0.91× 1595= 1451 kN

Actual factored loading is 1.35× 300 + 1.5 × 400 = 1005 kNTherefore the wall is adequate to carry this load.

7 CONCLUSIONS

This paper has described an experimental programme that has investigated thebehaviour and design of composite walling when subjected to axial loading.

The axial load capacity was found to be influenced by the local bucklingof the component plates in the steel sheeting and by the profiled shape of theconcrete cross-section. Reduction factors to account for these two phenom-enon have been presented. It can be seen that, in using these factors, the designof composite walling is only marginally more complex than required for con-ventional reinforced concrete.

It is believed that the longitudinal bending stiffness of composite wallingoffers substantial benefits in maintaining stability and preventing the earlyoccurrence of global buckling. Further experimental work on more slenderwalls is required to quantify this benefit.

It has also been suggested that purpose-designed profiled steel sheetingcould ensure that composite walling achieves equal structural efficiency withconventionally reinforced walls. The desirability of this course of action willdepend upon the necessity for composite walls, in practice, to carry highaxial loads.


ACKNOWLEDGEMENTS

This work has been carried out under an Engineering and Physical SciencesResearch Council contract with additional funding provided by the Steel Con-struction Institute. All of the decking manufacturers provided sheeting samplesfree of charge. The author would like to thank all of these organizations fortheir assistance.

REFERENCES

1. Wright, H. D., Evans, H. R. and Harding, P. W., The use of profiled steel sheetingin floor construction.Journal of Constructional Steel Research, 1987,7, 279–295.

2. Wright, H. D. and Gallocher, S. C., The behaviour of composite walling underconstruction and service loading.Journal of Constructional Steel Research, 1995,35, 257–273.

3. Wright, H. D. and Essawy, M. Bond in thin gauge steel concrete composite struc-tures.Engineering Foundation Conference on Composite Construction: CompositeConstruction III, Irsee, Germany, 10–14 June 1996.

4. Wright, H. D. and Velijkovic, M. Towards a numerical procedure for compositeslab analysis.13th International Speciality Conference on Cold Formed Steel, StLouis, Missouri, USA, 17–18 October 1996.

5. British Standards Institute Structural use of Concrete, Part 1. Code of practice fordesign and construction BS8110, Part 1, 1985.

6. Wright, H. D., The buckling of plates in contact with a rigid medium.The Struc-tural Engineer, 1993,71, 209–215.

7. Eurocode 4, Design of composite steel concrete structures: Part 1.1. General Rulesand Rules for Buildings, DD ENV 1994-1-1, BSI London.

8. Wright, H. D. The axial and bending behaviour of composite walling. AmericanSociety Civil Engineers, Structures Journal, in press.

Documents

The axial load behaviour of composite walling