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Out-of-plane resistance of concrete masonry infilled panelsJ . L. D A W EA N D C . K . S EAH

Department of Civil Engineering, University of New Brunswick, Fredericton, N.B ., Canada E3B 5 A3Received September 2, 1988

Revised manuscript accepted May 11, 1989Nine large-scale concrete masonry infilled panels (3 .6 X 2.8 m ) were tested to destruction under uniformly distributed latera l

pressure applied in small increm ents. Load-deformation curves of the infills and the enclosing steel frame showed linear elasticbehaviour prior to first cracking . Nonlinear behaviour due to cracking and archin g action of infill was prominent in thepostcracking range. Param eters investigated experimentally included the effects of boundary supports, joint rein forcement,panel thicknesses, panel opening, and characteristics of construction. In parallel with the testing program, computer-aidedanalytical techniques were developed to predict the first crack and ultimate loads. First crack prediction was based on a finiteelement analys is for bending of thick p lates, while ultimate load prediction was based on a yield-line technique modified toaccount for the arching action of infill confined within a flexible frame. Having been verified by comparison with test results, thepostcracking analysis program was used to conduct a parametric study. It was found that infill compressive strength, paneldimensions, and frame rigidity have significant effect on ultimate loads. While central openings do not affect the ultimatestrength, they do, however, reduce postcracking ductility.Key words: masonry , infill panel, steel fram e, experimen tal, out-o f-plan e,behaviour, strength, arching , yield-line techniqu e,cracking.Neuf panneaux de grandes dim ensions avec remplissage de maqonnerie (3 ,6 X 2,8 m) ont CtC soumis, dans le cadre d'essa is dedestruction, A une pression lattr ale rCpartie uniformCment. Les courbes charge-deformation des matkriaux de remplissage et ducadre en acier rkvklent un comportement Clastique lineaire avant la premikre fissuration . Le com porte men t non linCaire dii A lafissuration et A l'effet de voiite du remplissage Ctait evident durant la pos tfissu ration . Les paramktres vCrifiCs expCrimentalementincluaient les effets des appuis de pourtour, de l'armature des joints, de 1'Cpaisseur des pannea ux, de l'ouverture des panneaux e tdes caractCristiques de construction. Des techniques d 'analyse informatisCes ont CtC ClaborCes en mim e temps que le prog ramm ed'essai afin de prCdire la premikre fissuration et les charges ultimes. La prediction de la premiere fissuration Ctait basCe surl'analyse des ClCments finis pour la flexion des plaques Cpaisses, alors que la prkdiction des charges ultimes Ctait basCe sur latechnique de la ligne d'ClasticitC modifiCe pour tenir compte de l'effet de voiite du matCriau de remplissage entour6 d'un cadreflexible. Aprks comparaison avec les rksultats d'e ssai, le programme d'analyse de la postfissuration a et6 utilisC dans le cad red'une Ctude paramCtrique. on a constate que la rksistance A la compression du matkriau de remplissage, les dimensions dupanneau et la rigidit6 du cadre avaient un effet important sur les charges ultimes. Bien que les ouvertures centrales n'affectentpas la resistance ultime, elles rkduisent cependant la ductilitk de la postfissuration.Mot s cle's :maq om erie, panneau de remplissage, cadre d'acie r, expkrimental, hors-plan, comporteme nt, resistance, effet devoiite, technique de la ligne d'elasticiti, fissuration. [Traduit par la revue]Can. I . Civ. Eng. 16,854-864 (1989)

IntroductionWhile ma sonry infilled structur al steel or reinforced concreteframes are a common form of construction, there remains ageneral lack of conclusive research and design information o nthe lateral strength and b ehaviour of the panels themselves. Forexample, the Canadian code for masonry design for buildings(Canadian Standards Association 1984) contains no specificguidelines for the design of infilled panels. A critical review ofcurrent design methods conducted by Essawy and Drysdale(1987) reveals a great international diversity in design philoso-phies, ranging from working stress design based on elasticanalysis to ultimate limit states design based on yield-lineanalysis. T he majority of experiments conducted (Haseltine and

Hodgkinson 1973; West and H o d m s o n 1976; West et al . 1979;Anderson 197 6, 1985; Drysdale and Essawy 1988) to validatepossible design methods also vary greatly. In most cases, testspecimens were fabricated with well-defined, but somewhatunrealistic, boundary conditions. In the experimental workspresented herein, all test specimens were fabricated by experi-enced mason s using comm only used techniques. It is believedthat these specimens can realistically reflect the behaviour ofpractical infilled panels.NOTE:Written discuss ion of this paper is welcomed and will bereceived by the Editor until April 30, 1990 (address inside frontcover).

Tests of clay brick infilled panels conducted by Tho(1953) show ed consid erable additional lateral capacity beyfirst cracking due to arching of an infill confined within a fraHendry (1 973) also recognized the strength enhancem ent darching action and West et al. (1973) have demonstratedimportance of frame-to-panel interface conditions on the lastrength of brickw ork panels.Theoretical treatment of the arching sh-ength of maspanels was scanty. McDowell et al . (1956) proposedanalytical method to predict the lateral strength of one-spanning brickwork beam s w ith r igid supports due to archAnderson (1 984) proposed a theory for predicting the behavof one-way spann ing unreinforced m asonry walls subjecteout-of-plane loading. Anderson's theory may also includeeffects of shrinkage, initial boundary g aps, and abutment sness. Prior to the work presented herein, there has beeanalytical method to determine cracking and ultimate capof a two-way spanning m asonry infilled panel confined withflexible steel fram e.In the investigation presented herein, nine specimens, consisting of a hollow concrete masonry block infill panelsteel frame, were subjected to gradually increasing unifnormal pressure applied to the wall surface. A finite elemelastic analysis was used to predict the response up to cracking, w hile a yield-line techn ique, modified to accounthe interaction o f infill with a flexible frame, wa s used to pre


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DAWE AND SEAH . 855

W200x46Panelinfill

///////////////////////////////////it T ~ r d a n sat 1200 mm4FIG. la.Test setup: front view.

the postcracking behaviour. Verified by comparisons with testresults, the proposed analytical technique for ultimate loadprediction was used to conduct a study of several designparameters. Based on this parametric study, expressions suit-able for design purposes are proposed.Experimental program

Test setupA typical test specimen consisted of a 3 .6 m x 2.8 m m asonrypanel laid up in a steel frame. T he steel frame mem bers shown inFig. 1a consisted of W250 x 58 columns base-supported on arigid floor beam and connected to the top to a W 200 X 46 roofbeam. The rigid floor beam was fabricated from two W 3 10 x 52

sections welded continuously along their flange tips andfastened to the laboratory strongfloor by anchor bolts a.t 1220mrn intervals. The w ebs of the floor beam and roof beam wereparallel to the plane of the panel and those of the columns wereperpendicular to it. Figure 1b shows an end profile of the testsetup. The masonry infill was loaded by inflating air bagsenclosed by a reaction panel. Supports for the reaction panelwere provided by a series of longitudinal angles and shortadjusting angles bolted to the reaction fram e. Air bag pressurewas monitored with a water manometer and infill and framedeformations were monitored w ith dial gauges, with precisionsof '0.1 kPa and '0.01 m m , respectively .Description of test specimensA typical test .panel consisted of ungrouted, unreinforcedhollow concrete blocks laid up in a steel frame in running bondpattern w ith Type S m ortar. ace-shell mortar bedding was usedthroughout. Tab le 1 sum marizes the properties of the concreteblock units used in this investigation. T ype S mortar used wasproportioned by volume using 1 part masonry cement, 1 2 partportland cement, and 44 sand. Water was added to achieve therequired workability as determined by masons. All specimenswere continuously kept moist by spraying the specimens withwater on both faces for approximately 2 min each day for aperiod of 7 days. They were then left to cure in laboratorycondition for an additional week before testing.Table 2 summ arizes the important characteristics for each ofthe nine specimens. Specim ens WE1 and WE 2 were designed toevaluate the effect of bed joint reinforcement for infilled panelssnugly fitted to the surrounding steel frame with mortar.Specimen W E3 was a dry stacked panel designed to assess thecontribu tion of tensile bond stren gth of mortar joints in the infill.Effect of panel thicknesses was d etermined by specimens WE 4and WE5, w hile specimens WE 6, W E7, and WE8 were used to

Test SpecimenReaction Panel

+ 1;


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85 6 CAN. . CIV.ENGTA BL E . Properties of concrete masonry units*

. VOL. 16, 1989TA BL E . Auxiliary test results: material properties

ParameterActual dimensions:Length (mm)Width (mm)

Height (mm)Face shell (mm)Minimum w eb thickness (mm)% solidMass density (kg/m3)Unit compressive strength (MPa)

Nominal wall thickness (mm) Orientation

*Based on 10 test samples for each block type.

TA BL E . Characteristics of test specimens--

Testspecimen Test specimen characteristicsThickness = 190 mm (190 X 190 X 390 mmunits) .Trust type joint reinforcement at alternate courses.All sides mortared to frame members.Thickness = 190 mm (190 X 19 0 X 390 mm units).Plain masonry panel.All sides mortared to frame members.Thickness = 190 mm (190 X 19 0 X 390 mm units).Dry-stack panel.All four sides of panel mortared to frame members.Thickness = 140 mm (140 X 19 0 x 39 0 mm units).Plain masonry panel.All sides of panel mortared to frame members.Thickness = 90 mm (90 X 190 X 390 mm units).Plain masonry panel.Vertical edges restrained from slipping.Thickness = 190 mm (190 X 19 0 x 390 mm units).Plain masonry panel.Vertical edges restrained from slipping.20 mm gap at roof beam to panel interface.Sam e as specimen WE 6 excep t for provision of truss typejoint reinforcement at alternate courses.Same a s specimen WE 4 except for provision of restraintsagainst slipping on four sides of panel.Thickness = 190 mm (190 X 19 0 X 390 mm units).Plain masonry panel.Central window opening (1.6 m X 1.2 m).All sides restrained from slipping.

determine the effect of boundary conditions. The effect of apanel with a central window opening was evaluated byspecimen W E9.Testing proceduresUniform pressure normal to the panel surface was appliedgradually in increme nts of 0.2 kPa prior to first cracking and 0.4kPa thereafter. At the end of each load increment, the pressurewas maintained at a constant level to allow the system tostabilize before deflection readings were taken. Recordedreadings consisted of the deflections of the infill at variouslocations and deformation of the columns at mid-height.

Normal to ParaProperty bed joint bedCompressive strengthMean (MPa)

No. of test specimensCoefficient of variation (%)Elastic modulusMean (MPa)No. of test specimensCoefficient of variation (%)Modulus of rupture - lain panelMean (MPa)No. of test specimensCoefficient of variation (%)Modu lus of rupture-Bed joint reinforcementMean (MPa)No. of test specimensCoefficient of variation (%)Poisson's ratioMean 0.21No. of test specimens 45 -Coefficient of variation (%) 22.3 -

*No experimental data available.

Additionally, crack patterns and cracking loads corresponto the appearance of each new crack were recorded.Auxiliay testsQuality control tests on mortar (Canadian Standards Asstion 1976) and concrete blocks (Canadian Standards Asstion 1985) were performed to ensure standards were Com pressive strength and elastic modulus in directions paand normal to bed joints were determined by compressionon two-unit high concrete block prisms.Flexural tensile strength norma l to bed joints was determusing a bond wrench (Brow n and Palm 1982). Test speciconsisted of two-unit high prisms with face-shell mortar ding. A truss-type joint reinfore ment w as embedded into thjoint to evaluate the effect of such inclusion.Flexural tensile strength (m odulus of rupture) parallel to joint was evaluated using two-unit-long and two-unit-highspecimens. The test specimens were tested as beams third-point loading and simple end supports. Effect of reinforcem ent was evalua ted by specim ens with truss-typereinforcement embedded in the bed joint. A summary oauxiliary test results is given in Table 3 .

Experimental resultsA typical curve of uniform pressure versus deflection oinfill central point, CP, is shown in Fig. 2. CP doesnecessarily correspond to the point of maxim um deflectioninfill load-deflection behavio ur can be decom posed intostages. Stage I is characterized by lin ear elastic behaviour to initial cracking, while in stage 11,propaga tion of initial cand development of a yield-line failure mechanism occurstage 111, arching of infill confined in the flexible steel fcauses the load to increase to a maximum above that predby standard yield-line analysis. In stage IV , the load drop


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DAWE AN D SEAH : I

00 2 0 40 60 80 100 120

D e f l e c t i o n ( mm 1FIG. 2. Typica l load-defle ction of infill.

Air -bag pressure

I!,4 P8 - Rotation

N= orth Column S = South Column

0 0-8 -C 0 C 8 -20 -10 0 10 20Deflection ( m m 1 Rotation (a 0.0001 rad. I

FIG. 3. Column deformations.due to gradual crushing of masonry at the yield lines andboundaries and, finally, ultimate collapse occurs.Load-deformation beha viour of the colum ns illustrated inFig. 3 indicates relatively small deformation prior to firstcracking. The abrupt increase in the rates of lateral deflectionand rotation occurs at the onset of first cracking. It is believedthat the infill resists app lied lateral loads by flexural action priorto first cracking, while arching action is the main load-resisting

( a ) Leeward face

crushing ofinfill

(b ) Win d ward f aceFIG. 4. Typical crack patterns.

mechanism in the postcracking range. The column load-defor-mation behaviour gives a good indication of the level of lateralin-plane thrust developed in the infill due to arching.A typical crack pattern as shown in Fig. 4 exhibits thecharacteristics of yield-line mechanisms for reinforced concreteslabs. The major crack pattern was fully formed at the end ofstage 11, while subsequent increase in loading increased thecrack width on the leeward face and the depth of crushing on


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858 CAN. J. CIV. ENG. VOL. 16, 1989the windward face. Occasionally, some secondary cracking(remote from the major yield lines) also occurred, but nopermanent load drop was associated with these. Ultimate loadwas reached when a critical limit of crushing of concrete blocks 1at the crack lines and at the interface with the confining framewas reached. R i l - First iterationI 'Elastic behaviour prior to thejrst crackingA finite element analysis using a thick plate elementincorporating flexural and transverse shear deformations (Hin-ton and Owen 1984; Cope and Clark 1984) was adapted herein.The quadrilateral element has x- and y-rotational and transversetranslational degrees of freedom at each of the four comernodes, four mid-side nodes, and one-mid-element node. Themasonry panel was treated as a voided plate, (i.e., a voidsandwiched between two face shells). Plate rigidity factorsproposed by Basu and Dawson (1978) were modified to accountfor web misalignment typical of face-shell bedded hollowconcrete blocks in running bond.

Specimen WE3 consisted of a dry-stacked infilled paneldesigned to qualitatively evaluate the effect of arching. The .-entire load-deflection curve exhibits stage I11 type behaviour 12:(Seah 1988). Its ultimate load, although lower than the other 2 R i 2panels, was nevertheless significant. 0MFirst crack load prediction 12:

5 P2 P3 '4 '5 'crApplied Pressure

FIG. 5. Seca nt method for determining critical load.

i \.Jj !1 \.-----+----. iteration: : .I I; I; ; \. .

TABLE4. First crack loadsFailure criteriaThe four failure criteria relating to masonry tensile crackingas developed by Essawy and Drysdale (1986) were used in theanalysis. These criteria correspond to (1) debonding along bedjoints, (2) simultaneous bond failure through head joints andunits, (3) a stepped failure through head and bed joints, and (4)splitting directly through the units. These proposed criteria wererearranged as functional relations of the form[ l ] F i ( u , r ) = O

First crack load (H a )Testspecimen Experiment Theory R * Failure modeWE1 2.6 2.9 1.12 Debonding of bed jointWE2 3.8 3.4 0.89 Debonding of bed jointWE3 - - - -WE4 2.4 2.2 0.92 Debonding of bed jointWE5 2.2 1.8 0.82 Debonding of bed jointWE6 5.2 4.1 0.79 Splitting through head j

where u represents a set of applied stresses, r represents and blocksstrength parameters, and i = 1, 2, 3, 4 for the four different WE7 6.8 7.2 1.06 Splitting through head joand blocksfailure modes. Each relation is satisfied identically in the critical WE 8 4.8 3.7 0.77 Splitting through headjosituation, while in a noncritical condition, a residual, Ri, and blockscorresponding to an amount by which the critical failed to be WE^ 3.2 3.4 1.06 Stepped failure throughmet, can be determined. Further details of the implementation head and bed jointsof these failure criteria are available elsewhere ( ~ e a h 988).Analytical procedureIn the analytical procedure, an arbitrary assumed appliedpressure and the weight of the panel are considered incalculating shear stresses and extreme fibre stresses normal andparallel to bed joints. A mesh size that guarantees satisfactoryconvergence is used. Once the four failure criteria are checkedand the residuals calculated, the initial applied pressure isincremented by a small amount and the stress analysis repeatedto obtain a new set of residuals for each failure mode. When twosets of residuals are obtained, the true critical load is approachediteratively using a secant method as illustrated in Fig. 5. In thisfigure, residuals Ri for each failure mode i are plotted againstapplied pressure. A solution is found when Ri s less than orequal to a prescribed minimum representing the precision of thetechnique. This is repeated for all four failure modes, and thefailure mode giving the lowest critical load is taken as thegoverning mode of failure. First crack load, mode of tensilecracking, and the location of first cracking can be predictedusing this technique.

* R = theoretical value divided by experimental value.Analysis of test resultsThe predicted first crack loads and failure modes compared with those obtained experimentally in Table 4. ratios of predicted to experimental values vary from 0.77 to with a mean of 0.93 and a coefficient of variation of 15%. Fcracking appeared as bond failure between mortar and ualong bed joints when vertical edges of the panel alongcolumn webs were not restrained against slippage. Splithrough head joints and blocks was the predominant cracmode when such slippage was prevented. Since tensile bstrength of masonry construction is highly variable and cannodetermined precisely, the accuracy of first crack load pretions similarly suffers random variability.

Postcracking analysisArching action of injll ed panelsMany researchers (Park 1964; Desayi and Kulkarni 1Breastrup 1980) have employed a yield-line technique to ana


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DAWE AND SEAH : I 859

I 1 Type 1 Strip I 1

( a ) Horizontal strips 4L-ractureFIG.7. Type I strip in a deformed configuration.

I-------: kkl fk4depth of contact str ain stress

(a ) Ultimate strain not exceeded

( b ) Ver t ica l s t r ipsFIG.6. Subdivision into horizontal and vertical strips.

reinforced concrete slabs with rigid edge restraints. In theanalytical procedure presented herein, the conventional yield-line method is modified for the analysis of voided concretemasonry panels including arching action within flexible steelframe boundaries.To facilitate computations, the panel is subdivided into aseries of horizontal and vertical strips as shown in Fig. 6 (Park1964). The error involved in this approximation can beminimized by increasing the number of strips.A typical Type I strip is shown in plan view in its deformedconfiguration in Fig. 7. In this figure, E is the compressive strainin the strip, q is the applied uniformly distributed pressure, tland t2 are outward displacements of the columns at the stripelevation, c is the depth of contact at the fracture location and atstrip ends in contact with column webs, and p, a fractionbetween 0 and 1, defines the location of the fracture line in astrip.Over regions of contact defined by c in Fig. 7, a rectangularplastic stress block as used in reinforced concrete analysis isadopted (Park 1964). The strain and stress distributions on theface shell for various strain ranges are shown in Fig. 8. Themasonrv material is conse~ativelv ssumed ineffective when

cul t =0.0015 ( from test resutts )

-----,? p e p t h of crushing

--- 4

depth of contac t st rain stre ss(b) Ul t imate stra in exceeded

FIG. 8. Stress block at region of contact for various strain ranges.(tf = face-shell thickness.)

compressive tests, was used in the analysis of all large-scalespecimens.Based on the assumed stress distribution as well as geometricand equilibrium considerations, the following relationship canbe established for a single horizontal strip (Seah 1988):[21 {Wan $1 + tan $2) + (klk2f6LIAEm)x [p cos + (1- p) cos $21)~= d(tan $1 + tan $2)

- L{1 - p(c0s $1 - cos $2) - COS $2)- (tl + t2 + G1+ G2)

loaded beyond the ultimate strain, ,I ,. An average ultimate where, as shown in Figs. 7 and 8, L is the length, d is thestrain of 0.0015, as determined experimentally from prism thickness of a strip, kl and k2 are stress block factors each equal


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C A N . J. CIV. ENG. VOL. 16. 1989

to 0.85 (Park 1964), A is the cross-sectional area of a strip, f;is the masonry compressive strength, Em s the modulus ofelasticity of the infill, and +2 are rotations of strip segments(Fig. 7) , z is the strip deflection at the fracture line, tan =z/pL, and tan +2 = z/[(l - p)L]. Also included in the aboveexpression are G Iand G 2 , he initial boundary gaps between thepanel and the frame due to shrinkage or construction errors.For a given horizontal strip, t l and t2 are dependent onthe flexibilities of the frame members and in-plane thrustsdeveloped in horizontal strips. Typical strips, i and j, within-plane thrusts, Ni and Nj respectively, are shown in Fig. 9.The total flexibility at an end of strip i can be expressed as[3] F . . = F a .+ F b + F ! .1J rl rl IJwhere superscripts a, b, and t indicate flexibility contributiondue to axial elongation of the roof beam, flexural deformationof the column, and torsional deflection of the column, respec-tively. The flexibility coefficient, Fij , is the total framedisplacement at an end of strip i due to a unit thrust in strip j.These quantities have been evaluated elsewhere (Seah 1988) forboth horizontal and vertical strips. Using these flexibilitycoefficients, t , and t2 for use in [2] for a horizontal strip i arecomputed as follows:

;-11 1

/ I I1 1 ;I I

1 1 ',: r \I I1 1 I

! I I1 1 11 : ;

where n is the number of strips and F ; and Ff are theflexibility coefficients for the left and the right side of strip irespectively. A similar procedure is followed for vertical strips.

F~ axial deformation ofN N;. 6 4 bending of colum nStr ip i twisting of columnN; N; 1 1 1+ I I 1

Solution procedureA yield-line pattern consistent with boundary conditions isfirst established. By assuming rigid plate rotation withinyield-line boundaries, the entire deflected configuration isdefined by a given lateral deflection, z, at a convenient locationon the selected yield-line pattern. An iterative technique, similarto a successive displacement technique used for solving largesystems of equations (Al-Khafaji and Tooley 1986), is used tofind the applied load for a given deflected configuration. Thesolution procedure is as follows:1. Axial thrusts for all strips and frame displacements at allstrip locations are set to zero and a value of z is assumed.2 . For strip i , the depth of contact and axial thrust arecalculated based on the assumptions in step 1.

Strip j t- F~~; I ' Nj = 1 unit

FIG. 9. Evaluation of flexibility coefficients.

beam

( a ) actual system

( b ) equivalent system

( c ) equivalent system w it h bed joint reinforcementFIG. 10. In-plane thrust developed in a strip due to arching anequivalent actions.3. Boundary displacements in all previous strip locationsto the force computed in step 2 are calculated from [4] and 4. Boundary displacements computed in step 3 change athrusts for all strips, and therefore the axial thrusts of

previous strips are updated based on the most recent tboundary displacement computed in step 3.5. Steps 3 and 4 are repeated until no significant changaxial thrust of strips occurs between two successive iteratiThe axial thrust in strip i + 1 is then computed and the procedrepeated until all strips have been considered.6. By equating the external virtual work of the apppressure to the total internal virtual work at the yield lines fostrips, the applied load is computed for the value of a assumestep 1.Internal virtual work of a typical strip can be computedconsidering the statically equivalent system of forces as shin Fig. 10. This results in


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DAWE AND SEAH :TABLE . Ultimate loads

Ultimate load (kPa)Testspecimen Experiment Theory R * RemarksWE1 22.3 21.8 0.98 Slippage at roof beam to panel interface; two-way(21.8) (0.98) arching at ultimate.WE2 19.2 19.9 1.04 Slippage at roof beam to panel interface;

(19.3) (1.01) two-way arching.WE3 7.8 - - Dry-stacked panel; two-way arching.WE4 11.2 11.7 1.04 140mm panel; slippage at column webs at low load;(10.4) (0.93) slippage at room beam to panel interface atultimate; two-way arching.WE5 7.8 8.9 1.14 90 mm panel; no slippage at frame to infill interface;(8.6) (1.10) two-way arching.WE6 10.6 10.7 1 O1 20 mm top gap; no interface slippage;(9.9) (0.93) one-way arching.WE7 14.7 13.8 0.94 Joint reinforcement; 20 mm top gap; no interface(13 O) (0.88) slippage; one-way arching.WE8 13.4 13.5 1 O1 140 mm panel; no interface slippage; slippage of(13.5) (1.01) upper bed joints at ultimate; two-way arching.WE9 17.4 17.4 1 OO 1.6 m x 1.2 m central window opening; no interface(17.4) (1.00) slippage; slippage of upper bed joints at ultimate;two-way arching.NOTE:Values enclosed within parentheses are theoretical minimum collapse loads based on failure patternsobtained by trial and error.*R = theoretical value divided by experimental value.

where A is the virtual displacement at a fracture line, M and M'are resisting moments at a boundary and a yield line, respec-tively, N is the axial thrust, p is a fraction between 0 and 1, and Lis the total length of a strip. F or a strip with joint reinforcemen t,the internal work is appro ximated as

where M, = A,Fyd, (mom ent resistance of a cross section dueto reinforcement), A, is the sectional steel area of jointreinforcement, F y is the yield strength of steel, and d , is themoment arm of the steel wire.The external work fo r a strip of width w , is given by[81 Wext = qLwsAI2where q is the external applied pressure and L is the length of astrip.The load resisted by a fractured panel a t an assumed deflection,z , can be obtained by equating[91 C int =C extwhere the sum mation is performed for all horizontal and verticalstrips. Shear forces have no net work comp onent and thereforeare neglected in [6]-[8].The yield-line deflection, z , is next incremented and thecorrespond ing applied pressure computed by the abov e proce-

..'\

Crack patte rn (Spec.W E7)ij . 1 1 . 10 0 20 LO 60 8 0 100 120

Deflection ( m m IFIG.11. Comparison of a typical theoretical and experimentalload-deflection response of infill.

dures. T his pro cess is repeated until the entire load-deflectionresponse, including an unloading portion, is obtained.Comparison with test resultsPredicted ultimate loads are compared with the correspond-ing experimental values in Table 5. Ratios of predicted toexperimen tal values vary from 0.94 to 1.14 with a mean of 1.02and a coefficient of variation of 6% .A typical comparison oftheoretical and exp erimental load-deflection curves is shown inFig. 1 1 , where applied pressure is plotted against mid-paneldeflection. Actual observed failure patterns were used in theanalysis and, in most cases, excellent correlation betweenexperimental and predicted behaviour was obtained. In caseswhere failure patterns are unkn own, patterns that yield minimumcollapse loads may be found by trial-and-error procedures.Results of such analyses are sh own in paren theses in Table 5.


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862 CAN. 1. CIV. ENG. VOL. 16, 1989Parametric studies and experimen tal observations

Effects of bed joint reinforcementSpecimens WE1 an d WE7 incorporated 190 mm thick panelswith bed joint reinforcement placed at alternate courses.Specimens WE2 an d WE6 were identical to specimens WE1an d WE7 respectively, except for the omission of bed jointreinforcement. Both specimens without joint reinforcementfailed in a sudden, explosive manner, while specimens withjoint reinforcement showed con siderable ductility for the entireloading history.Specimens WE6 an d WE7 were constructed with a 20 mmgap between roof beam and top edge on panel. Since the panelswere restrained along the other three edges, the horizontal spanwas mobilized in resisting applied pressure.Comparison of specimens WE6 an d WE7 show s that inclu-sion of joint reinforcement results in a higher first crack load.This can be attributed to the higher modulus of rupture parallelto bed joints for panels with joint reinforcement (Table 3) . Th etwo vertical edges of specimens WE1 and WE2 were notprevented from slipping at the column wed-to-infill interface,and hence bending in the vertical span caused debonding failureof bed joints at relatively low loads. Inclusion of bed jointreinforcement had little effect on first crack loads of these twospecimens.Effects of boundary supportsSlippage at panel-to-column interfaces effectively causes thevertical span to resist the applied loading. This results in higherstresses normal t o the bed joint and causes debonding failure atsubstantially low er first crack load s. Slipp age may be preventedby using panel-to-frame ties or by filling the ga ps between paneland column flanges with mortar or wooden strips. Higher firstcrack loads, effected by the latter type of construction, wereobtained for specimens WE6, WE7, and WE8.Effects of lateral edge restraint on ultimate load a re indicatedby the load-deflection curve s show n in Fig . 12 for specimenWE4. This specimen w as constructed with 140 rnm thick hollowconcrete blocks, and a 10 0 mrn gap between the panel andcolum n flanges was left unfilled. Failure mod e "a" indicated inFig. 12 was eminent in the initial stage of loading. When theleeward face of the panel was restrained from further slippageby the column flanges, failure mode "b" ensued and substantialincrease in load cap acity was obtained.Figure 13 sho ws theoretical load-deflection curves of fourpanels with different boundary cond itions. Case 1 is representedby panels supported in a nonslip mode at the top and bottom.Case 2 consists of panels supp orted along three sides and free atthe top. Cases 1 and 2 allow only one-way arching to developand the ultimate loads are therefore lower. Case 3 is representedby panels supported on three sid es, but the absence of a gap atthe top boundary allows arching action to develop in the verticalspan. Slippage at the top interface, howev er, is not restricted.This boundary resulted in high er arching strength as comparedwith the two previous cases. A panel with all four sidessupported, as shown by Case 4 of Fig. 13 , results in asubstantially higher ultimate load .Effects of panel thickness and panel aspect ratioThick panels are mo re effective in developing arching actionand, as a result, the ultimate loads are higher. However, thiseffect is diminished very rapidly with increasing panel length.Results of a parametric study of these effects are shown in Fig.14 . Panels of three different types of block enclosed in a steel

0 0 40 8 I20Def lec t ion ( m m )

FIG. 12 . Load-deflection curves for specimen WE4: effect of bodary supports.

,-_-I, .... 4.-----

', Boundary conditionsi frame

// It interface gap

I'/ 2;

I interface; slippageIII 4 crack,

I/pat t ern

I /.A .-.-.-.-.-.- -I \. 3I..---..0 0 50 100 1

Deflection ( m m )FIG.13. Effect of boundary conditions on ultimate loads.

frame, w ith a stiffness com parable to that of the test frames uherein, were used in this stu dy. Theoretical ultimate loadsplotted on the vertical axis and the correspond ing aspect raare plotted on the horizontal axis. T he ultimate strength duarching diminishes rapidly with decreasing panel thicknesand for a thin panel with a large aspect ratio, arching strengtinsignificant.Effects of frame rigidityFigure 15 shows the results of a parametric study, conducfor a 19 0 mm thick panel restrained in a steel frame alongfloor beam and both co lum ns. Ultimate loads increase parabically w ith in-plane flexural stiffness of the colum ns. H owevthe rate of increase in ultimate load decreases rapidly wincreasing frame flexural stiffness. An increase in torsiostiffness of the columns was also found to increase ultimload, but only for frames with relatively high flexural stiffnEffects of panel openingA 1.6 m X 1.2 m opening was placed at the centerspecimen WE9. The opening was covered with a stiffe


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DAWE AND SEAH : I 863

01 O 1.2 1.4 1.6 1.8 2.0Aspect Rat io ( L / H 1

FIG.14. Effect of panel aspect ratio and panel thicknesses onultimate loads. ( t = panel thickness.)

J = L o ~ = ! ~m LE.200 GPaG= 77 GPa I

Column Moment of Inertia ( 8 1 )FIG.15. Effect of frame stiffness on ultimate loads.

plywood sheathing during the test. Load acting on the closedwindow was therefore transferred to the four sides of theopening. Test results indicated no significant decrease inultimate load as a result of the opening, but postultimateductility was reduced somewhat. The postultimate ductility ofmasonary panels infilled in steel frames may be attributed to theability of central strips to resist part of the applied loa ding whencrushing of perimeter strips occurs. A central openin g in a paneleffectively destroys th e ability of central strips to resist load an d,as a result, postultimate duc tility is reduced . Panel deflections atultimate load were significantly lower than for specimenswithout openings.Design recommendations

An extensive parametric study was conducted to evaluate theeffect on ultimate load of several parameters. The moreimportant parameters identified were panel thickness, infillstrength, boundary conditions, frame rigidity, and panel dimen-sions. Based o n the results of this study, the following empiricalrelations suitable for design are proposed:

(a) For a panel supported on three sides and free at the top,[ lo ] quit = 800(fk)0.75 t 2 a / ~ 2 . 5where

(b) For a panel supported on four sides,[ l 11 quit = 800(f &)0 .75t2{a/~2 .5P/H ' .~}where

In [ lo] and [ l l ] , quit is the ultimate load (kPa), t is the panelthickness (mm), L is the panel length (mm), H is the panelheight (mm), E and G are Young 's m odulus and shear modulusof the frame members respectively, I , and Ib are moments ofinertia of columns and beam respectively, and Jc and Jb ar etorsional constants of columns and beams respectively.The above equations are applicable to hollow concrete blockpanel infills within a pinned frame. Application of theseequations to infills in moment-resisting frames may result inconservative estimates of ultimate loads. Similarly, for infillssandwiched between two adja cent panels as frequently occurs ina multi-bay infilled frame syste m, the upper limits of a and P of[lo ] and [ l l ] may be used for a conservative estimate ofcollapse loads. The upper limits of a and P were establishedbased on a results of a parametric study conducted for infilledpanels enclosed within infinitely stiff frames.It should be noted that quit is the resistance due to archingaction of the panel as it deflects. A t low loads, lateral restraintspreventing th e panel from m oving out of the frame are requiredto enable arching action to begin. Fram e members are assumed

to exhibit adeq uate ductility and strength so that failure of framemembers may not occur before failure of the infill. For infillstrengths and thicknesses comparable to those used in the testprogram presented herein, frame failure would not be a majorconcern.The above relations are applicable for ultimate load predic-tion due to arching . Ela stic analysis and failure criteria proposedby Essaway and D rysdale (1986 ) and as implemented herein canbe used to predict first crack loads.Conclusions

As a result of this investigation, the following conclusionshave been reached:1. Prior to first major cra cking, the main lateral load resistingmechanism for infilled panels is by flexural action, and in thepostcracking range it is by arching action.2. Elastic analysis combined with the failure theories ofEssawy and Drysdale as implemented herein may be used topredict first crack loads.3 . Inclusion of archin g action in yield-line analysis improv espredicted capacities over that obtained by the conventionalyield-line technique.4. Deform ations of a flexible frame should be included in theanalysis to reflect the true behaviour of a system.5. Ultim ate loads increase parabolically w ith increasing panelthickness, but de crease with increasing panel length and height.


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86 4 CAN. J. CIV. ENG. VOL. 16, 19896. Th e ef fect iveness o f central s tr ips is not as s ignificant as H E N D R Y ,. W . 1973. The lateral strength of umeinforced brickwtha t o f pe r imete r s t r ips in deve loping a r ch ing ac t ion . Thus , a Structural En gineer, 51 (2): 43-50.relat ively small central opening does not reduce the arching HINTON, . , and OWEN,D. R. J. 1984. Finite element software

s trength s ignif icantly. plates and shells. Pineridge Press Limited, Swansea, UnKingdom.AcknowledgementsT h e a u t h o rs w o u l d l i k e t o t h a n k L . E. S h a w L t d . f o rpr ovid ing f inanc ia l suppor t f o r th i s r e sea r ch and to M r . Geo r geFor sy th of the At lan t ic Masonr y Assoc ia t ion f or h i s suppor t and

e n c o u r a g e m e n t .

AL-KHAFAJI , . W . , and TOOLEY,. R. 1986. Numerical methods inengineering practice. H olt, Rinehart and Winston, Inc., New Y ork,NY. pp. 118-120.ANDERSON,. 1976. Lateral loading tests on concrete block walls.Structural Engineer, Part 2,5 4( 7) : 239-246.1984. Arching action in transverse laterally loaded masonrywall panels. Structural Engineer, 62B(1): 12 -23.1985. Test on walls subjected to uniform lateral loading andedge loading. Proceedings of the Seventh International BrickMasonry Conference, Melbourne, Australia, pp. 889-899.BASU,A. K . , ~ ~ ~ D A W S O N ,. M . 1978. Orthotopic sandwichedplates- art 1: dynam ic relaxation treatment, P art 2: analysis andapplication to multicell and voided bridge decks. Proceedings, theInstitution of Civil Engineers, Supplementary, pp. 87-1 15.BRE AST RUP , . W . 1980. Dome effect in RC slabs: rigid plasticanalysis. ASCE Journal of the Structural Division, 106(ST6):1237- 1253.BROW N,R. H. , and PALM,B. D. 1982. Flexural strength of brickmasonry using the bond wrench. Proceedings, Second NorthAmerican Masonry Conference, University of Maryland, CollegePark, M D, pp. 1.1-1.15.C A N A D I A NTANDARDSSSOCIATION.976. Mortar and grout for unitmasonry. National Standard of Canada CAN3-A179 M76 , Rexdale,Ont. 1984. Masonry design for buildings. National Standard ofCanada CAN3-S304 M84, Rexdale, Ont.1985. Standard for concrete masonry units. National Standardof Canada CAN3-A 165 M85 , Rexdale, O nt .COPE,R. J. , and CLARK , . A . 1984. Concrete slabs: Analysis anddesign. Elsevier Applied Science Publishers, New York, NY .DESAYI ,. , and KULKA RNI,. B. 1977. Load-deflection behaviour ofrestrained R/C slabs. ASCE Journal of the Structural Division,103(ST2): 405-419.DRYSDALE,. G ., and ESSAW Y, . S . 1988.Out-of-plane bending ofconcrete block walls. ASCE Journal of the Structural Division,114(ST1): 121-133.ESSAW, A . , and DRYSDA LE,. G . 1986. Micro scopic failure criterionfor masonry assemblages. Proceedings, Fourth C anadian MasonrySymposium, Fredericton, N.B., pp. 263-277.1987. Evaluation of available design methods for masonrywalls subject to out-of-plane loading. Proceedings, Fourth NorthAmerican Masonry Conference, University of California, LosAngeles, CA, pp 32.1-32.18.HASELTINE,. A., and HODGKINSON,. R . 1973. Wind effects uponbrick panels walls - esign information. Proceedings, ThirdInternational Brick Masonry Conference, Essen, Germany, pp.399-406.HASELTINE,. A., WEST,H. W . H., and TUTT, . N. 1977. Design ofwalls to resist lateral loads. Structural Engineer, Part 2, 55(10):422-430.

MCD&ELL, E. L. , MCKEE,K. E . , and SEVEN , . 1956. Archaction theory of masonry walls. ASCE Journal of the StructDivision, 82(ST2): 915.1-915 .18.PARK,R. 1964. Ultim ate strength of rectangular concrete slabs unshort-term uniform loading with edges restrained against latmovements. Proceedings, the Institution of Civil Engineers, 125-150.S E A H ,C . K. 1988. Out-of-plane behaviour of concrete masoinfilled panels. M .Sc.E. thesis, Department of Civil EngineerUniversity of New B runswick, Fredericton, N .B.T H O M A S ,. G . 1953. The strength of brickwork. Structural EnginPart 2, 31 : 35-41.WEST, H. W . H. , and HO DGKINSON,. R. 1976. The resistancbrickwork to lateral forces - review of work at BCRA. FCanadianMasonry Symp osium, The U niversity of Calgary, CalgAlta. , pp. 275-291.WEST, H. W. H. , HODGKINSON,. R. , and WEBB,W. F. 1Lateral loading test on walls with different boundary conditiProceedings, Third International Brick M asonry Conference, EsGermany, pp. 180-186. .WEST,H. W . H . , HODGKI NSON,. R., and HASELT INE,. A. 1The resistance of brickwork to lateral loading. Structural EnginPart2,55(10): 411-421.

List of symbolscross-sect ional area of a s tr ipdepth of co ntac t a t f r ac tur e l ines tr ip thicknesse l a s ti c m o d u l u s o f f r a m e m e m b e r se las tic mod ulus of masonr y in fi l lfa i lure cr i ter ia functio nf lexibi l i ty coef f ic ientmasonr y compr es s ive s t r engths h e a r m o d u l u s o f f r a m e m e m b e r sin i t ia l boundar y ga psp a n e l h e i g h tm o m e n t o f i n e rt ia o f f r a m e m e m b e r stor s iona l cons tan ts o f f r ame m ember sstress block f ac torspane l l engthr es is t ing mome ntsin-plane thrus tsul t imate la teral pressureres idual of fa i lure cr i teria functionmasonr y p ane l th icknessla te ra l boun dar y d i sp lacementsex te r na l wor kin te rna l wo r kw i d t h o f a s tr ipdef lect ion of f r ac tur e l inev i r tua l d i sp lacemen tcompr es s ive s t r ainult imate s traina f r ac tion be tw een 0 a n d 1s tr ip rotat ions

Rpv Dawe Out of Plane

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