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Paper: Cheung et al~ ~~~

Ordinary MeetingA paper to be presented and discussed at the Institution of Structural Engineers on Tuesday 25 M ay 1993 at 5 p m

Behaviour of beam-column joints in seismically-loaded RC framesP. c. Cheung, MSc, PhD, CEng, MIStruc tE , MHKIE, MASCE

Professor T.Pauiay, OBE, DTechSc, BE, PhD, FRSNZ, FIPENZ, H ~ ~ M A C IUnivers i ty of Canterbury, New Zealand

Professor R. Park, ME, PhD , FEn g, FIStructE, FICE, FRSNZ, FIPENZ, FACI, FASCEUnivers i ty of Canterbury, New Zealand

Ove Arup & Partners , Hong Kong

Pak Chi0 (Patrick) Cheung is a senior engineerwith Ove Arup & Partners, Hong Kong. Hestudied engineering at the National C heng KungUniversity in Taiwan, Cornell University in the

Zealand. He has worked for Moh & Associates,USA, and University of Canterbury in New

Scott Wilson Kirkpatrick & Partners, andAustralia’s Victoria University of Technology.

who studied in Hungary and New Zealand, is theThomas Paulay, a retired teacher and researcher,

author and coauthor of numerous technicalpapers and three books. The behaviour ofreinforced concrete buildings exposed o largeearthquakes is his major interest. He is therecipient of national and international awards

International Association for Earthquake

and honours. He is the current President of the

IntroductionIt is only since the970s that the attention f structural engineers haseen

draw n to the critica l role of beam-column joints in reinforced concrete

frames subjected to earthqu ake effect^'.^. Traditiona lly, engineers hadplaced more emphasis n the esign of beams and columns,as can be seen,

for instance, in a pape r published in The Structural Engineer in 19344.

Concern for the structural adequacyf beam-column joints,however, has

been justified as a result of repeated field observations of joint failures

in recent earthquakes’.’, an exa mple being shown in Fig 1.

Fig 1. Example of beam-column joint failures in the 1985 Mexicoearthquakes6 (courtesy of the New Zealand National Society for

Earthquake Engineering)

’ RobertPark tudiedatCanterburyUniversityCollege, New Zealand, and the University ofBristol. His reasearch work, related primarily tothe design of concrete structures or buildingsand bridges, has been published in over 200

technical papers, book chapters and two books,and these have been recognised by 16 nationaland international awards. He is currently DeputyVice-Chancellor of the Universityof Canterbury.

W-SynopsisThe behaviour of beam-column join ts is discussed in thecontext of current design procedures f o r reinforced concreteductile fram es subjected to severe earthquake motions. Asplastic hinges are expected to develop in beam s, the beam -colum n join ts m ust be capable of transferring large shearforces across the joi nt cores. The mech anisms of shearresistance of jointcores comprise a diagonal concrete strutmechanism and a truss mechanism. A considerable amou nt o fjoin t core shear reinforcement is necessary to sustain the trussmechanism i f bond failure of longitudinal bars is avoided. T he

diameter of longitudinal beam reinforce ment in joi nt coresneeds to be restricted to ensure adequate anchorage in joi ntcores. The significant differences in detailing requirements ofbeam-column joints that exist b etween various concrete designCode s led to an internationa l collabora tive research projectinvolving the testing of full-scale beam-column-slab joi ntsubassemblages under quasi-static cyclic loading. The threesubassemblages designed to New Zealand practice performedvery well.

T h e 1993 129



Paper: Cheung et al

Fig 2. Some beam and joint reinforcement arrangements fo r d uctile

reinforced concrete fram es in New Zealand

The first research paper on the eismic behaviour of beam-column joints

was published in 1967'. Intensive study in New Zealand of the prob lem s

of joints began in 197lzS9, m ainly t the Universities of Canterbury andAuckland and at Work s Central Laboratories. Th e currentoncrete design

Cod e in New Zealand" contains seismic provisions pertinent to beam-

column joints which are significantly different rom those adopted in the

United States"'12. Thedifferentdesignapproachesan ddetailing e-

quire me nts used in New Zealand and the United States haveeen discuss-

ed onvariousoccasionsl2I6. Seismic Codesdeveloped forEuropean

cond i t i~ns '~~ ' 'nclude design provisions or beam-colum n join ts with many

similarities to tho se used in New Zealand".

Construction difficulties encountered ineavily reinforced beam-colum n

joint cores are evident from Fig 2. This paper does not address the Code

provisions in detail,but examines the mportant aspects of the behaviour

of beam-column joints subjected o seismic loading and reports on recent

research progress.

Design criteriaMultistorey reinforced concrete framesn seismicareas are sually designed

and detailed for d ~ctilitJ , '~" '. The ductile esign approach is associated

with what is generally known as 'strong column eak beam' beh aviour

whereby plastic hinges are designed to form in the eams rath er th an the

columns. Beam-column joints in such a moment-resisting framesee Fig

3) are subjected to large shear and bond forces. When such a frame is

designed for sustain ed ravity loads and transient wind forces, the beam -

column jointsmay need some attention in termsf adequate strength,with

exterior joints being more vulnerable than interior joints. Induced joint

shear forces in such case are relatively modera te when compared w ith those

generated by a severe earthquake. The situationecomes mor e critical when

cyclic reversalsof earthquake actions eed to be accounted forince beam-

column joints are p ron e to extensive cracking, patterns of which can be

seen in Fig 4. During a severe earthquake, when inelastic lateral fr amedisp,lacements take placemainly as a result f plastic hinging in the beam s,

the beams usually develop their flexural strengths at the column faces.

Columns above andelow a beam -column joint shou ld preferab lyemain

elastic. Thus primary attention in esign must befocused on he capability

of each beam-column joint to transmit the necessary shear forces, both

horizontally and vertically across the inevitably cracked joint core, without

jeopardising the esired ductile response f the frame. Therefo re, the joint

should be considered as an integral part of the column.

Design criteria adopted inEurope17." and New Zealand" are intended

to ensure that the strengthf a beam -column joint core shouldot be less

than that correspondingith the developmentof the selected p lastic hinge

mechanism in the rame and tha t the capacity o fcolumn should not e

jeopardised by possibletrength degradation of theoint.These

considerations in turn influence the form ulation f performance criteriafor laboratory testing and also the interpretation of test results".

Behavioural modelsUnderhorizontalearthquakeattack, hemomentsand hear orces

generated in the eams and columns of a building frame introduce internal

stress resultants at the facesf joint cores,as llustrated in Fig 4. The stress

resultants cause both horizontal and vertical shear forceso act on the joint

cores. As a result, internal diagonalensile and compre ssive stresses, shown

130

Fig 3 . Exterior and interior beam-column joi nts in a mom ent-resisting

frame subjected to lateral forces

(a)Exterior

(b) Interior joint

Fig 4. Forces acting on beam-column join ts under seismic actions .

as & an df ,, espectively, in Fig 4, occur, which, if large enough, will lead

to diagonal cracking f the core concrete. Unless adeq uate shear resistance

is provided, eventually failureof the joint c ore m ayccur along a corner

to corner diagonal plane.

For the purpose f this paper, the behaviourf a typical interior beam -

column jointof a seismic frame (Fig 4(b)) is discussed. The principles and

conclusions can be similarly applied to an exte rior joint (Fig 4(a)). For

simplicity, axial loads on the column and beams are presently ignored.

From equilibrium con ditions, the longitudinal hear force qh cross themid -dep th of th e joint core (Fig 4(b)) is

q h = T I + C,,+ C, - V:

. ( l a )

= Tz + C,, C,, V:

whereas the vertical joint shear force qvs given by

qv= T' + C:+ C - V,,

= T' + C: + C',- V,,

. . 2a )

Inhe case of the omm onmultilayered arrangemen t of column

reinforcem ent, he derivation of theverticalstress esultants is mo re

cumbersom e. By taking into ccou nt the distances between he various stress

resultants and themember dimensions, the following approxima tion fordesign purposes is considered to be acceptable":

qv=: v. hb . . . 2b)Ih h ,

where h, an d h, are the beam and column depths, respectively.

Designing for ductili ty mplies that plastichinges are expected o form

in the eams, generally at the olumn faces. When this conditions reached,

The Structural Engineer/Volume 71/No.8/20 pril 1993



Paper: Cheung et al

W 4Fig 5. Equilibrium criteria for an interior joint core23

*Cracks

ConcreteJ /

strerses

Fig 6. Diagonal com pression field induced by shear deform ations

in a diagona lly cracked core of a b eam-column joinF 3

1 I

Y(a) Concrete strut ( DC1 ( b) Diagona l compression field ( DS )

Fig 7 . Fundamental mechanisms of shear transfer in aninterior bea m-column oin t

the tensile stresses in the longitudinal beam bars attaining forces Tl and

T , (Fig 4(b)) can be significantly higher than hat given by the

characteristic yield strength of the steel, f , . The stress in the tension steelcan reach X$,, where X, is the overstrength factor accounting for

deviations fro m characteristic yield strength and also for strain-hardening

due to cyclic inelastic strain reversals. In New Zealand” and the United

States”, X, is taken to be 1.25 in the design of ductile frames.

For building frames with a regular layout, usually the earthquake-induc ed

shear forces from thebeams at theopposite sides ofa jointcore are similar.

It may therefore be assumed that V,, z V,, =: V,, as indicated in Fig 5 .

A similar acceptable approximation is V ; = V! = Vcol. oting that C,,

-k C,, = T 2 ,eqn. (la) can then be simplified to

V;.h = Tl + T, - Vcol . .. ( lb )

The magnitude of Vcorusually ranges between 12Vo and 20 Vo of ( T I +

It has been shown23 hat the behaviour of seismic beam-column joints

can be deduced from equilibrium principles relevant to diagonally cracked,

reinforced concrete elements. Considerations of shear deform ations of joint

cores23 see Fig 6) shows that shear stresses applied to the boundaries of

a joint core can be transferred by means of a diagonal compression field

(see also Fig 7(b)). In com parison to the tensile strains in the reinforce ment,

the concrete diagonal compression strains aregenerally negligible. Hence,

as Fig 6 suggests, there is a tendency for the joint core to dilate as seismic

actions continue. This then implies that both vertical and horizontal

reinforcement passing through the joint core must become longer.

T2).

The Structural Engineer/Volume 71/No.8/20April 1993

Joint shear strengthThe resistance of shear forces in a jointcore can be based on two p ostulated

mechanism^^,^, as illustrated in Fig 7.

The strut mechanism (Fig 7(a)) transfers shear forces via a diagonal

concrete strut which sustains compression only and is assumed to be inclined

at an angle close to that of the potential corner-to-corner failure plane.

The diagonal compression force is mobilised primarily by concrete

compression forces at the tw o corners of the joint core and also by some

bond forces transmitted from the beam and column reinforcement

approximately over lengths within the shaded area of Fig 7(a). The

contribution of this mechanism, sustaining a diagonal compression force

D,,s sometimes referred to as the ‘shear carried by the concrete’. T he

notation V,, an d V,, is used to represent the horizontal and vertical joint

shear forces resisted by this mechanism, respectively.

The trussmechanism (Fig 7(b)) consists of the contribution to the shear

resistance of th e vertical and horizontal reinforce ment inside the joint core.

Horizontal a nd vertical forces transferred by bond from the beam an d

column longitud inal bars are transm itted to the core concrete mainly outside

the shaded area of Fig 7(a). In Fig 7(b) these bond forces are idealised as

uniformly distributed shear flow. Despite extensive cracking in the joint

core,adiago nal compression field with a resultant force D , can be

sustained to transmit the bond forces, if ad equate transverse forces norm al

to the boundaries are provided through he presence of the oint corereinforcemen t. The truss mechan ism generating this compression field

involves the participation of horizontal reinforcem ent (normally in the form

of joint hoops), vertical reinforcement (normally in the form of column

intermediate bars), and num erous diagonal concrete struts. The contribution

of this mechanism is often referred to as the ‘shear carried by the shear

reinforcement’. The notation V,, an d V,, is sed to represent the

horizontal and vertical shear forces resisted by this mechanism.

The two mechanisms may then be superimposed to resist the total

horizontal an d vertical joint shear forces as follows:

I / h = + ‘,h . . . .(3a)

q v = VC, + v,, . . . .(3b)

Unless the axial compression load on the colum n is large, or the beam plastic

hinges are relocated away from th e column faces, the truss mechanism may

resist themajority of Yh nd V;.,due to the large bond forces to be

transferred within the joint. Consequently, a considerable amount of

transverse horizontal and vertical reinforcement may be required in the

joint core for shear resistance. This approach has been adopted in the New

Z ea la nd ” a nd E ~ r o p e a n ’ ~ ” ~esign Codes. The otal reas of shear

reinforcement required in the horizontal (Ajh)and vertical (Aj,) directions

are then simply calculated as follows:

Total area of horizontal hoops:

A j h = ‘sh’fyh . . . . 4a )

Total area of intermediate vertical column bars:

Ajv = J‘..v/fyv . . . (4b)

where f y h nd f , are the yield strengths of the horizontal and vertical

reinforcement, respectively.To ensure adequate joint shear strength it is also necessary to prevent

diagonal com pression failure of th e concrete in the joint core. T he presence

of tensile strains in both the horizontal and vertical directions will reduce

the compressive strength of the diagonal compression struts. Premature

diagonal compression failure canbe avoided by ensuring that the nominal

joint core shear stresses do not exceed kf k , where k is of the order 0.2 to

0.25 and f L is the concrete compressive cylinder strength. It should be

noted that the value of k is significantly smaller than the corresponding

value used in gauging the contrib ution of the concrete to flexural strength.

However, current Codes’o’li”sestrict joint shear stresses to a function o f

<f k because of an assumed dependence on the tensile strength of the

concrete. Codes need to be revised in this respect.

The actions at anexterior beam-column oint are similar to those discussed

above,but heconditionsare less critical (Fig 4(a)). With theproper

anchorage of barsas suggested in Fig 8, the diagonal concrete strut

mechanism will be enhanced by the end hoo ks of the beam, bars. Hence

some reduction in requirements for horizontal joint shear reinforcement

in exterior joint is possible.

Rather than adopting a rational model such as above, current ACI design

recomm endations””2 emphasise the need for he confinement of joint

cores by transverse reinforcement and framing beams. This ACI approach

disregards the issues of controlling diagonal tension and m aintaining shear

transfer m echanisms in joint coresI4.

131



Paper: Cheung et al

hc The lesserof0.5 h c Or 10db

h i T k i ?hinge

(a ) Without beam stub

Ignore

(b) With beam stub

Fig 8. Detailing requirements o r longitudinal beam bars in exterior

beam-column joints io

The role of transverse reinforcement in providing confinement in

reinforced concrete components is to develop significant passive lateral

pressure t o compressed concrete when strain ductility demand arises. This

pressure is mobilised by dilatation o f the concre te and is transverse to theapplied external compressive load. Confinem ent so achieved may then result

in two very desirable features of inelastic response of concrete’. Firstly,

it can convert th e relatively brittle material into a ductile one. Secondly,

it may enhance compressive strength so that , for example, the loss of the

con tribu tion to esistance of spalled concrete o utside a confined core ay

be more th an com pensated for by strength increase of the concrete within

the core. While there is need for confinement of the plastic hinge regions

of columns subjected to axial compression and bending above or below

a joint of a ductile frame, there s not the same need inside joint cores since

significant inelastic compression stra ins do not arise in the concre tewithin

a joint core.

Under seismic actions, the sign of both the column and thebeam bending

momen ts usually changes inside the joint core. T his feature is illustrated

in Fig 9 in which moment patterns f or a model structure idealised by line

members and a real frame with beam and column depths being taken intoconsideration can be compared. It is evident that the m oments within the

joint core of the real structure will always be less critical. Hence the need

to confine a joint core to the same extent as an adjacent potential plastic

hinge region of a column does not appear to be justified. Of course, it

is necessary to preserve the integrity of the core con crete which is subjected

to tensile strains in several directions. D iagonal sp litting cracks developing

in the non-prismatic diagonal co ncrete strut (Fig 7(a)) have to be restricted.

Usually, a nominal am ount of reinforcement, more appropriately referred

to as basketing or containment reinforcement, is used for this purpose.

It is evident that the main need for reinforcement within joint cores is to

provide shear resistance, not confinement.

Column Icentre-Line- Beam moments Column fac e

moments

(a) Model structure (b ) Real structure

Fig 9.A comparison of moment patterns fo r m odel and real

frame substructures

*it cal

4 Predicted resistance

Fig I O . A comparison of force-displacement hysteresis loops

It is sometimes argued that the presence of beams at the four faces of

joint cores of two-way frames results in an enhan cement of the performan ce

of joint coresi131’,This is indeed the case when the seismic forces act in

the direction of one axis of the building only, and hence plastic hinges do

not for m in all beams. However, the effects of earthquake groundmotions

in various directions cannot be ignored. With the form ation of plastic hinges

at all four sides of an interior rectangular column at various stages duringan eart hqu ake , he cracking in the beams at the colum n faces will reduce

the confinemen t of the joint core. Hen ce confinement by transverse beams

cannot be relied on . Laboratory tests which have demonstrated significant

confinement from transverse beams have generally loaded the specimens

in one direction only, leaving transverse stub beams unloaded or subjected

only to simulated gravity loads.

Anchorage of beam bars within joint coresThe above m echanisms of the shear resistance of beam-column joint cores

imply that the bond stresses due to the longitudinal bars of beams and

columns passing throug h joint cores play a very importan t role in the shear

behaviour of join ts. D uctile frames during a severe earthquake will develop

plastic hinges at the ends of the beams for the moment patterns shown

in Fig 9. It is necessary to ensure th at be am bars can develop tensile stresses

(with overstrength) on one side of an interior joint core and compressivestresses on the other side simultaneously, if the plastic hinges are to be

sustained. As a result, very high bond stresses can occur which could lead

to excessive slip or bond failure of the beam bars.

In the evaluation of ea rth qua ke resistance, energy dissipation capacity

of astructure is traditiona lly associated with theshape of the force-

displacement hysteresis loops’. The solidtline loop in Fig 10 therefore

represents a more desirable hysteretic response as opposed to the dashed-

line loop with ‘pinch ing’ characteristics. How ever, recent studiesz2 uggest

that som e variations in hysteresis oop shape may not have a major influence

on the inelastic dynamic response of a structure when subjected to severe

earthq uake excitations. Tha t is, hysteresis loops showing some pinching

or stiffness degrad ation, caused by, for instance, inelastic deform ations

du e to shear and bond mechanism s, will not necessarily lead o significantly

larger inelasticdisplacements, provided that the structure has some damping

of viscous ype and is capable of some further damping by hysteretic energy

dissipation . Th e inelastic response of structures with a sho rt fun dam ental

period of vibration depends to a greater extent on hysteretic energy

dissipation. Th us the extent to which shear an d bond mechanisms should

be perm itted to p articipate in the hysteretic behaviour is still a con troversial

matter . Mo reover, it is easier to repair dam age due to inelastic flexural

deformations a t a well-detailed plastic hinge of a member a nd topreserve,

albeit at a reduced level, structural stiffness than to restore shear and bond

strength within a joint core.

Bond degradation of beam bars in joint cores can be avoided as far as

possible by limiting the radio db/h, , where db s the beam bar diameter and

h, is the column (joint) depth, by the following relationship:

where CY is a coefficient. The m ost stringent current requiremen t is stipulated

in New Zealan d”, where it is required tha t

and the cylinder compressive strength of con crete,f , is not to be less than

20 MP a. Thu s it s required that d , /h , 1/36 when grade 430 U; = 430

MPa) steel is used for beam bars. It is often difficult to satisfy this bond

132 T h e Structural E n g i n e e r / V o l u m e 71 N o . 8 / 2 0 April 1993



Paper: Cheung e t a1

4I

At interiorolumns At exterior columns

Fig 11 . Effective width of T-beam subjected to negative bendingmom ent, within which the slab flexural reinforcement is

considered as being fully effective in tension"

IN IN IN

Cycles 1and 2 Cycle 4 Cycles 5 and 6

NS loading to 0.5V, EW loading to A y NS loading to 2 A y

Cycle 3NS loading to Ay

I N IN IN

si S I st

Cycles 7and 8 Cycles 9 an d 10 Cycles 11 an d 12

Bidirectional loading NS direction to 4A y Bidirectional loading

to 2Ay or 0.02dr i f t to 4 A y or 0.02dr i f t

Fig 12. Bidirectional displacement history used for quasi-static cyclic

loading tests of the US-New Zealand-Japan-China collaborativeresearch projec t on reinforced concrete beam-column joint?4

criterion since it leads to large column sections and/or relatively smalldiameter longitudinal beam bars.

Detailing requirements for the anchorage of b eam bars in exterior joints

are suggested in Fig 8".

Current ACI Code provisions"*I2 for the anchorage of beam bars in

joint cores are much less restrictive. Test resultsI6 usually indicate the

development of inferior bond strengths and pinching response of units

detailed according to AC I equirements. Severe bond deg radation implies

that the trad itional assum ption of flexural section analysis, whereby strains

in each fibre are the same fo r both the steel and the concrete, is grossly

violated at the beam-co lumn interface. Moreo ver, the contribution of beam

bars in com pression to flexu ral resistance at the critical section may vanish

completely. However, an immediate benefit is that joint shear resistance

due to the strut mechanism (Fig 7(a)) may be significantly enhanced.

Floor slabs as tension flangesWhen cast monolithically with beams, reinforced concrete floorslabs will

act as tension flanges and thus enhanc e the flexural strength of beam s. This

enhancem ent must be aken into account if beam-colum n oints and columns

are to rem ain essentially elastic when beam p lastic hinges develop flexural

over strength^'^*' "̂^. Fig 11 shows recommended tension flange widths,

formulated in New Zealand an d based largely on engineering judgment

more than 10years ago". There has been a need to explore more fully the

mechanisms of slab contributions under seismic actions.

l

150 Section 1- 1 Section 2 - 2

(a ) Two-way interior join t specimen

Unit 20 - E

2025 l 4 4 0 5 2 H

1150 I

Section 2 -2 Section 1- 1

(b) Two-way exterior oint specimen

Fig 13 . Details of bearn-column oint test subassemblages

International collaborative research projectTo address the design problems of beam-column joints a nd differences in

Code approaches, the University ofCanterbury and theUniversity of Texas

at Austin initiated, in 1984, a collaborative research project involving

engineers from New Zealand, the United States, Japan , and China. Each

country was to undertake seismic load tests on beam-column-slab

subassemblages designed according to the Codeof the country. Guidelineson the general dimensions of the subassemblages, and on the simulated

seismic loadin g to be appl ied were established by consensus. Fig2 illustrates

the gen eralised quasi-static lateral displacement history which was adop ted

in terms of the ideal strength of thenit or theateral yield

displacement Ay, imulating bidirectional earthquak e. This international

agreement enabled meaningful comparisons of test results to be made.

Papers resulting from this collaborative research have been published in

a recent Special Publication volume by the American Concrete Institute24.

The Structural Engineer/Volume 71 /No.8/20 April 1993 133



Paper: Cheung et al

(a ) One-way interior joint specimen to ductility of p = 10

~~

(b ) Two-w ay interior join t specimen to ductility of p = 8

. I

(c ) Two-w ay exterior join t specimen to ductility of p = 13

Fig 14 . One-w ay and two-way beam-column join t subassemblages with

froorslab tested under quasi-static cyclic oading simulating severe

earthquake forces

New Zealand test results

The New Zealand tests using three full-scale isolated beam-column-slab

joint subassemblages of one-way and two-way building frames have been

reported in de ta i l e l~ewhere~~*”.ig 13 shows the dimensions of two test

units. The third unit (ID -I) represented a n interior one-way joint and was

similar to unit 2D-I (see Fig 13(a)) except that the nor th and south beamswere omitted. To avoid furth er complexity in the construction of the loading

rig, no axial forces were applied to the columns. Thus the influence of

vertical axial compression stresses on joints, co nsidered to be beneficial,

were not explored in these tests.

The primary aim of the tests was to examine the behaviour of the test

units designed according to the New Zealand Codel’. In addition, he

effects of bidirectional displacements and of t he presence of transverse

beams an d floorslabs were to be investigated. The great m ajority of beam-

column join t tests eported since 19672.12 onsisted ofplane frame

subassemblages without floorslabs.

In the tests, all units designed according to the New Zealand Codel’

perfor med very satisfactorily in terms of strength , nergy dissipation, and

ductility capacity, when subjected to the cyclic lateral load o r displacement

history shown in Fig 12. Beam plastic hinges formed at the column faces,

but the joint remained fully functional, as can be seen from Fig 14 (a),

(b) and (C) howing the specimens at the final stages of testing. In the case

of unit 2D-I, deteriorationof the bottom beam bar anch orages within the

joint core eventually occurred, leading to bar slippage through the joint

core. Tests of theother two unitswere terminated after the imposition of

very large ductility demand s when the bottom beam b ars in the plastic hingeregions had buckled.

Fig 15 shows for the three test units the m easured horizontal force v.

displacement hysteretic responses n terms of the eq uivalent interstorey drifts

(displacements). The reference ideal strengths, expressed in terms o f the

column hearorces, Vi and V$ are based on measuredmaterial

properties. Vi includes also the contribution to flexural tension of the slab

reinforcement withi? the recommen ded” effective tension flange width

(see Fig 1 ), w hile I. :allows for the full participa tion of reinforcement in

tension over the entire slab width. Both the effects of strain hardening of

the steel and participation of slab bars in tension are evident. Also indicated

are displacement ductility levels p and corresponding interstorey drifts

expressed as ercentageof the storey height. Despitea gradual and inevitabledegradation of stiffness, the hysteresis curves exhibit stable energy

dissipation. D isplacement ductility factors of at least p = 8 and interstorey

drifts of at least 3.5 070 of storey height - ell in excess of usable limits

in ductile frames- ere attaine d, while streng th degrada tion was negligible.

The circled numbers represent the progression of cyclic displacements

following the loading history outlined n Fig 12. The hysteretic responses

shown, resulting fr om judiciou s detailing of critical regions, are considered

to be close to the optimal performance attainable in reinforced concrete

frames.

The ranges of response, of particular importance to a structural engineer,

are highlighted in Fig 15(a) for unit ID-I . In the elastic range, the dr ift

at first yield displacement (i.e. at p = 1) is ab out 0.45 070 (11220) of the

storey height. It has been dem onstrated2’ that the measured lateral

stiffnessof each test ubassemblagewas considerably ess than that estim ated

by conventional analysis techniques. I n addition to the deviation of actual

material properties fro m those specified, distortions of the beam-column

joint core made a majo r contribution t o the stiffness reduction. Indeed,’

the joint panels cannot be assumed to be infinitely rigid.

Limits fo r inelastic response (Fig 15(a)) are suggested in terms of th e

maximum likely ductility demand with p = 6 and a drift of 2.7 Yo,as well

as reserve displacement capacity a t which P 4 ffects are likely to become

critical. It is evident that, with appropriate detailing of beam plastic hinges

and provision of ade quate strengths of both the joint and column, ample

reserve ductility is available.

The full details of the test results may be seen elsewhere25. he following

brief review attempts to highlight the major observations.

(1) Deformations along the diagonals of the beam-column joint core of

unit ID-I are recorded in Fig 16. Along the diagonal 51-53, the observed

relatively large an d gradually increasing tensile (positive) strains an d theconsistently small compressive (negative) strains confirm that the joint core

gradually dilated. This expansion of the joint ore canbe readily explained

with the aid of the trussmechanism in Fig 7@). Jo int core expansions were

primarily the consequence of steel tensile strains developed within the join t

core, whereas the small compressive strains resulted from the essentially

elastic response of the concrete under diagonal compression forces.

(2) Distributions of beam bar strains for unit 2D-I from ductility p = 1

t o p = 4 are shown in Fig 17. Th e beam b ars were in compression on one

side of the joint core and in tension on the other side. Tensile strains in

the beam bars a t the central part of the joint core were consistently low.

Thus the bars were well anchored, allowing high bond stresses to be

sustained in the jointcore and theplastic hinges to be spread towards the

free ends of the beams.

(3) Typical strain U istributionsof column bars are shown in Fig 18 for barsC l an d C 2of unit 2D-I. For thecorner bar C l , residual tensile strains at

levels 1 and 4 were recorded, although loading conditions were expected

to impose compressive strains. The deviation from expected strains was

believed to be caused by the actionof intersecting beam bars a nd high local

bond forces introduced by the beam bars. For the intermediate bar C2,

consistent tensile strains of significant magnitudes, although below yield

level, were found from evels 1A to 4A. Thispattern did not conform with

the sense of flexural actions as implied by the m oment patterns in Fig 9,

The Structural Engineer/V olume 71/No.8/20 pril 199334



Paper: Cheung et al

3% 1% MElastic

Interstorey d rift

Ductility reserve Ductiiity

(a) One-way interior join t specimen

Unit 2 D - I 1*l. 2 1 3% 1%

Interstorey d rift

32 0as t - West direction

240ru nbe r

160

-5 80

v 10

>

5

g

0

S

2 -80

5- 1 6 0

-240

-320

Interstorey d rift

DuctiMyr-

Proboblo Limitof stohla response ( P - A effect ) 4(b) Two-w ay interior joint specimen

Unit 2 D - EInterstorey drif t

L2=/e

120

80-vf 40

h

5b -80

-120

-160

Eost-West direction

nr

Interst-orey driftExptctnd maximurnductility demand

W t y eserveA + - w i t y reserveProbabld imi t of stoble resuonsc ( P - A effect 1

(c ) Two-way exterior oint specimen

Fig IS.Horizo ntal orc e (storey shear) v. horizontal displacement response of three beam-colum n joi nt test subassemblages

T he Structural EngineerIVolume 71/No.8/20 pril 1993 135



Paper: Cheung et Q]

W

SX

C.-em

,---SlabUnit ID-I

Ductility p - 4 + 6 * 8 * l o q

Fig 16. Variation of strains along diagonal JI-J3 of the join t panel of

unit ID-I

I : 'm m . . = = -Top D24bar

8000

60009X

. 4000

c

Unit 2D-1

North -South

beam bars

0

-1000

(a ) Top firs t layer D24 bar

South275

BottomD24 bar

-\ \ \ \

W

ISXC.-ev)

6000

4000

2000

EY

0

-1000

(b ) Bottom layer D24 bar

Fig 17. North -south beam bar strains of unit 20-1

@Ductility, p =

since the ntermediate olumnbars needed to act s vertical shear

reinforcement across the joint core. The role of those vertical bars, and

the hoops in the joi nt core, participating in the truss mechanism of joint

shear resistance, is illustrated in Fig 19 (a). At one node (see Fig 19(b )),

vertical bond forces were sustained by a diagonal concrete strut and aensile

force in the horizo ntal leg of the joint hoo p, developed by means of a 135"

hook with the tail anchored in the core concrete. Horizontal bond forces

were resisted by a similar mechanism (see Fig 19 (c)).

To develop the tensile force in a column bar due to truss action, the bar

has to be anchored in the column beyond the joint core by means of bond

forces. Idealised distribution of stress necessary to sustain this truss action,

together with flexural stresses, are shown in Fig 19(d). The resulting

combin ed stress pattern is similar o thatmeasured during the test. It explains

the significant tensile strains measured in the intermediate column b ars at

the mid-depth of the joint core. Stresses in column bars in joints due to

bidirectional loading exhibit more complex and irregular patterns and are

not discussed here.

(4) Tensile strains in the legs of the horiz ontal joint ties of unit 2D-I ar e

shown in Fig 20. As the ductility levels imposed on the subassemblage

increased, and more diagonal cracks formed n the joint core, some of the

2

,*,South

HD 28 column bars

Top beam

bars

( S l Z ( N 1

0 Uni- directional

loading

p = I ( N - S )

Bottom

b beam

,v0 1000 2000 3000 0 1000 2000 3000 UnitD- I

Tensile strainx

Fig 18. Column bar strains in the jo int core of unit 20 -1

Detail Y '

Detail X '

Force

Verticalcolumn ba r

I n

4Th

Diagonal

(a) Shear transfe r by truss D concrete stru tmechanism of bond forces

from reinforcement passing Elevation Planthrough joint core

(b) Actions at detai lX 'Compressive

Vertical column bar(with flexural

actions omitted)

cBeam

bar

concrete strut

DEquilibrating

tensile force

(c) Actions at detail Y '

Top surfaceof beam1

lBeambar

Tensile

Due toruss D& to Combined

action only flexuralctions

action only

(d) Stresses in vertical olumn bar

Fig 19. Development of tensile strains in column bars participating in

truss mechanism of joint shear resistance

joint shear resistance was redistributed from he trut to the truss

mechanisms, and the tie stresses increased as a result. T he mid-depth ties

(layers 2 an d 3) were generally subjected to larger strains than the outer

ties (layers 1 an d 4).This pattern became distorted at higher ductility levels

as bidirectional loading was imposed. Type F legs exhibited larger stra ins

than types D an d E. This may be attributed to the beam width (400mm)

being smaller than the column width (600mm), and therefore thenner tie

legs (type F) were closer to the bond forces introduced to the beam bars

than the outer tie legs (type D). The strains measured in types D an d E

legs indicated tha t all hoops efficiently participated in the truss mech anism

without undue yielding. This aspect can be related to the controlled joint

shear distortions described previously. The predominantly elastic response

of joint shear reinfo rcem ent, even at large ductilities, suggests that the

amount of this reinforcement recommended by the New Zealand Code'"

could be reduced.

136 The St ruc tu r a l Eng inee r /Vo lume 71 /No .8 /20 Apr i l 1993



Paper: Cheung e t al

Column bar

\ \ \ \ \ \ \ \ \ \ ‘ . \ J ~ ’

Beam t op bars--->

Layer 1 -v)

I

Layer 2 Y

mPIIn

Layer 4

Beambottom bars’

Note :Xindicate s bi- directiona l Loading

0 1000 2000 0 1000 2000 0 1000 2(

Tensile strain x

Bar strains at positive(North-South) ductilities

Fig 20. Joint horizontal hoo p strains of unit 20-1

Fig 21 . Recommended effective widths of tensile flanges f o r cast-in-

place floor systems

In the light of the test results, revisions to the seismic,fiv--of-the -

New Zealand C ode” for beam-column joints have bken

They are briefly summarised as follows:

(1) The effective width of a slab monolithic with beams, b e , acting as a

tension flange, m ay be taken as the lesser of (see Fig 21):

1

- ne-quarter of the beam span at each side from the beam centreline- ne half or one-quarter of the distance to an adjacent parallel beam,

at each side of the beam centreline, at interior or exterior columns,

respectively- ne column w idth at each side from the beam centreline, at exterior

columns without edge beams

(2) For practical purposes, some loss of the quality of fram e performance

can be tolerated by allowing some relaxation in anchorage requirements

for beam bars in interior beam-column joints. H owever, a number of factors

additional to those considered in the current Co de are introduced. F or a

typical two-way frame with f, ‘ = 40 MP a and grade 430 steel for beam

bars when the column is subjected to a m inimum averag e axial compressive

stress of 0.2x, the limiting value of d , / h , increases to 1/25. This is a

significant relaxation on the current 1/36 ratio.

(3 ) By allowing some bond deterioration of beam bars inside joint cores,

the strut m echanism of jo int shear resistance is enhanced, thus reducing

the need for joint shear reinforcement. The minimum magnitude of V,,

(eqn. 3(a)) isof theorder of 0. 3 y h and increases with higher axial

compression load on the columns.

The Structura l E ngineer /Volume 71/No.8/20 pril 1993

. -. .. ..

ConclusionsBeam-column joint cores can be critical regions in the design of ductile

reinforced concrete momen t-resisting frame s. Goo d detailing of beam-

column joint core regions is essential if reinforced concrete fram es subjected

to severe seismicmotions are to respond in a satisfactory manner. Thevery

large shear forces acting on joint cores need o be resisted,primarily throughthe use of horizontal and vertical shear reinforcement. The diameter of

longitudinal column and beam reinforcing bars passing through jointcores

must no t be excessive to ensure adequate ancho rage and to avo id premature

bond failure.

Three full-scale beam-column-slab joint subassemblages designed

according to the New Zealand Code” performed extremely well when

subjected t o simulated seismic loading. Th e test results conformed t o the

intentions of the Code design philosophy an d supp orted the postulated

mechanisms of joint shear resistance. Some possible relaxations in the

seismic provisions of New Zealand design Code pertinent to beam-column

joints have been proposed.

AcknowledgementsSponsorship of the New Zealand part of the recent US/NZ/Japan/Ch ina

collaborative research project on beam-column joints was generously

provided by the following organisations: the Building ResearchAssociation

of New Zealand, the Ministry of Works & Development, the University

Grants Comm ittee, the University of Canterbury, the US/NZ Cooperative

Science Programm e, Pacific Steel Ltd, and the New Zealand National

Society for Earthquake Engineering.

References1. ACI 3 18-71 Building Cod e requirements o r rein orced concrete an d

Commentary on building Code requirements o r reinforced concrete,Detroit, Michigan, American Concrete Institute, 1971

2. Park, R., Paulay, T.: Reinforced concrete structures, New York, John

- - Wiley & Sons, 19753 .

4.

5 .

6.

7 .

8.

9.

10 .

ACI-ASCE Comm ittee 352: ‘Recommendations for design of beam-

column joints in monolithic reinforced concrete structures’, Journalof the American Concrete Institute, 73,No.7, July 1976, p375

Ham ann, C . W.: ‘An introduction to the earthquake resistance of

structures’, The Structural Engineer,XII, No. 12, D ecember 1934, p492

El-Asnam, lgeria arthquake of October IO, 1980 - a

reconnaissancendngineering report, Berkeley, California,

Earthquake Engineering Research Institute, January 1983

‘The September 1985 Mexico earthquakes: Final report of the New

Zealand reconnaissance team’, Bulletin of the New Zealand NationalSociety fo r Earthquake Engineering, 21,No.1, March 1988, p3

EEFIT: The Luzon, Philippines earthquake of 16 July 1990,A field

report by Earth quak e Engineering Field Investigation Team , Lon don,

Institution of Structural Engineers, 1991

Han son, N. W., Con nor, H. W.: ‘Seismic resistance of RC beam-

column joints’, Journal of the Structural Division, American Society

of Civil Engineers, 93, No.ST5, October 1967, p533

Paulay, T ., P ark , R., and Priestley, M. J . N.: ‘Reinforced concrete

beam-column joints under seismic actions’, Journal of the American

Concrete Institute, 75 , No.11, N ovember 1978, p585

NZS3101:1982 Part I : Code of practice for the design of concretestructures an d Part 2: Commentary on thedesign of concrete

structures, Wellington, Standards Association of New Zealand, 1982

(Amendment n o.1 was issued in December 1989, mainly to take into

account new reinforcing steel grades manufactured in New Zealand)

137



Paper: Cheung et a1 Informal study groups

11 .

12 .

13 .

14 .

15 .

16 .

17 .

18 .

19 .

20 .

21 ,

22 .

23 .

24 .

25 .

26 .

ACI 3 18-89:Building Code requirementso r reinforced concretean d

Commentary on building Code requirementso r reinforced concrete,

Detroit, Michigan, American Concrete Institute, 1989

ACI-AS CE Comm ittee 352: ‘Recommendations for design of beam-

column joints in monolithic reinforced concrete structures’,Journal

of the American Concrete Inst i tute, 2, No.3, May-June 1985, p266

ATC-11: Seismic resistance of reinforced concrete shear walls and

frame joints: implicat ionsf recent research o r design engineers, Palo

Alto , Calif ornia , Applied Technology Council, 1983

Park, R., Milburn, J. R.: ‘Comparison of recent New Zealand and

United States seismic design provisions for reinforced concrete beam-

column joints and est results from four units designed according to

the New Zealand Code’, Bulletin of the Ne w Zealand National Society

fo r Ear thquake Engineer ing , 16, No.1, March 1983, p3

Paulay, T.: ‘A critique of the special provisions for seismic design

of the building C ode requirements for reinforced co ncrete ACI

318-83)’, Journal of the American Concrete Institute,3, No.2, March-

April 1986, p274

Paulay, T.: ‘Seismic behaviour of beam-colum n joints in einforced

concrete space frames’, State-of-the-Art Report in Special Theme

Session SF on Inelastic Behaviour and Modelling of Concrete

Structural Compo nents under M ultidirectional Seismic Forces, Proc.

9 th Wor ld C onference on Earthquake Engineering, Tokyo-Kyoto,1988, VIII, p557

CEB: ‘Model Code for seismic design of concrete structuresBulletin

d’Information, No. 165, ComitC Eur o-Inte rnatio nal du B eton, April

1985

Eurocode 8: ‘Structures in seismic regions- esign- art 1 general

and building’, Report EUR 12266 EN , Luxem bourg, Com mission of

the Eu ropean Communities, May 1988 ed. (d raft for comment)

Rosenblueth, E. (ed.) Design of earthquakeresistantstructures,

Lon don, Pentech Press Limited, 1980

Par k, R.: ‘Capacity design of ductile reinforced concrete building

structures for earthq uake resistance’, The Structural Engineer, 70,

No.16, 18 August 1992

Paulay, T ., Priestley, M. J . N.: Seismic designof reinforced concrete

and masonry bui ldings, New York, John Wiley & Sons, 1992

Park, R.: ‘Ductility evaluation from laboratory and analytical testing’,

State-of-the-Art Report in Special Theme Session SG on Ductility

Evaluation a nd Design of Concrete Structures and Elements, Proc.

9th World Conference on Earthquake Engineering, Tokyo-Kyoto,

1988, VIII, p605

Paulay, T. : ‘Equilibrium criteria for reinforced concrete beam-co lumn

joints’, Structural Journal, American Concrete Institute, 86, No.6,

November-December 1989, p635

ACI SP-123: ‘Design of beam-column join ts for seismic resistance’,

Special Publication of American Concrete Institute, etroit, Michigan,

1991

Cheung, P. C., Paulay, T. , and Park, R.:Seismic design of reinforced

concre te beam-column jo in t s wi th f loor s lab , Department of Civil

Engineering Research Report 91-4, University of Canterbury,

Christchurch, New Zealand, 1991Cheung, P. C. , Paulay, T. , and Park , R .: ‘Some possible revisions

to the seismic provisions of the New Zealand concrete design Code

for mo ment-resisting frames’, Proc. Pacific Conference n Earthquake

Engineering, Auckland, 1991, 2, p79

Informal study groups

The purpose of the study group scheme is to createopportunities fo r mem bers of the Institution to exchange ideasand work on deepening and developing their knowledge ofstructural engineering, thu s stimulating a greater interest inand promoting the art and science of structural engineering.

or who require furthe r information about a study groupshould write to the appropriate Convener.

Members wishing to tak e part in th e work of a study group

History of Structural EngineeringConvener: Frank Newby, MA(Cantab) , FEng, FIStructE, HonFRIBA,

27 Mayfield Avenue, Londo n W4 1 PNThe Structural Engineer, Ma rch 1973, p1 10

Model Analysis as a Design ToolConvener: F. K. Garas, PhD, CEng, FIStructE, MICE,

Taylor Woodrow Construction Ltd., Taywood House, 345Ruislip Road , South hall, Middlesex UB1 2QXThe Structural Engineer, Febru ary 1977, p63

Qualitative Analysis of Structural

BehaviourConvener: D. Johnson, BSc(Eng), PhD, CEng, FIStructE, MICE,

Department of Civil & Structural Engineering, NottinghamTrent University, Burton Street, Nottingham NG1 4BUThe Structural Engineer, November 1978, p309

The Design of Steel Portal FramesConvener: L. J . Morris , BSc(Eng), PhD, ACGI, DIC, CEng, FIStructE,

Simon Engineering Laboratories, University of Manc hester,Manchester M1 3 9 PLThe Structural Engineer, Part A, June 1983, p170

Vibration Problems in StructuresConvener: J . W. Smith , BSc(Hons), PhD, ACGI, CEng, MIStructE,

Department of Civil Eng ineerin g, University of Bristo l, BristolBS8 1TRThe Structural Engineer, Part A , June 1983, p1 70

Computing in structural engineeringConvener: A. T. Humphrey, CEng, MIStructE, MIMechE,

Analysis & Test Division, G EC Re search, Marconi ResearchCentre, West Haningfield Road, Gt. Baddow, EssexThe Structural Engineer, Ma rch 1987, p83

Advanced Composite Materials and

StructuresConvener: P. R. Head, BSc(Eng), AC GI, CEng, MIStructE, MICE,

Maunsell Structural Plastics Ltd., Maunsell House, 160Croydon Road, Beckenham , Kent BR3 4DE

The Structural Engineer, Part A, June 1987, p221

Management and Maintenance of

BridgesConvener: G. D avison, BSc, CEng, MIStructE,

c/o The Institution of Structural Engineers, 11 Upper BelgraveStreet, London SW lX 8BHStructural news, 23 Jan uary 1990, p4

Documents

Beam Column Joint 01