beam column joint 01
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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
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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,
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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.
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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
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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.
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(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,
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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
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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.
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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.
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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
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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
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