Six full-scale nonseismically detailed reinforced concrete (RC)interior beam-wide column and beam-wall joints with zero to highaxial compression loads were tested to investigate the seismicbehavior of the joints. Quasi-static cyclic loading, simulatingearthquake actions, was applied. The overall performance of eachtest assembly was examined in terms of lateral load capacity, drift,stiffness, energy dissipation capacity, and nominal joint shearstress. Three levels of axial compressive column loads were investigatedto determine how this variable would influence the performance of thejoints. All the specimens failed adjacent to the joint panel withgradual strength deterioration and low attainment of structuralstiffness. The low attainment of stiffness and strength was attributed tothe slip of the longitudinal bars through the joint core. The test resultsshowed that RC interior beam-wide column joints and beam-walljoints with nonseismic design and detailing attained a drift ratio of2.0% without significant strength degradation. It was thusconcluded that such joints could also possess inherent ductility foradequate response to unexpected moderate earthquakes.
INTRODUCTIONExtensive experimental research1 on ductile beam-column
joint conducted in different countries throughout pastdecades has given a better understanding of ductile jointbehavior. A relatively limited database is available in literaturefor nonseismically detailed joints, however, with respect toductile detailed joints. In addition, nonseismically detailedjoints with wall-like wide columns have rarely been studied.The only available experimental study of such joints wasconducted by Li et al.2-3 The experiments carried out by Liet al.2 involved two full-scale nonseismically detailed interiorbeam-wide column joints to investigate the seismic behaviorof the joints. The two variables studied in the test specimenswere the amount of joint transverse reinforcement and the lapsplice details for column and beam reinforcements. The columnto beam width ratio of the specimens was approximately 3. Themaximum nominal horizontal shear stress in the joint corewas determined to be 0.15fc′ based on these experimentalresults. The joint without joint horizontal transverse reinforce-ment failed at a displacement ductility factor of 2, whichcorrelates well with the model proposed by Hakuto et al.4
This suggests that for reinforced concrete (RC) interior jointswithout joint transverse reinforcements, joint shear failureoccurs around a displacement ductility factor of 2, where thejoint shear stress is between 0.11fc′ and 0.17fc′ . Due to thepresence of joint transverse reinforcements, the reinforcedjoint specimen achieved a better ductility factor of 3.
Reinforced concrete structures consisting of wall-likewide column elements are very common in regions of low-to-moderate seismicity and are the predominant structuralsystem in Singapore. The BS 81105 code used in Singapore
does not specify any provision for seismic design ordetailing of RC structures. Therefore, it is of great concernthat the strength, ductility, and energy dissipation capacity ofthese structures may be inadequate to sustain earthquake-induced loads in regions of low-to-moderate seismicity. Theneed for evaluating and improving detailing of existingstructures is obvious. An experiment consisting of six full-scale RC interior beam-wide column joints and beam-walljoints typically found in framed structures designed withnonseismic detailing in Singapore has been undertaken atNanyang Technological University, Singapore (NTU) tobetter understand the seismic behavior of such joints. Thelevel of axial compression load exerted on the columns orwalls was also studied in this paper.
RESEARCH SIGNIFICANCEThe seismic behavior of nonseismically detailed beam-
wide column and beam-wall joints and the effects of columnaxial loads on the performance of such joints have rarelybeen studied. This study provides test observations andrelated analyses of such joints. Knowledge gained from thetest results can be used to develop theoretical models forseismic evaluation.
EXPERIMENTAL STUDIESSpecimens and test setup
Six full-scale nonseismically detailed interior beam-widecolumn and beam-wall joints designed based on the BS 81105
code were constructed and tested. These specimens weretypical as-built joints abstracted from the existing buildingsin Singapore. Figure 1(a) illustrates the schematic dimensionsof Specimens C1A, C1B, and C1C. These interior beam-widecolumn joints have a column-to-beam width ratio ofapproximately 3.56. The wall-to-beam width ratio ofSpecimens C2A, C2B, and C2C is approximately 7, asshown in Fig. 1(b). The C2 series specimens had a wallcross-sectional dimension of 1600 x 300 mm (63.0 x 11.8 in.)and a beam cross-section dimension of 230 x 600 mm(9.1 x 23.6 in.), whereas the C1 series specimens had a columncross-section dimension of 820 x 280 mm (32.3 x 11.0 in.) and abeam cross-section dimension of 230 x 300 mm (9.1 x 11.8 in.).All specimens met the criterion of strong column-weak beam.Three levels of axial loading—0.0, 0.1, and 0.35 Ag fc′—were investigated for both beam-wide column joints andbeam-wall joints.
Title no. 106-S54
Seismic Behavior of Nonseismically Detailed InteriorBeam-Wide Column and Beam-Wall Connectionsby Bing Li, Tso-Chien Pan, and Cao Thanh Ngoc Tran
ACI Structural Journal/September-October 2009592
A schematic view of the loading apparatus is shown in Fig. 2.Each specimen was subjected to quasi-static reverse cyclicloads that simulated earthquake loadings. A reversed horizontalloading was applied to the top of the column using a 1000 kN(224.8 kip) capacity long-stroke actuator which wasmounted on the reaction wall. The actuator was pinned at itsend to allow rotation during the test. This loading device wasmanually operated to have better control of the load increment.The bottom of the column was pinned to a strong floor. Both
beam ends were connected to the strong floor by steel linksthat allowed rotations as well as free horizontal beam movement.Vertical movement of the beam ends was restricted to providevertical beams with support reactions. The axial compressionload was applied using three small hydraulic jacks placedbetween column top end and the bottom suffix of the steel transferbeam. Four threaded rods were each fixed at four corners aroundthe test unit to balance the applied column axial load.
MaterialsThe specimens were built with identical reinforcements
and cast with Grade 20 concrete (2.9 ksi [20 MPa]). InSingapore, the buildings are designed to sustain high axialloading of 0.35. The concrete compressive strengths ofSpecimens C1A, C1B, C1C, C2A, C2B, and C2C at thetesting days were 18.9, 18.4, 19.2, 19.0, 20.0, and 20.5 MPa(2.7, 2.7, 2.8, 2.8, 2.9, and 3.0 ksi), respectively. Due to thelimited capacities in the hydraulic jacks, such low-strengthconcrete was used to achieve the targeted axial loadings.High-strength deformed reinforcement bars Y10, Y13, Y20,Y22, Y25, and Y28 were used as main bars in the test units,whereas mild steel bar R10 was used as stirrups. The propertiesof reinforcing bars are shown in Table 1.
ACI member Bing Li is an Associate Professor in the School of Civil and EnvironmentalEngineering at Nanyang Technological University, Singapore. He received his PhD fromthe University of Canterbury, Christchurch, New Zealand. His research interestsinclude reinforced concrete and precast concrete structures, particularly the design ofearthquake- and blast-resistant structures.
Tso-Chien Pan is a Professor and Dean of the College of Engineering at NanyangTechnological University. He received his PhD from the University of California atBerkeley, Berkeley, CA. His research interests include damage assessment of buildingssubjected to dynamic loading and vibration isolation for structures and equipment.
Cao Thanh Ngoc Tran is a PhD candidate in the School of Civil and EnvironmentalEngineering at Nanyang Technological University, where he received his BEng. Hisresearch interests include reinforced concrete structures, particularly the design ofearthquake-resistant structures.
Fig. 1—Reinforcement details of Specimens: (a) C1A, C1B, and C1C; and (b) C2A, C2B,and C2C. Dimensions in mm. (Note: 1 mm = 0.04 in.)
ACI Structural Journal/September-October 2009 593
Test procedure and instrumentationAxial load was slowly applied to the column or wall and in
balanced steps until the designated level, 0.1fc′Ag for SpecimensC1B and C2B or 0.35fc′Ag for Specimens C1C and C2C, wasachieved. Specimens C1A and C2A were tested with zeroaxial load. During each test, the axial load was maintained bymanually adjusting the flat jacks after each load step. Thelateral load was applied cyclically through the actuator in aquasi-static fashion at the top end of the column, as shown inFig. 2. The loading procedure, consisting of displacement-controlled steps, is illustrated in Fig. 3. The fraction numbersgiven in Fig. 3 were the targeted drift ratios.
To obtain the test results that explained most of the observationqualitatively, instrumentations such as the actuator, straingauges, linear variable displacement transducers (LVDTs),and displacement transducers were installed on all specimensduring the test setup. The behavior of reinforcing bars wasmonitored by the strain gauges installed on the bars prior tocasting of the specimen on locations as shown in Fig. 4.Figures 5(a) and (b) show the arrangement of displacementtransducers and LVDTs that measure the shear deformationsof beams and joints and the flexural deformations of beams andcolumns, respectively. The transducers and LVDTs weremounted on steel brackets.
DESCRIPTION OF TEST BEHAVIORLoad-displacement hysteresis responses
The theoretical lateral strength of C1 specimens of 19.2 kN(4.3 kip) was calculated based on the theoretical flexuralstrength of the beams, utilizing the approach prescribed inthe NZS 3101 Code6 with actual material strengths andassuming a strength reduction factor of 1.0. Figures 6(a) to (c)show the load-displacement hysteresis loops of C1 specimenswith the column-to-beam width ratio of approximately 3.56.The hysteresis loops showed the degradation of stiffness and
load-carrying capacity during repeated cycles due to thecracking of the concrete and yielding of the reinforcing steel.The low attainment of stiffness and strength was attributed tothe slip of the beam longitudinal bars through the joint core.Specimens C1B and C1C reached their peak strength in thefirst cycle at a drift ratio of 2.0%. Specimen C1A obtained itspeak strength at a drift ratio of 3.0%, whereas at this driftratio, the resistance of Specimens C1B and C1C started todrop by this point. At the next drift ratio of 4%, the peaklateral loads attained by Specimens C1B and C1C were only71.1% and 74.2% of the maximum recorded values of eachspecimen, respectively. The test was ceased at this point. Atthis drift ratio, however, Specimen C1A did not suffer anysignificant drop in its strength. As illustrated in Fig. 6(a) to (c),only the maximum lateral load attained by Specimen C1Aexceeded its theoretical lateral load calculated from themeasured material properties. Compressive reinforcementsin beams of Specimens C1B and C1C attained tensile stressat a drift ratio of 1.33% due to the bond deterioration, whichsignificantly reduced the flexural moment capacity of thebeams obtained by Specimens C1B and C1C. The bonddeterioration in Specimens C1B and C1C, which could beattributed to the presence of column axial loading in thesespecimens, has resulted in smaller maximum lateral loadsobtained by these specimens than Specimen C1A. In general,the specimens from the C1 series with the column-to-beamwidth ratio of approximately 3.56 illustrated limited energydissipation characteristics when drift ratio was increasedprogressively throughout the test. Their energy dissipationswere the greatest at a drift ratio of 3.0%. Limited energieswere dissipated beyond a drift ratio of 3.0%, as shown inFig. 6(a) to (c). The lower flexural strength of the beams than
Fig. 2—Test setup (in mm). (Note: 1 mm = 0.04 in.)
Fig. 3—Loading procedure.
Fig. 4—Typical strain gauge locations (in mm). (Note: 1 mm =0.04 in.)
that of the columns resulted in significant beam damage. Inaddition, the energy dissipation of the specimens was mainlycontributed by the beams.
The theoretical lateral strength of C2 specimens of 145.0 kN(32.6 kip) was calculated based on the theoretical flexuralstrength of the beams, using the approach prescribed in theNZS 3101 Code6 with actual material strengths andassuming a strength reduction factor of 1.0. Figures 6(d) to(f) illustrate the load-displacement hysteresis loops of the C2specimens with the wall-to-beam width ratio of approximately7. All three specimens reached their peak strength in the firstcycle at a drift ratio of 3.0%. At the next drift ratio of 4%, thestrength of Specimens C2B and C2C dropped to 9.2% and9.0%, respectively. At this drift ratio, however, Specimen C2Adid not suffer any significant drop in strength. Specimen C2Aobtained a higher maximum story shear force than SpecimensC2B and C2C. As illustrated in Fig. 6(d) to (f), Specimens C2Band C2C did not reach their theoretical lateral capacitiescalculated from the measured material properties in eitherpositive or negative loading directions, whereas Specimen C2Areached its theoretical lateral capacities in both directions. At adrift ratio of 2.0%, Specimens C2B and C2C exhibited a
significant bond deterioration which reduced the flexuralmoment capacity of the beams. Due to the bond deteriorationattributed to the presence of column axial loading, themaximum lateral resistance of Specimens C2B and C2Cwere less than that of Specimen C2A. Similar to the C1specimens, the C2 specimens with the wall-to-beam widthratio of approximately 7 had shown limited energy dissipationcharacteristics. The C2 specimens attained their greatest energydissipation at drift ratio of 3.0%. Significant bond slip withinthe joint core area of the C2 specimens resulted in pinchingof the hysteresis loops, leading to limited energy dissipationcapacity and is illustrated in Fig. 6(d) to (f).
Cracking patternsThe crack patterns of the specimens from the two groups
differed from one another. The behavior of the C1 specimenswas controlled by the flexural mechanism of the weak beam.In contrast, the specimens from the C2 series exhibitedsignificant cracks on the side face of the columns.
Figure 7 illustrates the formation of cracking patterns thatoccurred during the experiment on Specimens C1A, C1B,and C1C. For the C1 specimens, most of the crack damagewas concentrated in the beams near the column. The largestflexural cracks occurred at the interfaces of the beam ends.By the end of the test, these cracks were excessive, and thebeam flexural bars were observed to have slipped throughthe joint due to loss of bond. This could result in the gradualstrength deterioration and low attainment of structural stiffnessof the specimen during drift ratios of 3.0 and 4.0%. Diagonalflexural cracks were also found at beam bottom and top of allspecimens during a drift ratio of 0.4%. It was noticed thatcracks at the beams propagated rapidly when drift ratio wasincreased to 1.0%. When a drift ratio of 1.33% was attained,cracks were found to propagate rapidly at the beam top.Cracks at beam bottom of all specimens propagated rapidlyat a drift ratio of 3.0% and eventually met with cracksformed along the beam top. The presence of the axialcompression load had an obvious influence on the crackingpattern in the column of the specimens. Throughout the test,no cracks were found on the columns of Specimens C1B andC1C, whereas flexural cracks were found on the column ofSpecimen C1A at a drift ratio of 3.0%. The presence ofcolumn axial loads in Specimens C1B and C1C helped toclose up the cracks formed on the columns of these specimensand delayed their occurrence. Bond-splitting cracks along thebeam longitudinal bars of all specimens started to occur at adrift ratio of 4.0%. No shear cracks were observed at theprotruded joint of all specimens during the test due to thethick layer of concrete at the joint core.
The typical cracking patterns of C2 specimens are shownin Fig. 8. All specimens suffered severe cracking on the sideface of the columns as illustrated in Fig. 8. These cracksbegan to develop rapidly at drift ratios of 2.0 and 3.0%. Thepresence of the axial compression load had a significantinfluence on the cracking patterns. Fewer cracks wereobserved on the side face of column of Specimen C2C ascompared to Specimens C2A and C2B. The axial compressionload helped to close up these cracks and delayed their occurrence.Specimens C2A, C2B, and C2C had similar cracking patterns.When a drift ratio of 0.67% was exceeded, flexural crackswere found at beam bottom. The inclined shear cracks atbeam bottom, which were formed at a drift ratio of 1.0%,were believed to be extensions of these flexural cracks. Itwas noticeable that flexural cracks on the beams propagated
Fig. 5—Typical LVDT locations of specimens (in mm).(Note: 1 mm = 0.04 in.)
ACI Structural Journal/September-October 2009 595
Fig. 6—Hysteretic loops of Specimens C1A, C1B, C1C, C2A, C2B, and C2C.
Fig. 7—Typical cracking patterns of C1 specimens at DR of 4.0%.
596 ACI Structural Journal/September-October 2009
rapidly when the drift ratio was increased to the 1.0% level.When a drift ratio of 1.33% was attained, several diagonalcracks were found within the joint core area and flexuralcracks on the beams were also found to propagate rapidly.Limited new cracks were observed at the beam bottom,whereas more cracks were formed within the joint core areaafter a drift ratio of 1.33% was reached. A significant amountof new cracks formed at beam top when the drift ratio wasincreased to 2.0% and 3.0%, respectively. As illustrated inFig. 8, the cracks on the side face of the 1600 mm (63 in.)long column propagated aggressively throughout the test.These cracks were initiated at a drift ratio of 0.67% and theydeveloped rapidly as drift ratios increased to 1.33 and 2.0%.These cracks, which were formed mainly at the lowerportion of column, were caused by the fixed end moment ofbeams which rotated about the long column. The shearcracks at the column front face, as shown in Fig. 8, werebelieved to be an extension of these cracks formed at thecolumn. At a drift ratio of 4.0%, crushing of concrete at thefixed end of the beams due to flexure was observed.
Distribution of steel strains along beamFigures 9(a) to (f) show the measured strain distribution
along the beam top longitudinal bars. It was observed that thedistribution of strain along the reinforcement vary considerablywith an increase in lateral load. The degradation of anchorageresistance in beam bars passing the joint occurred in the C1specimens as the drift ratio increased to a drift ratio of 1/150.Thus, the compressive stress in the compressive reinforcementbars decreased and changed to tension as shown in Fig. 9(a) to (c).In contrast, the C2 specimens still displayed a good anchorageresistance characteristic up to a drift ratio of 2.0%, asexplained in Fig. 9(d) to (f). With reference to these strainprofiles, yielding of the longitudinal bars has extended intothe joint core area, thus indicating the bond deterioration inthese specimens. The first yield in beam longitudinalreinforcement of most of the specimens was observed ata drift ratio of 1.33%. The largest tensile strain of Specimen C2Awas detected at the beam-column interface, whereas forSpecimens C2B and C2C, it was within the joint area. Thus,it could be concluded that the plastic hinges of Specimen C2Awere successfully confined to the beam end.
At interior beam-column joints, extremely high bondstresses can be developed when an RC frame sustains largeinelastic deformations due to seismic motions. Beam barsmay be forced to yield in tension at one column face and besubjected to a large compressive stress at the opposite
column face. Also, yield penetration along a beam bar fromeither face of an interior column may considerably reducethe effective anchorage length of the bar and increasedeformations. Thus, the limit for the ratio of bar diametersto the column depth is intended to ensure that a beam barwill not slip prematurely through the joint core during cyclicreversed inelastic displacements. Currently in the NewZealand practice,6 it is regulated that
(1)
where db is the diameter of the beam bars; hc is the columndepth; αf = 1.0; a0 = 1.25; fc′ is the concrete compressivestrength; and fy is the steel tensile strength. In the Americanpractice,7 however, a less strict regulation is applied
(2)
Bond condition is determined mainly by the ratio of thebeam and column bar diameter to the column and beamdepths. The ratio of beam bar diameter to the column depthof all specimens, db/hc = 1/22, does not satisfy the requirementgiven by NZS 31016 as shown in Table 2. Hence, it is notsurprising that bond deterioration occurred along the beambars at the drift ratio of 4%.
Distribution of steel strains along columnIt was observed that the strains along the column steel bars
in C1 specimens were very small; this was supported by littleflexural cracks observed at the column during the test. Whenthe beam reached its flexural strength, the column was stillin the elastic range, indicating a strong column-weak beamresponse. As such, the ratios of column flexural strength tobeam flexural strength were large.
Similar to the C1 specimens, the measured strains of thelongitudinal reinforcement of the C2 specimens had neverexceeded its elastic limit. The presence of compressivecolumn axial load had an obvious effect on the strains of thecolumn steel bars. Smaller tensile strains of Specimen C2Cthan Specimens C2A and C2B were observed during the test.No bond deterioration was observed along the column barsof all tested specimens because the column was still withinits elastic range.
DISCUSSION OF TEST RESULTSDecomposition of interstory drift
The total interstory drift recorded at the top of the columnsconsisted of several components. The major componentscomprised of lateral displacements due to the beam flexural,beam shear, and column flexural deformations, as well as the
db
hc
----- 3.3αffc′
a0 fy
---------≤
db
hc
----- 0.05≤
Fig. 8—Typical cracking patterns of C2 specimens at DRof 4.0%.
Table 2—Beam bar diameters used in tested specimens
db, mm hc, mm db/hc
db/hc regulated by
NZS6db /hc regulated by
ACI-ASCE 3527
13 280 0.046 0.019 0.05
20 300 0.066 0.019 0.05
25 300 0.083 0.019 0.05
Note: 1 mm = 0.04 in.
ACI Structural Journal/September-October 2009 597
beam-column joint shear distortion. Data captured byLVDTs mounted on the specimens were used to derive thedifferent sources of deformations, following the proceduresdescribed by Wu.8 In general, the total calculated lateraldisplacements due to the contributing components were lessthan the measured interstory drift. The uncounted lateraldisplacement in C2 specimens could mainly be attributeddue to the shear cracks occurred outside the plastic hinge inthe beam, which were not measured during the test.
For C1 specimens, column deformation might havecontributed to the total displacement, but it was comparativelyinsignificant throughout the test of Specimens C1A, C1B, andC1C. The contribution to the total drift of beam flexure waspredominant; it contributed to more than 50% of the interstorydrift. This conformed to the visible crack patterns on the testunits, in which the flexural tension cracks in the beam and the
vertical cracks along the beam-column interfaces were thedominating cracks as shown in Fig. 7. The shear deformationin the beam was relatively insignificant when compared tothe beam flexure deformation. Due to the three-dimensionalnature of the specimen, the transducers placed diagonally inthe joint panel could not capture the joint shear deformation.It explained the contribution to the total drift from the jointof Specimens C1A, C1B, and C1C at a drift ratio of 4.0%was approximately 1.0%.
The contribution to the total drift of beam flexure, beamshear, column flexure, and joint of Specimen C2B at 4% driftratio were 39.0%, 2.5%, 9.2%, and 0.0%, respectively. Thecontribution to the total drift of beam flexure was predominant—it varied from 26.6 to 39.0%. Due to the three-dimensionalnature of the specimen, the transducers placed diagonally inthe joint panel could not capture the joint shear deformation.
Fig. 9—Local strains along beam top bars of specimens.
598 ACI Structural Journal/September-October 2009
It explained zero contribution to the total drift from the joint.Specimen C2C showed a similar trend as Specimen C2B,where the contribution to the total drift of beam flexure,beam shear, column flexure, and joint at 4% drift ratio were36.1%, 4.0%, 5.0%, and 0.0%, respectively. The contribution tothe total drift of beam flexure of Specimen C2C waspredominant—it varied from 23.2 to 36.1%. For Specimen C2A,the flexural deformations of beam and column were the majorsources of lateral displacement, both of which contributed to atotal of more than 50% of the interstory drift. The presence ofthe axial compression load obviously reduced the flexuraldeformation of column of the specimens.
Energy dissipation capacityThe total energy dissipated from Specimens C2A, C2B, and
C2C were approximately 50.1, 51.2, and 45.1 kNm (443.7,453.5, 399.5 kip-in.), respectively, at a drift ratio of 4.0%.Compared to Specimens C2A and C2B, energy dissipationcapacity of Specimen C2C was less favorable, although theenergy dissipation was generally increased constantlythroughout the test. The column axial load level did not signifi-cantly affect the energy dissipation capacity of the specimens.Similarly, the column axial load level only slightly affected theenergy dissipation capacity of Specimens C1A, C1B, and C1Cwith a column to beam width ratio of approximately 3.56.
Nominal joint shear stressTo assess the maximum joint shear stress, the joint shear
force was calculated as
Vjh = T + Cc′ + Cs′ – Vcol (3)
where Cc′ and Cs′ are the compressive force carried by theconcrete and top beam reinforcing bars, respectively, and T isthe tension force induced by top beam bars as shown in Fig. 10.9
Maximum joint shear stress can be calculated as follows9
(4)
where hc is column depth, and bj is the effective joint widthas defined in Fig. 11.9
The maximum joint shear stress coefficient γ is calculatedas follows
(5)
According to NZS 3101,6 for interior beam-column jointswith nonseismic detailing, the maximum joint shear stress isbetween 0.11fc′ and 0.17fc′ , as shown in Table 3. This correlateswell with that of the experimental results of the specimens withcolumn-to-beam width ratio of approximately 3.56. This isconsistent with similar test results conducted by Li et al.2,3 atNTU, when the specimens, which had a beam-to-columnwidth ratio of approximately 3, were tested with a zerocolumn axial load. Based on the limited test data, using thelower value of NZS range,6 0.11fc′ , gives the lower boundfor the maximum shear stress of the nonseismically detailedjoints with a column-to-beam width ratio within from 0 to3.56. The effects of column axial load level were not significantin the comparison of Specimens C1A, C1B, and C1C, as the γvalue achieved was similar in all specimens. It is known thatthe column flexural strength to beam flexural strength ratiossignificantly affect the global as well as the local behavior ofthe specimens. As shown in Table 3, the column flexuralstrength to beam flexural strength ratios of Specimens C1A,C1B, and C1C were 4.62, 5.23, and 6.36, respectively, whichsatisfied the strong column-weak beam criterion. The failureof beam in a weak beam-strong column combination couldhave overridden the effect of column axial loading. Thisobservation was supported by the absence of major diagonalshear cracks in C1 as beam failure was the main cause.
CONCLUSIONSThe present investigation is concerned with the assessment of
RC interior beam-wide column joints, which are often found
vjhVjh
bjhc
---------=
γvjh
fc′---------=
Fig. 10—External and internal actions of interior beam-column joint.9
in framed structures with nonseismic detailing in Singapore,and the effects of column axial loadings on the seismicbehavior of such joints. The research has shown that thejoints attained a drift ratio of 2.0% without significantstrength degradation. The low attainment of stiffness andstrength was attributed due to the slip of the longitudinal barsthrough the joint core. It is concluded that such RC interiorbeam-wide column joints with nonseismic design anddetailing might also possess the inherent ductility toadequately respond to unexpected moderate earthquakes.
For nonseismically detailed RC interior beam-widecolumn joints, column to beam width ratio of approximately3.56 under zero, 0.1fc′Ag, and 0.35fc′Ag applied axialcompression loads, the maximum nominal horizontal shearstresses in the joint core was approximately 0.11fc′ . The axialcompression loading did not significantly affect the stiffness,energy dissipation capacity, and maximum nominal horizontalshear stresses. The failure of beam in a weak beam/strongcolumn combination could have overwritten the effect of theaxial compression loading. The maximum nominal horizontalshear stresses correlate well with NZS 3101,6 which suggestedthat for interior beam-column joints with nonseismic detailing,the maximum joint shear stress is between 0.11fc′ and0.17fc′ . This is consistent with the conclusions drawn byLi et al.2,3 Based on the limited test data, the lower valueof NZS 3101 range,6 0.11fc′ can be used to give the lowerbound for the maximum shear stress of the test results.
ACKNOWLEDGMENTSThe financial assistance provided by Building & Construction Authority
and Ove Arup, Singapore, is gratefully acknowledged.
NOTATIONAg = gross sectional area of columnbj = effective joint width
bw = width of beamDR = story drift ratiofc′ = concrete compressive strengthfy = yield strength of reinforcementhc = column depthIg = moment of inertia based on uncracked gross concrete areaVc = column shear forceVjh = joint shear forcevjh = nominal joint shear stressγ = joint shear strength coefficient
REFERENCES1. Kim, J., and LaFave, J. M., “Key Influence Parameters for the Joint
Shear Behavior of Reinforced Concrete (RC) Beam-Column Connections,”Engineering Structures, V. 29, No. 12, 2007, pp. 2523-2539.
2. Li, B.; Wu, Y. N.; and Pan, T.-C., “Seismic Behavior of NonseismicallyDetailed Interior Beam-Wide Column Joints—Part I: Experimental Resultsand Observed Behavior,” ACI Structural Journal, V. 99, No. 6, Nov.-Dec.2002, pp. 791-802.
3. Li, B.; Wu, Y. N.; and Pan, T.-C., “Seismic Behavior of NonseismicallyDetailed Interior Beam-Wide Column Joints—Part II: Theoretical Comparisonsand Analytical Studies,” ACI Structural Journal, V. 100, No. 1, Jan.-Feb. 2003,pp. 56-65.
4. Hakuto, S.; Park, R.; and Tanaka, H., “Seismic Load Test on Interiorand Exterior Beam-Column Joints with Substandard Reinforcing Details,”ACI Structural Journal, V. 97, No. 1, Jan.-Feb. 2000, pp. 11-24.
5. BS 8110, “Structural Use of Concrete, Part 1. Code of Practice forDesign and Construction,” British Standard, 1997.
6. NZS 3101, “Concrete Structures Standard (1998): Part 1—The Designof Concrete Structures,” 1998.
7. Joint ACI-ASCE Committee 352, “Recommendations for Design ofBeam-Column Connections in Monolithic Reinforced Concrete Structures(ACI 352R-02),” American Concrete Institute, Farmington Hills, MI, 2002,37 pp.
8. Wu, Y. M., “Experimental and Analytical Study of ReinforcedConcrete Interior Beam-Wide Column Joints for Seismic Performance,”thesis, Nanyang Technological University, Singapore, 2001.
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