ductile details for welded unreinforced moment connections subject to inelastic cyclic loading

Engineering Structures 25 (2003) 667–680www.elsevier.com/locate/engstruct

Ductile details for welded unreinforced moment connectionssubject to inelastic cyclic loading

James M. Ricles∗, Changshi Mao, Le-Wu Lu, John W. FisherDept. of Civil and Environmental Engineering, Lehigh University, 117 ATLSS Drive, Bethlehem, PA, 18015-4729, USA

Abstract

The results of a 3-D finite element study of welded unreinforced flange beam-to-column moment connections in steel specialmoment resisting frames are presented. Computer models of connection subassemblies were developed using the general-purposenonlinear finite element programabaqus. Several issues were addressed in the study: (1) geometry and size of the weld accesshole; (2) benefit of a welded beam web; (3) control of inelastic panel zone deformations; and (4) effectiveness of continuity platesin reducing local demand on the connection. The analytical results provided information related to basic performance and the effectsthat these connection parameters have on inelastic cyclic performance, thereby furthering the current understanding of weldedmoment connection behavior under seismic loading conditions and leading to improved design criteria. Based on the results of theanalytical study, recommendations for the seismic design of connections are given. The incorporation of the recommendations intoan experimental test program showed good connection ductility in the test specimens. 2002 Elsevier Science Ltd. All rights reserved.

Keywords: Beam web attachment; Connection; Continuity plates; Cyclic tests; Ductility; Finite element analysis; Fracture; Inelastic rotation; Panelzone; Plastic strain; Weld access hole; Welded

1. Introduction

Numerous welded beam-to-column moment connec-tions in steel moment resisting frames (MRFs) failedduring the 1994 Northridge Earthquake [1]. The failuresraised many questions regarding the validity of designand construction procedures used at the time for theseconnections. Since the earthquake, several extensiveanalytical and experimental studies have been conductedto investigate the various aspects believed to be associa-ted with the failure observed in the pre-Northridge con-nection and to improve connection performance [2–9].This paper presents the results of recent research con-ducted under Phase 2 of the SAC Steel Project thatfocused on the seismic resistance of welded unreinforcedflange moment connections with improved details forspecial MRFs. The improvements in the connectiondetail include the use of notch tough electrodes, properlycontoured and sized weld access holes, beam web fullpenetration welds, sufficient column panel zone strength,and column continuity plates with adequate thickness.

∗ Corresponding author. Fax.:+1-610-758-5553.

0141-0296/03/$ - see front matter 2002 Elsevier Science Ltd. All rights reserved.doi:10.1016/S0141-0296(02)00176-1

Numerous weld access hole fractures have beenreported in post-Northridge earthquake inspections andin laboratory tests, where conventional access hole con-figurations were employed in the connections. Previousfinite element analysis [6] has shown that small accessholes result in less strain concentration around the holes.However, the reduced hole diameter restricts access dur-ing welding, which tends to increase the size of the welddefects resulting from incomplete fusion at the root ofthe flange groove weld. For the beam lower flange thecritical location of major defects often occurs near theaccess hole, where fracture may initiate. Additional stud-ies are therefore needed to further examine the effect ofthe size and geometry of the weld access hole on thefracture potential of the material near the hole.

The typical shear tab of a pre-Northridge moment con-nection is welded to the column flange as well as boltedto the beam web. The shear tab is designed to resist thebeam shear force, with the welded beam flanges resistingthe beam bending moment. Previous experimental stud-ies [5,10] have shown that web supplemental fillet welds,which are provided along three edges of the shear tab,and full penetration groove welds between the beam weband column face enhance the strength, ductility and

668 J.M. Ricles et al. / Engineering Structures 25 (2003) 667–680

energy absorbing capacity of the connection. While thecontribution of web welds are considered to be ben-eficial, it is not clear how much contribution they offer.Also, it is not clear what type of modification is suf-ficient to enhance the ductility of the connection.Additional experimental and theoretical studies areneeded to develop a better understanding of the behaviorand effect of bolted versus a welded shear tab onmoment connection behavior.

Panel zone yielding has been shown [11] to provideconsiderable ductility in inelastic deformation. Recentbuilding codes had placed increased emphasis upon util-izing this ductility in seismic design. Deierlein and Chi[8] investigated the effects of panel zone deformationson the fracture toughness demands in welded beam-to-column connections. Their analyses concluded that largepanel zone deformations in connections with weak panelzones could considerably increase the fracture toughnessdemands in the connection. This suggests the need toreevaluate the AISC seismic design provisions [12,13],which may result in significant panel zone deformationsin buildings.

At the time that this research study was initiated,FEMA [14] recommended the use of continuity platesfor seismic design in all cases. However they were notrequired if connection test results can show that otherdesign features of a given connection are effective inreducing or redistributing flange stresses such that theconnection performs satisfactory without them. AISC[12,13] requires that for seismic resistant design conti-nuity plates be provided to match the tested connection.Thin continuity plates will likely result in lower residualstresses, which will reduce the potential of weld crackingbecause of less heat input during welding. Some labora-tory tests of connections show good performance withoutthe use of any continuity plates. Thus, continuity plateseismic design requirements needed to be reevaluated.

To investigate these issues, nonlinear finite elementmodels were used to conduct a parametric study toinvestigate the effects of these details on the inelasticresponse of the connection. The analyses were conductedapplying monotonic increasing static load and cyclicvariable amplitude load to the models, respectively. Theanalytical study was then followed by an experimentalstudy consisting of full-scale test specimens. This paperdescribes the main results and conclusions derived fromthese studies.

2. Finite element parametric study

2.1. Finite element model

The connection subassembly shown in Fig. 1 was usedin the analytical parametric study. This subassembly isthe same as that utilized in the experimental study for

Fig. 1. Specimen T1: (a) test setup, and (b) connection detail.

exterior connections by Ricles et al. [15] under Phase 2of the SAC Steel Project. Specimen T1 was chosen asthe control specimen for the analysis. It consisted of aW36 × 150 beam connected to a W14 × 311 column.These member sizes are typical in MRFs in Californiafor newer construction involving Grade 50 steel. Thebeam flanges are joined to the column flange with fieldgroove welds using the flux core arc welding procedurewith E70TG-K2 electrodes. This weld metal has highnotch toughness, with a typical toughness of 27–54Joules at �28 °C. The bottom backing bar is removed,the root pass of the groove weld back-gouged, andreinforced with a notch tough fillet weld, while the topbacking bar is left in place and reinforced with a closure

669J.M. Ricles et al. / Engineering Structures 25 (2003) 667–680

fillet weld. The beam web is connected to the columnflange using a groove weld with supplemental fillet weldaround all edges of the shear tab. The shear tab servesas an erection device during construction (through theuse of erection bolts) and as a backup bar for the webgroove welds. Continuity plates are located on both sidesof the column web. Changing some of the details ofSpecimen T1 creates different connection configurations.These include the weld access hole geometry, beam webattachment detail, panel zone doubler plates, and conti-nuity plates.

The three-dimensional finite element models of theconnection subassembly have been developed for theparametric study using the general-purpose nonlinearfinite element analysis program abaqus [16]. The finiteelement model is shown in Fig. 2, and consists of eight-node brick elements that utilize standard integration(element C3D8 in the abaqus element library). Themodels include details such as all welds, the shear tab,bolt holes and bolts, and the top backing bar (the bottombacking bar is removed in the connection). In the para-metric study the analyses are conducted by applying aprescribed monotonic displacement to the top of the col-umn to achieve a story drift of 0.03 radians of inelasticrotation, the amplitude of inelastic rotation that connec-tions must achieve in qualification tests prescribed byAISC [12] and SAC [14]. More recent requirements[13,17] require a total story drift capacity (i.e., elasticplus inelastic deformations) of 0.04 radians, which isabout the same as the older requirements since the elasticcomponent of drift of test specimens is typically 0.01radians.

The bottom of the column and the end of the beamin the model has pin and roller boundary conditions,respectively. The sub-model shown in Fig. 2(b) wasdeveloped to obtain increased accuracy in the computedlocal stress-strain state of the tension beam flange regionof the connection, where the connection has the highestfracture potential. The displacement results of the global

Fig. 2. Three-dimensional finite element model.

model shown in Fig. 2(a) are used for the boundary con-ditions around the perimeter boundary of the sub-model.The analyses account for material nonlinearities usingthe von Mises yield criterion. Isotropic hardening isassumed for the monotonic analysis, whereas kinematichardening is assumed for the cyclic analysis. Geometricnonlinearities are accounted for through a small strain,large displacement formulation. The members in themodel are assumed to be A572 Gr. 50 steel. In the analy-sis the measured stress-strain properties of the materialsreported by Ricles et al. [15] were used. The stress-straincurves for the materials are shown in Fig. 3. Elastic andinelastic convergence studies have been conducted toevaluate and arrive at the final mesh for the finiteelement models. The finite element model was verifiedby comparing the measured cyclic response of SpecimenT1 [15] with the predicted cyclic response. The detailsof these studies and how the material and geometric non-linearities are implemented into the analyses are reportedin Ricles et al. [15].

Unstable crack propagation is not addressed in theanalytical study. Rather, the study is concerned with the

Fig. 3. Stress-strain relationships used for finite element analysis.


potential of ductile fracture through the development ofstress and strain states that would facilitate fracture if aflaw or other irregularity exists. Furthermore, as notedabove, the parametric study is conducted for monotonicloading conditions only. As in prior analytical studies onwelded moment connections performed by El-Tawil etal. [6], it is assumed that the conclusions drawn from thisstudy are qualitatively applicable to cyclic conditions.

2.2. Response indices

In the analyses cracks are not explicitly modeled. Toevaluate and compare the different analyzed connectionconfigurations for ductile fracture potential, a ruptureindex is computed at different locations of the connec-tion. Others [6] have used the same approach in analyti-cal connection studies. The rupture index (RI) isdefined as:

RI �εp /εy

exp��1.5σm

σeff� (1)

where �p, �y, σm, and σeff are, respectively, the equivalentplastic strain, yield strain, hydrostatic stress, and equiv-alent stress (also known as the von Mises stress). Therupture index was motivated by the research of Hancockand MacKenzie [18] on the equivalent plastic rupturestrain of steel for different conditions of stress triaxiality.The process of ductile fracture initiation is caused byhigh tensile triaxial stresses (i.e., high tensile hydrostaticstress) that result in damage accumulation throughmicrovoid nucleation and coalescence. The ratio ofhydrostatic stress-to-von Mises stress (σm/σeff) thatappears in the denominator of (1) is called the triaxialityratio (TR). High triaxiality can cause a large reductionin the rupture strain of a material, thereby limiting itsductility [19]. Thus, locations in a connection withhigher values for RI have a greater potential for fracture.

The ratio of equivalent plastic strain-to-yield strainthat appears in the numerator of (1) is called the plasticequivalent strain (PEEQ) index. This index is a measureof the local inelastic strain demand, and is also usefulin comparing the different analyzed configurations. ThePEEQ index is computed by:

PEEQ Index ��2

3εp

ijεpij

εy(2)

where �pij are the plastic strain components.

The triaxiality ratio and PEEQ index are also com-puted at different locations of the connection to provideadditional means of comparing the analyzed connectionconfigurations. The locations determined to have thehighest fracture potential in the different analyzed con-figurations [15] are in the weld access hole region, near

the interface of the weld metal and base metal at thebottom surface of the beam tension flange, and the beamweb groove weld (see Fig. 4). Comparisons are made atthese locations when the inelastic story drift (θp) is at0.03 radians.

3. Weld access hole

The parametric study for the weld access holeinvolved evaluating the nine different weld access holeconfigurations shown in Fig. 5. The primary geometricalparameters of the weld access hole are the overall length,length of the flat portion, slope, and height. Configur-ation 1 is the configuration commonly used in US con-struction and in recent tests by Engelhardt [20] and Sto-jadinovic et al. [9] with post-Northridge connectiondetails. It is similar to one of the standard weld accesshole geometries prescribed by AISC [21] for rolled sec-tions, and is the minimum permitted hole size. The diam-eter of the circular portion of the hole is 20 mm with alength equal to 1.5 times the thickness of the beam web.Configuration 2 is one without a weld access hole, whichis not practical for actual fabrication but is studied forthe purpose of comparison. The size and geometry ofthe other configurations were developed from the stan-dard hole with the intent to minimize plastic straindemand in order to reduce the potential for fracture ofthe beam flange near the hole region. The diameter ofthe holes in all the configurations was 20 mm, which isprobably the smallest diameter that still permits properwelding of the beam flanges. The prototype connectionconfiguration used is that of Specimen T1, where thebeam web is welded directly to the column flange witha groove weld and supplementary fillet welding is placedalong the edges of the shear tab.

The profile of the PEEQ index value across the beamflange on a line that passes through the root of the weldaccess hole (point A in Fig. 5) is shown in Fig. 6 forfour selected configurations. Configuration 1 of the weldaccess hole is shown to have a high concentration ofplastic strain in the middle of the flange, at the root ofthe weld access hole where initiation of fracture occurredin the connection tests conducted by Engelhardt [20] and

Fig. 4. Regions of connection with highest fracture potential.


Fig. 5. Various weld access hole configurations investigated.

Fig. 6. PEEQ index across beam tension flange (root of the accesshole) for selected weld access hole configurations, θp � 0.03 radians.

Stojadinovic et al. [9]. The maximum value of the PEEQindex that developed in the weld access hole region (thelocation is identified for each configuration in Fig. 5)is summarized in Fig. 7 for all nine weld access holeconfigurations. Note that the PEEQ index of the standardhole (configuration (1)) is twice that of configuration (6),which appears to be the best access hole configuration.A summary of the triaxiality ratio and rupture index forweld access hole configurations 1 and 6 is given in Table1. Configuration 1 was found to have the highest triaxial-ity ratio and rupture index among the weld access holegeometries studied, while configuration 6 was found tohave the lowest. The maximum rupture index in the holeregion of configuration 6 is equal to 45% of that of con-figuration 1. Thus, configuration 6 would have the lowestpotential for crack initiation among the different holeconfigurations. Configuration 6, referred to as the modi-

Fig. 7. Maximum PEEQ index at the critical location within theaccess hole region, θp � 0.03 radians.

Table 1Comparison of triaxiality ratio and rupture index for weld access holeconfigurations 1 and 6, θp � 0.03 radians

Configuration Triaxiality ratio Rupture index

1 0.62 2556 0.55 114


fied weld access hole, was used in fabricating the con-nection specimens tested by Ricles et al. [15]. The accessholes were ground to a surface roughness of about 6.3µm to remove any notches. These specimens had nofracture occur in the access hole region of the beamflange during testing, with inelastic story drifts of up to0.05 radians being achieved [15].

4. Beam web attachment detail

To investigate the effect of the beam web attachmentdetail on connection performance, four different webattachment details were studied. These included thebolted and welded web details that are summarized inTable 2. Web attachment detail I is the same as Speci-men T1 (Fig. 1), where the beam web is groove weldedto the column flange and 10-mm supplemental filletwelds are placed around the shear tab on the beam web.Detail II is similar to detail I, except that the supplemen-tal fillet welds are omitted. Detail III has the beam webattached to the shear tab by fillet welding, with nogroove weld connecting the beam web to the columnflange. Detail IV is a bolted beam web connection usingten 25-mm diameter A325X bolts to connect the beamweb to the shear tab. The shear tab for all cases wasfillet welded to the column flange. All the details hadthe modified access hole, column web continuity plates,and no doubler plates. Doubler plates were not requiredwith the current AISC seismic provisions [12,13].

A summary of the analysis results are given in Fig.8, where the values of the PEEQ index, triaxiality ratio,and rupture index at the middle and edge of the interfaceof the weld metal and base metal of the beam flangetension flange, weld access hole region, and edge of thebeam web are shown. These locations, identified in Fig.4, have the greatest potential for fracture.

Fig. 8(a) shows that the web attachment associatedwith detail III has the largest PEEQ index, which occursat the edge of the fillet weld attaching the shear tab to thecolumn flange. The PEEQ index is shown to be largest atthe edge of the beam flange for details II and IV, andat the edge of the beam web groove weld for detail I.For all cases the triaxiality ratio is shown (Fig. 8(b)) tobe the largest at the edge of the beam web. The restraint

Table 2Analysis matrix of beam web attachment details

Case Beam web attachment detail

I Beam web groove weld with 10 mm fillet weld alongshear tab edges

II Beam web groove weldIII 10 mm fillet weld along shear tab edgesIV Bolted beam web using 10 bolts

Fig. 8. PEEQ index, triaxiality ratio, and rupture index at selectedlocations for various beam web attachment details, θp � 0.03 radians.

by the beam web groove weld and supplemental filletweld in detail I result in the largest triaxiality ratioamong the attachment details. The corresponding valuesfor the rupture index of the four details are given in Fig.8(c), where details I and III are seen to have a relativelylarge value at the beam web edge. In detail I this is atthe end of the beam web groove weld, while for detail IIIthis occurs in the shear tab-to-column flange fillet weld.

Overall, detail I has the smallest PEEQ and ruptureindex among the four details at the interface of the beamtension flange and weld metal, as well as the access holeregion. The rupture index at the beam tension flange andweld metal is reduced by about 45% for the all weldedweb detail (i.e., detail I) compared to the bolted web


detail (detail IV). At the access hole region the reductionis about 34%. The reduction in the rupture index is dueto the beam web groove weld and supplemental filletweld in detail I creating a stiff web attachment,restraining out-of-plane local buckling of the beam webnear the column face as well as providing a load transferpath for the beam moment. These results suggest thatthe fracture potential of the beam flanges is reducedwhen the beam web is groove welded to the column andsupplemental fillet welds are used. However, the fracturepotential at the beam web groove weld is increasedbecause of the high restraint. The fracture of the beamweb weld will alter the stress and strain distribution pat-tern and impose more demand on the beam flanges, poss-ibly leading to less desirable performance of the connec-tion.

In all the web attachment details no doubler plateexist. As will be discussed next, the high local strain atthe edge of the beam web groove weld in detail I iscaused by the weak panel zone, which has undergonesignificant inelastic shear deformations. The ratio ofpanel zone-to-beam moment (Mpz/Mbm) is 1.02, whereMpz is the beam moment developed at the face of thecolumn based on the panel zone nominal design strengthRu (Eq. (3) given below) and accounting for the columnshear, and Mbm is the beam moment at the column facedue to the beam flexural capacity Mp developing in theplastic hinge at a distance equal to one-half the beamdepth from the column face.

5. Panel zone strength

Inelastic shear deformations develop in the panel zonewhen it yields. Panel zones with a larger strength willforce more of the inelastic behavior to occur in thebeams that are attached to the connection, and therebyreduce the amount of inelastic shear deformations in thepanel zone. The panel zone design shear strength Ru

given by the current AISC seismic provisions [12,13] isequal to:

Ru � φv0.6Fycdctp�1 �3bcft2cf

dbdctp�, (3)

where Fyc, bcf, tcf, dc, db and tp are the panel zonematerial yield strength, column flange width, columnflange thickness, column depth, girder depth, and panelzone thickness, respectively, and φv is the resistance fac-tor (equal to 0.75). Eq. (3) is based on the work of Kraw-inkler [11], where the panel zone is assumed to haveundergone an average shear strain that is four times theyield shear strain γy when the full panel zone strengthis developed. The second term in Eq. (3) accounts forthe increase in panel zone shear force beyond yieldingdue to the bending restraint of the column flanges. The

1997 AISC seismic provisions [12] require in design thatRu need not exceed the panel zone shear force determ-ined from 80% of the sum of the expected momentcapacity Mpe of the beams framing into the connection.Supplement No. 2 to the 1997 AISC seismic provisions[13] bases the minimum required panel zone strength on100% of the summation of the expected beam capacities,while stipulating that φv � 1.0 in Eq. (3). For Grade 50steel, Mpe is equal to 1.1 Mp. The beam moment Mbm

used throughout this paper is equal to approximately 1.1Mp, and therefore is also equal to Mpe.

The effect of panel zone deformations on the ductilefracture potential of the connection was investigated,where the configuration and details of Specimen T1 areagain adopted, with three different panel zone strengths.The following three panel zone details are used: (1) nodoubler plate; (2) one-13 mm thick doubler plate with ayield stress of 345 MPa (i.e., A572 Gr. 50 steel); and(3) one-25 mm thick doubler plate with a yield stress of345 MPa. The corresponding Mpz/Mbm ratio based onusing nominal yield strengths and φv � 0.75 in Eq. (3)is 1.02, 1.30 and 1.58, respectively, for the three panelzone designs.

The PEEQ index, triaxiality ratio, and rupture indexfor all the three panel zone designs are plotted in Fig. 9as a function of the Mpz/Mbm ratio. The locations corre-sponding to the results plotted in Fig. 9 are at the edgeof the beam web weld, the interface of the weld metaland base metal at the middle of the beam tension flange,and the weld access hole region (see Fig. 4). The PEEQindex in the beam web weld is shown to be reduced inFig. 9(a) by a factor of about 3.4 when the panel zonestrength is increased from Mpz /Mbm � 1.02 to 1.58.Increasing the panel zone strength, however, is foundnot to significantly change the PEEQ index in the accesshole region and beam flange. The triaxiality ratio is thehighest at the beam web groove weld compared to otherlocations in the connection (see Fig. 9(b)) due to thehigh local restraint at the beam web-column flange inter-face. A slight decrease occurs in the triaxiality ratio atthis location with a stronger panel zone. The ruptureindex at the beam web weld is shown in Fig. 9(c) tosignificantly decrease with an increase in panel zonestrength, whereas the rupture index at the beam flangeand weld access hole region are almost constant withpanel zone strength.

These results indicate that connections with a strongerpanel zone, which consequently have a reduced amountof inelastic shear deformation, have less fracture poten-tial than connections with a weak panel zone. The frac-ture potential of the groove welded beam web attach-ment detail is significantly reduced, making thisconnection detail more appealing. These results supportthe finding by Deierlein and Chi [8].

To reduce the inelastic panel zone deformations, it isproposed that the panel zone be designed whereby at the


Fig. 9. Effect of panel zone strength on PEEQ index, triaxiality ratio,and rupture index at selected locations in connection, θp � 0.03 radi-ans.

onset of panel zone yielding the beams that are adjacentto the connection reach their plastic moment capacity Mp

in the beam plastic hinge. The onset of panel zone yield-ing is associated with a shear strain of γy, and thereforethe first term in Eq. (3):

Ry � fv0.6Fycdctp (4)

Eq. (4) should be used with 100% of the beammoments (i.e., the factor of 0.8 that is applied to thebeam moment demand should not be used).

Using this approach, a panel zone with a 13-mm thickGrade 50 steel doubler plate would be required forSpecimen T1. For this configuration Mpz/Mbm equals1.07, where Mpz is based on nominal material propertiesand φv is equal to 0.75 in Eq. (4). It was found that thecontribution of the panel zone deformation is reducedfrom 70% to 30% of the total inelastic story drift at aθp of 0.03 radians in the model when adding the 13-mmthick doubler plate. The corresponding rupture index atthe beam web weld for this configuration is about 172.

The current AISC seismic design provisions would notrequire a doubler plate for Specimen T1, and the corre-sponding rupture index would be equal to 392. There-fore, basing the panel zone strength Ru on only the firstterm in Eq. (3) results in a significant reduction of therupture index at the beam web groove weld.

6. Continuity plates

To investigate the effects of continuity plates on con-nection behavior a finite element model of an interiorconnection subassembly was developed, and is shown inFig. 10. The exterior connection model (Fig. 2) was alsoutilized in the continuity plate study. The finite elementmodel for the interior connection was generated usingthe same brick element used in the exterior connectionmodel. Due to computer capacity limitations, however,portions of the beams and columns away from the con-nection that remained elastic were modeled using 3-Dbeam-column elements. Sub-models were alsodeveloped for the interior connection to acquire a finermesh and more accurate result at the beam flange-col-umn interface and continuity plate region.

The bottom end of the column and ends of the beamsaway from the connection in the model had pin androller boundary conditions, respectively. Cyclic load washorizontally applied to the top of the column of themodel, following the SAC loading protocol [22] untilreaching a story drift of 0.05 radians. Both geometricand material nonlinearities were included in the modelto account for the effects of cyclic yielding and localbuckling. A nonlinear cyclic strain hardening constitut-ive relationship was employed in the model by combin-ing the nonlinear kinematic and isotropic strain harden-ing relationships obtained from cyclic coupon test resultsof the steel.

Eight different cases of various continuity plate thick-ness were analyzed, as indicated in Table 3. In Table 3

Fig. 10. Global finite element model of interior connection.


Table 3Analysis matrix of continuity plate thickness

Case Test Specimen Column Size Connection Continuity Plate thickness bf,bm/tf,col

1 C1 W14 × 398 Interior 0 4.22 C2 W14 × 398 Interior tf,bm 4.23 C3 W27 × 258 Interior 0 6.84 – W27 × 258 Interior 0.5tf,bm 6.85 C4 W27 × 258 Interior tf,bm 6.86 – W14 × 311 Exterior 0 5.37 – W14 × 311 Exterior 0.5tf,bm 5.38 T1 W14 × 311 Exterior tf,bm 5.3

bf,bm is equal to the beam flange width and tf,col the col-umn flange thickness. Of the eight cases, the first-fiveinvolved the analysis of an interior connection. A rangeof column flange thickness tf,col was considered, with thebeam size equal to a W36 × 150 for all cases, resultingin a beam flange width-to-column flange thicknessbf,bm/tf,col ratio of 4.2–6.8. The continuity plate thicknessranged from zero to one beam flange thickness tf,bm forthe various cases, as indicated in Table 3. A number ofthe analysis cases in Table 3 have been tested in thelaboratory, as indicated by the presence of a test speci-men label in column two of Table 3. Cases 1 through 5had a strong panel zone to match the interior connectionspecimen conditions, whereas cases 6 through 8 had aweaker panel zone to match the exterior connectionspecimen conditions of the test matrix.

Shown in Fig. 11 are the accumulated equivalent plas-

Fig. 11. Accumulated PEEQ across the beam flange near the root of the groove weld at 0.05 radians story drift.

tic strain (PEEQ) across the beam flange at the root ofthe groove weld where the strain demand was largest(see Fig. 4). Fig. 11(a) compares the results of cases fora stiff column flange (W14 × 398), where bf,bm / tf,col �4.2. It is apparent in Fig. 11(a) that the maximum

accumulated PEEQ at the center of the beam flange isthe same for both cases (cases 1 and 2 in Table 3) andthat there is little difference between the results. Theresults for cases 3, 4, and 5 involving a flexible columnflange (W27 × 258, bf,bm / tf,col � 6.8), and an intermedi-ate flexible column flange (W14 × 311, bf,bm / tf,col �5.3) for cases 6, 7, and 8 are shown in Fig. 11(b) and

(c), respectively. It is apparent in Fig. 11(b) and (c) thatthe distribution of accumulated PEEQ across the beamflange width becomes more uniform when continuityplates are added to the model. For the W27 × 258 col-umn the maximum accumulated PEEQ is also reduced


whereas for the W14 × 311 column no change in themaximum value of accumulated PEEQ occurs whenadding continuity plates. Furthermore, the distribution ofaccumulated PEEQ across the beam flange are similarfor the cases when the continuity plate thickness is either0.5tf,bm (12 mm) or 1.0tf,bm (24 mm).

A plot of the analysis results for the maximumaccumulated PEEQ at the center of the beam flange asa function of the bf,bm/tf,col ratio is shown in Fig. 12.Fig. 12 indicates that for the cases studied that continuityplates are not required when the bf,bm/tf,col ratio is lessthan 5.2, for the maximum accumulated PEEQ is thesame value for all cases with and without continuityplates. However, for larger values of the bf,bm/tf,col ratiothe use of continuity plates reduces the maximumaccumulated PEEQ. The use of a reduced continuityplate thickness of 0.5tf,bm � 12 mm is shown to be effec-tive in reducing the maximum accumulated PEEQ whenbf,bm / tf,col � 6.8.

The analysis results showed good agreement with theexperimental results of Specimens C1, C2, C3 and C4noted in Table 3. The test results of Specimens C1 andC2 showed little difference in local behavior at thebeam-column interface, despite Specimen C1 not havingcontinuity plates. Specimen C4 showed some improve-ment in behavior and connection ductility with the useof continuity plates, compared with Specimen C3 whichdid not have continuity plates and the sameW27 × 258 section for the column. More details relatedto the continuity plate analysis are given in Ricles etal. [15].

7. Experimental study

The results from the analytical study were evaluatedin the experimental study by Ricles et al. [15]. Theexperimental study involved the inelastic cyclic testingof eleven full-scale unreinforced welded moment con-nection specimens. A description of the connection detail

Fig. 12. Accumulated PEEQ vs bf,bm/tf,col ratio.

of each of the specimens is given in Table 4. All thespecimens had details similar to the details of SpecimenT1 described earlier and each had the modified accesshole geometry. Variations were made in the web attach-ment, panel zone strength, and continuity plates to createdifferent connection configurations. Specimens T1through T6 were all exterior connection specimenswhereas Specimens C1 through C5 were interior connec-tion specimens. The specimens were each subjected toa history of cyclic displacements following the loadingprotocol established by SAC [22]. The specimen details,performance and inelastic behavior are described indetail by Ricles et al. [15].

A summary of the inelastic story drift achieved by thespecimens is given in Fig. 13. A fracture did not occurin the weld access hole region in any of the specimens.Fracture typically occurred in either the heat affectedzone of the beam flange or the fusion line between theweld metal and base metal of the beam flange. In somespecimens no fracture occurred at all, and the test wasstopped due to limitations of the test setup. In SpecimensT1 and T4 cracking occurred in the beam web to columnflange groove weld and shear tab to column flange filletweld, respectively.

Specimens with a groove welded beam web and sup-plemental fillet weld attachment detail (all specimensexcept for T2, T3, and T4) achieved an inelastic storydrift greater than 0.035 radians in at least one completecycle prior to when either fracture occurred or specimenstrength had deteriorated to 80% of the nominal flexuralbeam strength. The benefit of the supplement fillet weldcan be seen by comparing the results for Specimen T1with T2, where in the latter the supplemental fillet weldwas omitted and the inelastic story drift was conse-quently reduced to 0.025 radians. A further reduction inconnection ductility occurs in Specimen T3, which hadthe beam web fillet welded to the shear tab and an inelas-tic story drift capacity of only 0.02 radians. The boltedbeam web (Specimen T4) is shown to have the lowestductility (0.018 radians of inelastic story drift). Speci-mens T1 through T4 each had Mpz/Mbm equal to 1.09,using measured material properties with Mpz based onEq. (3) and φv � 1.0. Specimen T5 was similar to Speci-men T1 except that Specimen T5 had a stronger panelzone, with Mpz/Mbm equal to 1.55 based on measuredmaterial properties and with φv � 1.0 in Eq. (3). Conse-quently, the inelastic panel zone deformations werereduced in Specimen T5 compared to Specimen T1,resulting in an increase of the overall ductility in theformer. Specimen T6 was similar to Specimen T2,except that Specimen T6 had a stronger panel zone(similar to Specimen T5) and a heavier shear tab withstronger welds. The connection detail for Specimen T6is shown in Fig. 14. This detail has the heavy shear tabshop welded to the column using a groove weld, wherethe beam web is fillet welded to the shear tab in the field,


Table 4Test specimen matrixa,b

Specimen Connection Column size Beam web attachment detail Doubler plate Continuity plate Mpzc/Mbm

d

configuration thickness (mm) thickness (mm)

T1 Exterior W14 × 311 Groove welded web; supplemental 0 25 1.09fillet welds

T2 Exterior W14 × 311 Grooved welded web 0 25 1.09T3 Exterior W14 × 311 Fillet welded shear tab 0 25 1.09T4 Exterior W14 × 311 Bolted shear tab 0 25 1.09T5 Exterior W14 × 311 Groove welded web; supplemental 1@12 0 1.55

fillet weldsT6 Exterior W14 × 311 Groove welded heavy shear tab, 1@12 0 1.55

fillet welded beam webC1 Interior W14 × 398 Groove welded web; supplemental 2@19 0 1.22

fillet weldC2 Interior W14 × 398 Groove welded web; supplemental 2@19 25 1.22




fillet weld

a A572 Grade 50 steel for all columns.b Beam size was W36×150 of A572 Grade 50 steel for all specimens.c Mpz is based on Eq. (3) with φv � 1.0 and measured material properties and dimensions.d Mbm is based on Mp in the beam plastic hinge(s) and measured material properties and dimensions.

Fig. 13. Inelastic story drift achieved in test specimens.

thereby reducing the fabrication and inspection costscompared to the web attachment detail of Specimen T1.Specimen T1 requires that the beam web groove weldbe inspected by ultrasonic testing. The stronger panelzone and web attachment detail in Specimen T6 enabledan inelastic story drift of 0.05 radians to be achieved.The hysteretic response of Specimens T1, T5, and T6are compared to each other in Fig. 15, where the nor-

Fig. 14. Specimen T6 connection detail with heavy shear tab.


Fig. 15. Beam moment at column face-inelastic story drift response of Specimens T1, T5, and T6.

malized beam moment at the column face-inelastic storydrift (M /Mp�θp) relationship is plotted. In Fig. 15, Mp

is the plastic beam moment capacity based on measuredmaterial properties and section dimensions. The steadyincrease in specimen capacity shown in Fig. 15 forSpecimen T1 is associated with the weaker panel zone,where yielding occurred both in the panel zone andbeam, however, no pronounced local buckling developedin the beam. This behavior was typical of the specimenswith a weaker panel zone (i.e., Specimens T1, T2, T3,and T4). The deterioration in specimen capacity inSpecimens T5 and T6 is associated with the response ofa connection with a strong panel zone. In these speci-mens some panel zone yielding occurred followed by thedevelopment of pronounced inelastic deformations andinelastic cyclic local flange and web buckling in thebeam that led to the degradation in beam capacity. Thisbehavior was typical of all specimens with a strong panelzone (i.e., Specimens T5, T6, C1, C2, C3, C4, and C5).

The interior connections had Mpz/Mbm equal to 1.22(Specimens C1, C2, and C5) and 1.51 (Specimens C3and C4) based on measured material properties and with�v � 1.0 in Eq. (3). The ductility of these interior con-nections with the stronger panel zones was exceptional,as shown in Fig. 13, where inelastic story drifts between0.039 and 0.052 radians were achieved.

The test results presented in Fig. 13 are consistentwith the finite element parametric study, which indicatedthat connections with a strong panel zone, a groovewelded beam web and supplemental fillet weld attach-ment, and modified weld access holes can develop betterductility by reducing the amount of inelastic shear defor-

mations in the panel zone. For the specimens with astrong panel zone (i.e., Specimens T5, T6, and all of theinterior connections) the panel zone strength relative tothe beam strength has a ratio of Mpz/Mbm ranging from0.92 to 1.07 when Mpz is based on using Eq. (4) alongwith nominal material properties and φv � 0.75. There-fore, by ignoring the increase in panel zone strength dueto the bending restraint by the column flanges the nomi-nal value for the ratio of Mpz/Mbm is closer to 1.0 forthese specimens.

8. Recommendations and conclusions

Four critical issues were examined that have a strongeffect on the ductility of welded unreinforced flangemoment connections and which should be carefully con-sidered in design. The issues are: (1) size and geometryof weld access holes, (2) beam web attachment detail,(3) control of panel zone deformations, and (4) the thick-ness of continuity plates.

The key conclusions and recommendations from theanalytical and experimental studies presented in thispaper may be summarized as follows:

1. In some connection assembly tests conducted in thepast a fracture was observed to initiate at the toe ofthe weld access hole. An effort was made in the cur-rent study to investigate the influence of the size andgeometry of the access hole on the potential of ductilefracture initiation near the holes. Nine different accesshole configurations were included in the analytical


study. The results indicate the importance of selectinga proper hole configuration. The particular configur-ation shown as configuration (6) in Fig. 5 is rec-ommended. This configuration was adopted in fab-ricating eleven full-scale interior and exterior weldedunreinforced beam-to-column test assemblies. Nofracture near the access holes was observed duringthe testing of these specimens, which responded withgood ductility.

2. Based on the analysis results, full penetration groovewelds at the beam web-column interface in conjunc-tion with supplemental fillet welds around the edgesof the shear tab reduce the accumulated equivalentplastic strain demand in the connection. It is rec-ommended for special MRFs that this beam webattachment be used in order to develop adequateinelastic deformation in the connections. The boltedshear tab detail is recommended only for ordinaryMRFs. The performance of a test specimen with aheavy shear tab, groove welded to the column flangeand fillet welded to the beam web, was exceptional.This detail may be used in lieu of the groove weldedbeam web with supplemental fillet welds.

3. In some of the exterior connection tests, initial frac-ture occurred in the vertical welds connecting theshear tab or beam web to the column flange. The testspecimens in this series had a relatively weak columnpanel zone. Limiting the amount of panel zone defor-mation can control the problem of beam web weldfracture, as indicated by the finite element analysisand verified by the experimental study, where allspecimens with a stronger panel zone had a largerinelastic story drift capacity. Limiting the panel zonedeformations would require strengthening of the panelzone with doubler plates, or using a column with athicker web in order to reduce the inelastic defor-mations in the panel zone. The connection rotationcontributed by panel zone deformation can be con-trolled by designing the panel zone such that whenthe onset of panel zone yielding occurs, the beamsattached to the connection develop their flexuralcapacity Mp in the plastic hinge zones.

4. The finite element analysis and experimental testingindicate that continuity plate requirements can likelybe relaxed for seismic design, but need further evalu-ation. All specimens tested in the experimental studywithout continuity plates developed an inelastic storydrift greater than 0.039 radians. These specimens hadcolumns of different sizes and a beam flange width-to-column flange thickness ratio ranging from 4.2 to6.8. Furthermore, the connections in the specimenshad a strong panel zone, where the inelastic behaviorwas primarily in the beams.

The results and conclusions presented are primarilybased on a finite element parametric study of various

connection designs and an experimental study of elevenfull-scale exterior and interior connection assemblies.Further research is necessary to study the behavior ofconnections having different details, including beamswith thicker flanges. Additional issues may emerge fromthis research.

Acknowledgements

The research described in this paper was supported bygrants from the SAC Joint Venture and the Departmentof Community and Economic Development of the Com-monwealth of Pennsylvania through the PennsylvaniaInfrastructure Technology Alliance. Dr. Eric J. Kauf-mann of the ATLSS Center, a specialist in welding met-allurgy, was most helpful in providing insight to someof the welding related issues.

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ductile details for welded unreinforced moment connections subject to inelastic cyclic loading

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