The Professional Journal of the Earthquake Engineering Research Institute PREPRINT

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1 The Professional Journal of the Earthquake Engineering Research Institute PREPRINT This preprint is a PDF of a manuscript that has been accepted for publication in Earthquake Spectra. It is the final version that was uploaded and approved by the author(s). While the paper has been through the usual rigorous peer review process for the Journal, it has not been copyedited, nor have the figures and tables been modified for final publication. Please also note that the paper may refer to online Appendices that are not yet available. We have posted this preliminary version of the manuscript online in the interest of making the scientific findings available for distribution and citation as quickly as possible following acceptance. However, readers should be aware that the final, published version will look different from this version and may also have some differences in content. The DOI for this manuscript and the correct format for citing the paper are given at the top of the online (html) abstract. Once the final, published version of this paper is posted online, it will replace the preliminary version at the specified DOI.

2 Cyclic Performance of Welded Unreinforced Flange Welded Web Moment Connections Sang Whan Han, a) Ki Hoon Moon, a) and Jin Jung a) The Welded Unreinforced Flange Welded Web (WUF W) moment connection is a prequalified connection for the special moment frame (SMF) specified in AISC In this study, inelastic cyclic tests of four WUF-W specimens were conducted to evaluate the seismic performance of WUF-W connections with different beam depths and panel zone strength ratios. The specimens were made to satisfy the design and detailing criteria specified in AISC and AISC Test results showed that the WUF W connections with a beam depth of 692 mm passed the acceptance criteria required for SMF connections, whereas the WUF W connections with a beam depth of 89 mm did not meet the criteria. INTRODUCTION Steel moment frames have been popularly used to resist seismic forces in various seismic regions due to their great expected energy dissipation capacity. In ASCE/SEI 7-1 (ASCE 21), moment frames are classified as special, intermediate or ordinary moment frames (SMF, IMF, and OMF, respectively) according to their ductility and energy dissipation capacity. The SMF has the highest ductility and greatest energy dissipation capacity among them. For this reason, the most stringent detailing and design requirements are imposed on SMFs (AISC 21a). In AISC (AISC 21b), seven different connections were listed as prequalified connections for use in SMFs and IMFs. The welded unreinforced flange welded web (WUF W) connection is an all welded moment connection as shown in Figure 1. The WUF-W is one of the prequalified connections for use in SMFs. The beam flanges are welded to the column flange using complete joint penetration (CJP) groove welds that meet the requirements of demand critical welds in AISC (AISC 21a), along with the requirements for treatment of backing and welding tabs. The beam web is welded directly to the column flange using a CJP groove a) Hanyang University, Seoul , Korea

3 weld that extends the full depth of the web. This is supplemented by a single plate connection, wherein a single plate is welded to the column flange and is then fillet welded to the beam web. There have been numerous reports on welded unreinforced flange bolted web (WUF B) connections from both experimental and analytical research (Popov and Stephen 1972, Popov et al. 1985, Tsai and Popov 1988, Anderson and Linderman 1991, Engelhardt and Husain 1996, Roeder and Foutch 1996, Stojadinovic et al. 21, Lee and Foutch 22. Han et al. 27). However, only limited tests have been conducted to investigate the behavior of WUF W connections. The prequalification of the WUF W moment connection is based on the results of two major research programs conducted at Leigh University (Ricles et al. 2) and the University of Minnesota (Lee et al. 25a, b). They tested 24 WUF-W connection specimens to evaluate the seismic performance of WUF W connections. Column d CJP beam flange to column Flange weld CJP beam flange to column Flange weld Single plate to column Flange weld Beam Single plate to Beam web weld Erection bolts in standard holes or horizontalshort slots are permitted as needed for erectionloads and safety CJP continuity plate to column Flange weld Fillet Continuity plate to column Web weld a. WUF-W Connection detail (ANSI/AISC 358, 21) 6mm minimum, 12mm maximum. 5mm minimum 5mm minimum. 3 (± 1 ) 12mm minimum, 25mm maximum b. Shear Connection detail (ANSI/AISC 358, 21) e d f c tbf a Note ; b a. Bevel as required for the WPS. b. tbf or 12 mm, whichever is larger (+½ tbf, or - ¼ tbf). c. The minimum dimension shall be ¾ tbf, or 2mm, Whichever is greater. The maximum dimension shall be tbf (+6mm) d. 3/8 in [1mm] minimum radius (-, +unlimited). e. 3tbf(±12mm) f. Tolerances shall not accumulate to the extent that the angle of the access hole cut to the flange surface exceed 25º c. Weld Access hole detail (AWS D1.8/D1.8,29) Figure 1. Details of WUF W connections Mao et al. (21) conducted three dimensional finite element analyses (3D-FEA) to investigate the inelastic behavior of WUF-W connections with four different beam web to column flange attachment details. They showed that the effect of the beam web attachment detail was significant on the fracture potential of the beam flanges near the interface of the weld metal and base metal. This study also investigated the influence of access hole 2

4 geometry on connection fracture near the access hole. The main parameter of the access hole geometry is the slope of the flat transition region between the beam flange and the access hole. With a decrease in the slope of the access hole, the plastic strain demand at the toe of the transition decreased, which resulted in delaying the onset of low cycle fatigue cracks at this location. Based on the results of FEA, a modified weld access hole geometry with a slope of 13 was proposed. Ricles et al. (22) conducted inelastic cyclic tests of 11 WUF W connections with the modified access hole proposed by Mao et al. (21). The specimens were made with a high toughness weld metal. They demonstrated that all tested specimens had excellent seismic performance and produced an inelastic rotation of at least.3 rad prior to failure. Lee et al. (25a,b) tested WUF W connections to investigate the effects of the doubler plate, continuity plate and weak panel zone. All specimens satisfied the rotation requirements for SMF connections. The access holes of the specimens were similar to those proposed by Mao et al. (21). The slope of the weld access hole was 15. Figure 2 shows the weld access hole geometries used by Ricles et al. (22) and Lee et al. (25a,b). Figure 2. Weld access hole geometry According to Section 8.5(2) of AISC 358-1, access hole geometry and quality requirements shall conform to the requirements of Section of AWS D1.8/D1.8M-9 (AWS 29). As seen in Figure 2c, AWS D1.8/D1.8M-9 specifies a range for the slope of the access hole, which is less than 25, instead of a single value. No experiments were conducted herein to evaluate the performance of WUF W connections with steep access hole slope as close as 25. In this study, experiments were conducted to investigate whether WUF W connections designed and detailed according to AISC 341-1, AISC and AWS D1.8/D1.8M-9 provided a sufficient rotation capacity to meet the requirements for SMF connections. WUF W connections with relatively steep access holes were considered. 3

5 TEST SPECIMENS Experiments were conducted using four WUF W connection specimens. The test variables were beam depth and panel zone strength ratio. Roeder and Foutch (1996) reported that the panel zone strength ratio and beam depth both influence the cyclic behavior of moment connections. Table 1 summarizes the information about specimens. Figure 3 shows a cantilever type exterior connection specimen in which a beam was connected to the column flange of a pin ended column with a height of 4. m. The column beam moment ratio ( * * M pc M pb ) of all specimens exceeded 1 that satisfied the strong column weak beam requirement (Section E3.4a of AISC 341-1). Figure 3 also shows the details of Specimen D7 B. All specimens in Table 1 satisfied the detail and design criteria required by AISC and AISC Specimens D9-S and D9-B with a beam depth of 89mm had the lowest ratio ( L / d ) of clear span to beam depth among WUF-W specimens, which was n b According to Section (5) of AISC 358-1, L / d for SMFs should exceed 7. It is note that WUF W specimens tested by Ricles et al. (22) and Lee et al. (25a,b) were 9.84 and Thus, in view of L / d, the beam sections used in this study were more critical that n b those used by previous studies. WUF-W specimens had a beam depth of either 692mm or 89mm, and had either strong or balanced panel zone. The panel zone strength ratio is the ratio of design panel zone shear strength to the required panel zone strength. The required panel zone shear strength ( R ) can be determined from the summation of u the moment at the column faces as determined by projecting the expected moments at the plastic hinges of the connected beams to the column faces [Eqs. (1), (2)]. According to Section 8.7 of AISC 358-1, the plastic hinge location of WUF W connections can be assumed at the face of the column. M f 1 L 1 Ru = Vc = M f db db L dc /2 h M f( L/( L dc /2)) Vc = (2) h where M f is the maximum moment expected at the face, V c is the column shear force, h is the column height, L is the beam length, ( L d c /2 ) is the clear beam length, and dc is n b (1) 4

6 the column depth. The design panel zone shear strength can be calculated according to Section 9.3 of AISC (AISC 21a). The panel zone shear strength ratio was used to classify the panel zones of WUF-W connection specimens as strong or balanced panel zones. WUF W connection specimens with a balanced panel zone were expected to experience significant yielding in both the beam and panel zone during the test, whereas specimens with a strong panel zone were expected to experience significant yielding at the beam plastic hinge, but minor yielding in the panel zone. In Eq. (1), M f can be calculated as follows: M f = Mpr = CprZxFye = 1.4( ZxRyFy ) (3) where C pr is a factor that accounts for the peak connection strength, Z x is the beam plastic section modulus, and R y is the ratio of the expected yield stress F ye to the specified minimum yield stress F y. Since the plastic hinge locations of a WUF W connection is taken to be at the face of the column (Section 8.7 of AISC 358-1), M f is equal to the probable maximum beam moment M pr. In this study C pr was set to 1.4 according to Section 8.7 of AISC The shear strength of the panel zone was calculated using Eq. (4) according to Section J1-11 of AISC 36-1 (AISC 21c). where Fyc 2 3bt cf cf φvrv = φv.6fycdctp 1+ ddt b c p is the specified minimum yield stress of the column, t p is the panel zone thickness including doubler plates, b cf is the column flange width, t cf is the column flange thickness, and φv is the strength reduction factor which is 1., as specified in Section E3.6e of AISC In this study, WUF W connection specimens with a balanced panel zone had a panel zone strength ratio ( φ Rv Ru ) of approximately 1. whereas specimens with a strong panel zone had φrv Ru greater than 1.2, an arbitrarily chosen value. The specimens were fabricated using the detailing criteria as recommended in Section J6 of AISC and Chapter 3 of AISC The beam flanges were welded to the column flange using an E71TG 1C electrode for CJP groove welds. At the top flange, the backing bar was not removed and the backing attached to the column was reinforced using a (4) 5

7 fillet weld. At the beam bottom flange, the backing bar was removed and the root pass was back gouged using the air arc procedure and reinforced with fillet weld. The runoff tabs were removed and the corners were ground smooth. The beam web was also welded directly to the column flange using an E71TG 1C electrode for CJP groove welds. A single plate was welded to the column flange and was then fillet welded to the beam web. Thus, the beam web was connected to the column flange with a CJP groove weld and a welded single plate connection. The E71TG 1C electrode satisfies the requirement of demand critical welds specified in AISC 341-1, namely, Charpy V Notch toughness of 27J at 29 C. The CJP groove welds for the beam flanges and web were inspected using the ultrasonic test procedure. For Specimens D9 S, D7 B, and D7 S, doubler plates were welded to the column web using CJP groove welds and attached to the panel zone using plug welds (Figure 3b). Plug welds are permitted by Section E3.6e.3 of AISC to prevent local buckling in the panel zone. The continuity plates were the same thickness as the beam flange and were welded to the column flanges using CJP groove welds and to the column web using fillet welds. In Section E3.6 of AISC 341-1, for one sided connections, continuity plate thickness shall be at least one half of the thickness of the beam flange. Thus, in view of continuity plates, specimens were conservatively constructed. Table 1. List of WUF W specimens Specimen Beam Size d b t t ) ( b f w f Column Size d b t t ) ( b f w f t dp (mm) t cp (mm) M pc / M pb Panel zone φ R / R v n u D7 B H H D7 S H H D9 B H H D9 S H H dp D7(9) B(S) (a) (b) t : Doubler plate thickness, nominal yield strength ( y ) (a) 7 and 9 stand for db of 692 and 89mm, respectively. (b) B and S stand for balanced and strong panel zones, respectively = R F with t cp : Continuity plate thickness, 1 using expected yield stress ( y y) F, 2 using tensile coupon yield stress ( yact, ) F (Table2) 6

8 Coulmn Lateral Restraint Hydraulic Actuator H428x47x2x35 Coulmn CJP E71T-1C E71T-1C 3 CJP E71T-1C 3 PL-652x314x8(SM49) Doubler plate PL-358x193.5x2(SM49) Continuity plate Plug Weld D16 E71T-1C Weld steel backing to 11 column 11 E71T-1C 3 Steel backing to remain remove weld tab D22 holes (for F1T) 3 E71T-1C 11 E71T-1C 11 PL-64x146x15(SS49) Shear tap 13 E71T-1C Strong Wall Beam 4 E71T-1C E71T-1C After root is cleaned and inspected E71T-1C Remove weld backing, Badk-gauge, runoff tab Strong Floor 925 E71T-1C Unit : mm CJP E71T-1C 8 H692x3x13x2 Beam 3 E71T-1C 3 13 E71T-1C 11 E71T-1C (a) Test setup (b) Details of specimens (D7-B) Figure 3. Test setting and details of specimens MATERIAL TESTS Material tests were conducted to investigate the mechanical properties of steel materials used for beams, columns, stiffeners, and doubler plates of connection specimens. For each component of the connection specimen, three round sectional-shaped coupons were taken. All rolled sections for beams and columns were made with SS4 and SM49 steels, respectively. The specified minimum yield strengths of SS4 and SM49 are 235 and 325MPa, respectively. Doubler plates of thickness of 8, 1, and 16mm, as well as continuity plates, were made from SM49. Figure 4 shows the stress-strain curves of SS4 and SM49. From the results of material tests, mean material strength and elongation were calculated, and summarized in Table 2. According to AISC 341-1, the specified minimum yield stress of steel to be used for members in which inelastic behavior is expected shall not exceed 5 ksi (345 MPa). In Commentary A.3 of AISC 341-1, material specifications for seismic application are specified as: (1) a pronounced stress strain plateau at the yield stress, (2) a large inelastic strain capability; tensile elongation of 2% or greater in a gauge with a length of 5 mm, (3) good weldability, and (4) for structural wide flange shapes, additional supplementary 7

9 requirements of ASTM A992 and ASTM A913 provide a limitation on the ratio of yield stress to tensile stress; namely, this ratio should not exceed.85. The steel materials used in this study satisfied the foregoing requirements for seismic application of structural material. First, as mentioned above, the steel materials SS4 and SM49 had respective minimum specified yield strengths of 235 and 325, smaller than the 345MPa requirement. Second, among the coupon test results, the highest ratios of yield stress to tensile stress for SS4 and SM49 were.71 and.68, respectively, which were smaller than a limiting value of.85. Third, the minimum tensile elongations for SS4 and SM49 were respectively 21% and 2%, which were greater than or equal to the 2% requirement. Finally, the stress strain curves for SS4 and SM49 included pronounced yield plateaus at the yield stress as shown in Figure (a) (b) stress (MPa) stress (Mpa) Fy RyFy Fy RyFy Fu RtFu 1 Fu RtFu SS4 SM strain strain Figure 4. Coupon test results: (a) SS4, (b) SM49 8

10 Table 2. Material properties Specimen Member Coupon 1Yield Strength (MPa) 2Tensile Strength (MPa) 3Yield Ratio= 1/2(%) Elongation (%) D7 B D7 S D9 B D9 S Beam H Column H Beam H Column H Beam H Column H Beam H Column H Flange Web Flange Web Flange Web Flange Web Flange Web Flange Web Flange Web Flange Web TEST SETTING AND TEST PROCEDURE The test setup and geometry of the specimen are shown in Figure 3a. The ends of the column were pin connected to simulate the point of inflection under a lateral loading condition. It was assumed that the column inflection points were located at the mid-height of a story. To prevent out of plane movement and twisting of the specimen, a lateral support was provided at a distance of 1.25m (1/3 of the beam length) from the loading point as shown in Figure 3a. Testing was conducted under displacement control, following a standard stepwise cyclic loading protocol based on Section K2 of AISC (AISC 21a). The cyclic loading was applied with a loading velocity of 1mm/s to the beam tip at a distance of 3.45m from the column centerline using a hydraulic actuator. 9

11 Strain gauges and displacement transducers were installed on each specimen according to Appendix F of SAC/BD 97/2. The specimens were whitewashed before testing to allow the visualization of yield patterns on the surface of the specimen. TEST RESULTS Figure 5 shows hysteretic curves for specimens. In this figure, the ordinate and abscissa represent M f M p and, respectively, where M f is the beam moment measured at the column face, M p is the beam plastic moment (= Z x F ye ), θ t is the displacement ( Δ ) at the beam tip divided by the beam length (L). Figure 6 shows the relationship between the panel zone deformation and M f M p. As expected, specimens with less panel zone strength (D7 B and D9 B) experienced greater panel zone deformation than corresponding specimens with stronger panel zone (D7 S and D9 S). Specimen D7 B, which had 692 mm beam depth and a balanced panel zone, experienced its first yielding in the beam bottom flange at a drift ratio of.75%. At a drift ratio of 1%, flaking of whitewash was noticed in the top and bottom flanges of the beam and the panel zone, indicating that yielding had occurred. At a drift ratio of 4%, the first local beam flange buckling was detected. After the loading cycle of 4% drift ratio, the strength of the specimen had deteriorated due to beam flange local buckling (Figure 5a). At the second cycle of 6% drift ratio, the specimen failed by a fracture at the toe of the weld access hole in the beam top flange (Figure 7a). Specimen D7 S, with a beam depth of 692 mm and a strong panel zone, behaved similarly to Specimen D7 B. At a drift ratio of.5%, yielding began in the beam bottom flange. At a drift of 3%, yielding was first detected in the panel zone, and local buckling occurred in beam flanges. Over the subsequent cycles before the drift ratio reached 6%, significant local buckling occurred in the beam flange, gradually reducing the connection strength; then during the first cycle of a drift ratio of 6%, the bottom flange fractured at the heat affected zone (HAZ) (Figure 7b). Specimens D7 B and D7 S had drift capacities greater than 4%. At a drift ratio of 4%, the strengths of the specimens exceeded 8% of the beam plastic moment (= FZ y x ). The specimens also 1

12 completed the two cycles at a drift ratio of 4%. Thus, Specimens D7 B and D7 S satisfied the requirement for SMF connections specified in AISC In Specimen D9 B, which had a beam depth of 89 mm and a balanced panel zone, the beam bottom flange near the column face started to yield at a drift ratio of.5%. Subsequently, the panel zone yielded at a drift ratio of 1%. After the first cycle at a drift ratio of 3%, local buckling occurred in the beam top and bottom flanges. At a drift ratio of 4%, a significant reduction in connection strength was observed due to local buckling. At the first cycle at a drift ratio of 4%, the specimen failed in the beam top flange near the access hole toe (Figure 7c). Normalized beam moment, M f /M p Normalized beam moment, M f /M p (a) D7-B Fracture of Beam Top Flange Panel Zone Yielding Normalized beam moment, M/M p Beam Yielding Normalized beam moment, M/M p Figure 5. Hysteretic curves of connection specimens (b) D7-S Fracture of Beam Bottom Flange Panel Zone Yielding Beam Yielding -1 Initiation of Flange -1 Initiation of Flange Local Buckling Local Buckling (c) D9-B Panel Zone Total story drift angle, θ t (radian) (d) D9-S Total story drift angle, θ t (radian) Yielding 1 1 Panel Zone Fracture of Beam Yielding Bottom Flange Beam Yielding Beam Yielding Fracture of Beam Top Flange -1 Initiation of Flange -1 Initiation of Flange Local Buckling Local Buckling Total story drift angle, θ t (radian) Total story drift angle, θ t (radian) 11

13 Normalized beam moment, M f /M p Normalized beam moment, M f /M p 1.5 (a) D7-B Normalized beam moment, M f /M p 1.5 (b) D7-S (c) D9-B Panel zone story drift angle, θ t (radian) (d) D9-S Panel zone story drift angle, θ t (radian) Normalized beam moment, M f /M p Panel zone story drift angle, θ t (radian) Panel zone story drift angle, θ t (radian) Figure 6. Relationship between M f Mp and panel zone deformation (a) D7 B (b)d7 S (c) D9 B (d)d9 S Figure 7. Fracture of specimens 12

14 Specimen D9 S, which had a beam depth of 89 mm and a strong panel zone, had similar cyclic behavior as Specimen D9 B. The specimen failed in the beam bottom flange near the access hole toe during the first positive cycle at a drift ratio of 4% (Figure 7d). Specimens D9 B and D9 S did not complete one cycle at a drift ratio of 4% that are required for qualifying SMF connections according to AISC As mentioned earlier, L n / d of Specimens D9-B and D9-S is 7.76, which is relatively low for SMF systems. b Note that the limiting value of L / d is 7. for the SMF system (AISC 2b). Roeder and n b Foutch (1996) reported that with a decrease in L / d, the rotation capacity of moment connections decreased. Figure 8 shows test specimens at the completion of the test. Specimens D7 B and D9 B with balanced panel zones had significant yielding throughout the whole area of their panel zones (most of the whitewash fell off), whereas Specimens D7 S and D9 S with strong panel zones had minor to moderate yielding in their panel zones, as further confirmed by their hysteretic curves in Figure 6. n b Figure 8. Pictures of WUF W specimens at the completion of testing 13

15 CONNECTION ROTATION CAPACITY AND STRENGTH Envelope curves were plotted in Figure 9 to compare drift capacities and strengths of specimens, which were extracted from the hysteretic curves in Figure 5. The results are summarized in Table 3, which includes total maximum story drift ratio, plastic rotation capacity, the ratio of maximum value of M f ( M f,max ) to M p, the contributions of the beam, the column and the panel zone to the connection plastic rotation capacity, and the location of the fracture. The total maximum story drift ratio and plastic rotation capacity were measured at the last completed cycle before connection fracture occurred. The plastic rotation was calculated using Eq. (5) according to SAC/BD 97/2: Δ P/ Kel θp = L + d /2 n c (5) where P is the force produced by the actuator and Kel is the elastic stiffness of the specimen. As shown in Table 3, for Specimens D7 B and D7 S, the values of M f,max M are p 1.43 and 1.38, respectively; for Specimens D9 B and D9 S, M,max f M are 1.41 and p 1.44, respectively. The average value of M,max f M p is 1.42, which is close to pr C of 1.4 specified in Section 8.7 of AISC According to Table 3, Specimens D9 B and D9 S, which had a beam depth of 89 mm, had a total maximum story drift ( θ t ) of 3% and a plastic rotation capacity ( θ p ) of 1.8%. These quantities did not exceed.4 for θ t and.3 forθ p, which are required by AISC On the contrary, Specimens D7 B and D7 S, which had a beam depth of 692mm, had θ t greater than 4% and θ p greater than 3%. 14

16 Table 3. Summary of test results Specimen D7 B D7 S D9 B D9 S Total story drift, (rad) Total plastic rotation, (rad) Panel zone maximum plastic rotation,, (rad) Beam maximum plastic rotation,, (rad) Column maximum plastic rotation,, (rad) M M f,max Location of fracture p 1.5 Beam top flange in beam HAZ Beam bottom flange in beam HAZ Beam top flange near access hole toe Beam bottom flange near access hole toe Normalized beam moment, M/M p D7B -1 D7S D9B D9-S Total story drift angle, θ t (radian) Figure 9. Envelop curves for connection specimens STRAIN DISTRIBUTION ON THE BEAM FLANGE The strain demand due to the beam depth and panel zone strength of connection specimens was investigated. Figure 1 shows the distribution of strain along the width of beam flanges at a drift ratio of 3%. The strain gauges of the beam flange were placed at 5 mm from the column face. The maximum strain demand of D7 B with a balanced panel zone was 15% less than that of D7 S with a strong panel zone, whereas the maximum strain demand of Specimen D9 B with a balanced pane zone was 25% larger than that of Specimen D9 S with a strong panel zone. At a drift ratio of 3%, Specimen D9 B experienced more significant yielding that was concentrated near the connection; this yielding induced a great deal of localized inelastic strain demand. Specimen D9 B failed at a drift ratio of 3%, whereas Specimen D9 S reached a drift ratio of 4%, but failed during the first cycle at this level. 15

17 For Specimen D9 S, the first crack detected was in the bottom flange at a drift ratio of 3%, as shown in Figure 11. During the first cycle at a drift ratio of 4%, the bottom flange fractured. As shown in Figure 1, the strains of the deeper beams ( d b = 89 mm) were generally greater than those of the shallower beams ( d b = 692 mm). For Specimen D9 B with a deeper beam and a balanced panel zone, the maximum strain ( ε max ) was.4 at the bottom flange whereas for Specimen D7 B with a shallower beam and a balanced panel zone, ε max was.2 in the top flange. As for specimens with strong panel zones, ε max in Specimen D9 S reached.32 in the bottom beam flange whereas ε max in Specimen D7 S was.23 in the top flange. The distribution of strains along the beam length was measured. The locations of strain gauges were shown in Figure 12. The greatest strain was measured from the strain gauge placed closest to the column face. As the location of strain gauges is greater distance from the column face, strain diminished. This indicates that yielding would initiate at the column face. Thus, for WUF-W connections, it is adequate to assume that the plastic hinge is located at the column face (Section 8.7 of AISC 358-1). With an increase in loading amplitude, the strain became greater, leading to an increase in the area of yielding. Figure 1. Strain demand in the beam flanges at a drift ratio of 3% 16

18 Column flange Partial expansion Beam bottom flange Initial crack Figure 11. Cracks in the bottom beam flange in Specimen D9 S at 3% drift loading cycle Figure 12. Strain distribution along the distance from the column face EFFECT OF WELD ACCESS HOLE DETAIL ON STRAIN DISTRIBUTION The WUF W specimens of a beam depth of 89mm tested in this study did not pass the acceptance criteria for SMF connections, whereas the WUF W specimens with a similar beam depth (W36 section) to those used by Ricles et al. (22) (Specimens T1 and T5) satisfied the SMF connection criteria. The total maximum story drift ratio and plastic rotation capacities of Specimen T1 were.5 rad and.35 rad, respectively, and Specimen T5 produced a total maximum story drift ratio of.5 and a plastic rotation of.35 rad. The most prominent difference between the specimens used by these two studies was the 17

19 geometry of the access hole, particularly the slope of the access hole to the beam flange. Table 4 summarizes the detailed dimensions of the access holes used in these two studies. Specimens T1 and had a lower slope of access hole to the beam flange than Specimens D9 B and D9 S. Note that in AISC and AWS D1.8/D1.8M-9, the dimension parameters for weld access hole configuration ( a through g in Table 4) are given as ranges rather than specific values. Nakashima et al. (1998) reported that the geometry of the weld access hole affects the stress and strain in the vicinity of the toe of the access hole. In 3D FEA conducted by Mao et al. (21), the inelastic strain demand decreased with a decrease in the slope of the access hole to the beam flange. Figure 13 shows the strain distribution in the beam top and bottom flanges of WUF W specimens at a drift ratio of 3%, as measured from strain gauges placed 5 mm from the column face. Specimens D-9-S and D9-B with an access hole slope of 21 experienced a greater strain demand than Specimens T1 and T5 with an access hole slope of 13. Table 4. Access hole dimensions b e f d a g tbf Specimen a(mm) b(mm) c(mm) d(mm) e(mm) f(mm) g ( ) D9 B, S (This study) T1, T5 (Ricles et al. 22) Figure 13. Comparison of strain distributions of WUF W connection specimens (T1, T5 of Ricles et al. (22); D9 B and D9 S of this study) 18

20 ENERGY DISSIPATION CAPACITIES The total amount of cumulative energy dissipated by each WUF W specimen was estimated. Table 5 summarizes the total cumulative energy dissipated by connection specimens and their component. Figures 14 and 15 illustrate the cumulative energy dissipated by each specimen at each drift loading cycle. In Specimen D9 B with a balanced panel zone, the energy dissipated by the panel zone is almost the same as that dissipated by the beam. In Specimen D7 B, the panel zone dissipated as much energy as the beam until a drift ratio of 3%. Beyond a drift ratio of 3%, the energy dissipated by the beam increased rapidly with an increase in drift ratio, whereas the amount of energy dissipated by the panel zone was nearly constant regardless of the amplitude of drift ratio. The rapid increase in energy dissipated by the beam was attributed to the occurrence of significant local buckling at drift ratios greater than 3%. The total amount of energy dissipated by Specimen D9 B was 5% less than that dissipated by Specimen D7 B. For Specimens D9 S and D7 S, which had strong panel zones, most energy was dissipated by the beams throughout the test. The total amount of energy dissipated by Specimen D9 S was 67% of that dissipated by Specimen D7 S. Among the tested specimens, D7 B dissipated the most energy, whereas Specimen D9 B dissipated the least. Note that specimens with a beam depth of 692 mm survived more loading cycles with high drift ratios than specimens of beam depth 89 mm. Table 5. Energy dissipation capacities of overall specimens and of their components Specimen Total dissipated energy(kj) Energy dissipated by beam (kj) Energy dissipated by panel zone (kj) Energy dissipated by column (kj) D7 B (72%) 42 (24%) 64 (4%) D7 S (88%) 91 (7%) 72 (5%) D9 B (52%) 397 (48%) 2 (%) D9 S (8%) 167 (19%) 1 (1%) 19

21 Cumulative Dissipated Energy (kj) Cumulative Dissipated Energy (kj) % 1% 2% 3% 4% 5% 6% 7% % 1% 2% 3% 4% 5% 6% 7% Story Drift Ratio (%) Story Drift Ratio (%) Figure 14. Cumulative energy dissipated by Specimens D7 B and D7 S at each drift loading cycle Cumulative Dissipated Energy (kj) (a) Specimen D7-B (a) Specimen D9-B Cumulative Dissipated Energy (kj) Figure 15. Cumulative energy dissipated by Specimens D9 B and D9 S at each drift loading cycle CONCLUSIONS Cyclic tests were conducted on four WUF W specimens designed and detailed according to AISC and AISC The test variables were the beam depth and the panel zone strength ratio. From the test results, the following conclusions were made: (1) In this study, WUF W specimens with deeper beam depths were incapable of accommodating a drift capacity of 4% required for SMF connections specified in Section E3.6b.(1) of AISC The specimens failed by beam flange fracture near the access hole toe. (2) Test results showed that the maximum strain of specimens with a beam depth of 89 mm was greater than that of specimens with a beam depth of 692 mm, which increased the potential of failure in specimens with a deeper beam depth. Note that all specimens had the same span length (b) Specimen D7-S (b) Specimen D9-S % 1% 2% 3% 4% 5% 6% 7% % 1% 2% 3% 4% 5% 6% 7% Story Drift Ratio (%) Story Drift Ratio (%) 2

22 (3) Among specimens with a beam depth of 692 mm, Specimen D7 S with a strong panel zone had a lower drift capacity than D7 B with a balanced panel zone. However, in the case of specimens with a beam depth of 89 mm, Specimen D9 S with a strong panel zone had a slightly greater drift capacity than Specimen D9 B with a balanced panel zone. Considering the test results of this study, the drift capacity of WUF-W connections did not vary consistently according to panel zone strength ratio. (4) Specimens D9-S and D9-B tested in this study did not pass the acceptance criteria for SMF connections, whereas Specimens T1 and T5 with similar beam depths tested by Ricles et al. (22) satisfied the criteria for SMF connection. The main difference between the specimens used in these two studies was the access hole slope. Specimens D9-S and D9-B had 21 o whereas Specimens T1 and T5 had an access hole slope of 13 o. It is noted that an access hole slope of 21 o is smaller than the limiting value of access hole slopes specified in AWS D1.8/D1.8M-9, which is 25 o. This suggested that the limiting value of access hole slopes should be more stringent to reduce beam strain demands. ACKNOWLEDGEMENTS The research described in this paper was supported by grants from the National Research Foundation of Korea (No. 212R1A2A2A645129). The valuable comments offered by reviewers are also greatly appreciated. REFERENCES American Institute of Steel Construction, 21a. Seismic Provisions for Structural Steel Buildings, ANSI/AISC 341 1, Chicago, IL. American Institute of Steel Construction, 21b. Prequalified Connections for Special and Intermediate Steel Moment Frames for Seismic Applications, ANSI/AISC 358 1, Chicago, IL. American Institute of Steel Construction, 21c. Specification for Structural Steel Buildings, ANSI/AISC 36 1, Chicago, IL. American Society of Civil Engineers, 21. Minimum Design Loads for Buildings and Other Structures, ASCE/SEI 7 1, Reston, VA. American Welding Society, 29. Structural Welding Code Seismic Supplement, AWS D1.8/D1.8M, Miami, FL. Anderson, J. C., and Linderman, R.R., 1991, Post Earthquake Repair of Welded Moment Connections, CE 91 4, Department of Civil Engineering, University of Southern California. 21

23 Engelhardt, M. D., and Husain, A. S., 1993, Cyclic loading performance of Welded Flange Bolted Web connections, ASCE Journal of Structural Engineering 119(12), Han, S. W., Kwon, G. U., and Moon, K. H, 27. Cyclic behavior of post Northridge WUF B connections, Journal of Constructional Steel Research 63(3), Lu, L. W., Ricles, J. M., Mao, C., and Fisher, J. W., 2. Critical issues in achieving ductile behavior of welded moment connections, Journal of Constructional Steel Research 55, Lee, D., Cotton, S. C., Hajjar, J. F., Dexter, R. J., and Ye, Y., 25a. Cyclic behavior of steel moment resisting connections reinforced by alternative column stiffener details I. connection performance and continuity plate detailing, AISC Engineering Journal 42(4), Lee, D., Cotton, S. C., Hajjar, J. F., Dexter, R. J., and Ye, Y., 25b. Cyclic behavior of steel moment resisting connections reinforced by alternative column stiffener details II. panel zone behavior and doubler plate detailing, AISC Engineering Journal 42(4), Lee, K., Foutch, D.A., 22. Seismic performance evaluation of pre Northridge steel frame buildings with brittle connections, ASCE Journal of Structural Engineering 128(4), Mao, C., Ricles, J., Lu, L. W., and Fisher, J., 21. Effect of local details on ductility of welded moment connections, ASCE Journal of Structural Engineering 127(9), Nakashima, M., Suita. K., Morisako, K., and Maruoka, Y., Tests of welded beam column subassemblies I: global behavior, ASCE Journal of Structural Engineering 124(11), Popov, E. P., and Stephen, R. M., Cyclic Loading of Full Size Steel Connections. Steel Research for Construction, Bulletin No. 21, AISI. Popov, E. P., Amin, N. R., Louie, J. J. C., and Stephen, R. M., Cyclic behavior of large beam column assemblies. Earthquake Spectra, Earthquake Engineering Research Institute 1(2), Ricles, J. M., Mao, C., Lu, L. W., and Fisher, J. W., 2. Development and Evaluation of Improved Details for Ductile Welded Unreinforced Flange Connections, SAC Joint Venture, Report No. SAC/BD /24, Sacramento, CA. Ricles, J. M., Mao, C., Lu, L. W., and Fisher, J. W, 22. Inelastic cyclic testing of welded unreinforced moment connections, ASCE Journal of Structural Engineering 128(4), Roeder, C. W., and Foutch, D. A., Experimental results for seismic resistant steel moment frame connections, ASCE Journal of Structural Engineering 122(6), SAC Joint Venture Protocol for Fabrication, Inspection, Testing, and Documentation of Beam Column Connection Tests and Other Experimental Specimens, Report No. SAC/BD 97/2, Sacramento, CA. 22

24 Stojadinovic, B., Goel, S. C., Lee, K. H., Margarian., A. G., and Choi, J. H., 2. Parametric tests on unreinforced steel moment connections. ASCE Journal of Structural Engineering 126(1), Tsai, K. C., and Popov., E. P., Steel Beam Column Joints in Seismic Moment Resisting Frames. Earthquake Engineering Research Center, UCB/EERC 88/19, Richmond, CA. 23