Interior wide beam connections subjected to lateral earthquake loading

Engineering Structures 25 (2003) 281–291 www.elsevier.com/locate/engstruct

Interior wide beam connections subjected to lateral earthquake loading W.L. Siah, J.S. Stehle, P. Mendis ∗, H. Goldsworthy Department of Civil and Environmental Engineering, University of Melbourne, 3010 Parkville, VC, Australia Received 4 October 2001; received in revised form 28 August 2002; accepted 30 August 2002

Abstract The use of wide beam–column systems has become increasingly popular in low to moderate seismic regions, despite very little information being available on their performance under seismic action. Therefore, this investigation was conducted to determine the behavior and likely failure mode of wide beam connections when subjected to earthquake loading. In this investigation, two interior reinforced concrete wide beam subassemblages and one post-tensioned concrete wide beam subassemblage were tested under quasi-static cyclic loading up to a drift ratio of 3.5%. It was found that the wide beam connection is likely to experience severe torsion cracking in the beam portions located at the sides of the column when subjected to severe earthquake loading. A special detailing strategy was developed to inhibit the torsion cracking and was found to be effective in the subsequent tests. The experimental behavior of the subassemblages is reported in this paper.  2002 Elsevier Science Ltd. All rights reserved. Keywords: Wide beams; Joints; Earthquake-resistant structures; Torsion; Debonding

1. Introduction In non-seismic regions, the wide beam–column system has been very popular as a gravity load resisting system (GLRS). The main advantage of this system is the reduction in interstorey height when compared to normal-width beam–column systems. However, when used as a GLRS in seismic regions, the wide beam–column connections are likely to experience lateral deformations, and the performance of the connections under such behavior has not been fully understood. It has been recognized that a GLRS should have sufficient deformation capacity to move laterally at each level by the same amount as the lateral load resisting system (LLRS) [1,2]. Otherwise, damage is likely to occur at the connection and may lead to failure of the system. In the past, only a few investigations have been conducted on the performance of interior wide beam connections under seismic action [2–5]. However, those

investigations were limited to subassemblages with beam-width to column-width ratio (bw/bc) of less than 3.5, whereas wide beam–column systems with beamwidth to column-width ratios up to 4.8 are commonly used in practice in countries such as Australia. Current code provisions to control the use of wide beams in regions of high seismicity are aimed at limiting the width of the wide beam. The limit provided by ACI318 [6] is set at (bc ⫹ 1.5hb), which is based on standard design practice, not experimental studies [7,8]. NZS3101 [9] also provides a limit on the width of the beam, which is equal to the smaller of (bc ⫹ 0.5hc) and 2bc. Based on experimental studies, some researchers [2,4,7] have also proposed limits on wide beam widths, which are dependent on the depth or width of the column.

2. Experimental program 2.1. Prototype and specimen design

Corresponding author. Tel.: +61-3-83-44-72-44; fax: +61-3-8344-46-16. E-mail address: [email protected] (P. Mendis). ∗

The experimental part of this research involved the testing of three wide beam specimens, i.e. two reinforced concrete (RC) wide beam connections and one post-ten-

0141-0296/03/$ - see front matter  2002 Elsevier Science Ltd. All rights reserved. doi:10.1016/S0141-0296(02)00150-5

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Nomenclature q Ap bc bw db dom f⬘c fpu fpy fsy fy h hb hc heq Ld Mr

concrete strut / torsion crack angle area of active reinforcement width of column width of beam diameter of reinforcement effective depth of the slab averaged around the perimeter concrete compressive strength ultimate tensile strength of active reinforcement yield strength of active reinforcement yield strength of passive reinforcement reinforcement yield strength depth of section (beam or column) depth of beam depth of column equivalent viscous damping ratio debonded length column-to-beam moment capacity ratio

sioned concrete (PC) wide beam connection. All three specimens (designated specimen WBB-I1 to WBB-I3) represented half-scale models of the prototype. The prototype structure is a four storey, six bay moment resisting frame, as shown in Fig. 1. The floor system consisted of wide beams in the long direction and oneway RC slabs in the transverse direction. The design gravity loading on each floor consisted of beam and slab self weight plus 1 kPa superimposed dead load and 5 kPa live load, which is typical for office buildings. The lateral earthquake loading was based on Australian loading codes [10,11]. The prototype was designed as an ordinary moment resisting frame (OMRF), according to the Australian Concrete Code AS3600 [12], with details similar to a non-ductile RC frame in California, USA. For the PC wide beam connection, the prestressing was designed to balance 60% of the dead load. The effective concrete prestress level was 1.5 MPa. The half-scale test specimens were terminated at column mid-height and beam mid-span, corresponding to points of inflection of the bending moment diagram

Fig. 1.

Elevation of prototype structure.

under lateral loading. The overall dimensions of the specimens are shown in Fig. 2. The specimens had wide beams 200 mm deep and 1200 mm wide, and 250 mm square columns. The top columns and bottom columns had lengths of 750 and 950 mm, respectively. It should be noted that there were no transverse beams or slabs involved in the test specimen. The specimens had a beam-width to column-width ratio of 4.8. The nominal concrete strength of the columns and the beams were 40 and 32 MPa, respectively. The first specimen (WBB-I1) is a RC wide beam connection commonly used in practice and was detailed, as shown in Fig. 3. As for the second specimen (WBB-I2), it is similar to the first specimen except that a special detailing strategy (debonding strategy—explained later in this paper) was implemented to overcome a problem observed in the first test, i.e., torsion cracking in the outside beam portion. The bottom beam bars in the second specimen were also continuous through the joint, unlike that for the first specimen. The reinforcement details of

Fig. 2.

Dimension of test specimen.

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Fig. 3.

283

Details of specimen WBB-I1.

Fig. 5. Details of specimen WBB-I3.

the second specimen are shown in Fig. 4. The third specimen (WBB-I3) is similar to the first and second specimens except that it had a PC wide beam instead of a RC wide beam. The column for the third specimen was stronger than that for the first and second specimens. The debonding strategy was also adopted in the third specimen to determine the effectiveness of the strategy for PC connections. The reinforcement details of the third specimen and its cable profile are shown in Figs.

Fig. 4.

Details of specimen WBB-I2.

5 and 6 respectively. The location of the debonding region is shown in Fig. 7, in which the outside beam reinforcement further away than 100 mm from each column side face was debonded, for a length of 650 mm. The materials used in the specimens were all tested in the laboratory and the average yield and ultimate strengths are presented in Table 1. Key specimen properties evaluated using measured material properties are given in Table 2. It should be noted that the specimen beam width of 1200 mm is much greater than the beam width limits set by ACI-318 and NZS3101, which were found to be 550 and 375 mm, respectively. Not more than one-fourth of the beam reinforcement was anchored through the core in each specimen. The joint shear stress ratio (g) was determined assuming reinforcement yield strength of 1.0fy. Two values of joint shear stress ratio are given for each specimen, corresponding to the values determined using the effective joint shear area as pro-

Fig. 6.

Cable profile of specimen WBB-I3.

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Fig. 7.

Location of debonding region.

Table 1 Material properties Specimen

WBB-I1 (MPa)

WBB-I2 (MPa)

WBB-I3 (MPa)

Column Beam R6 R10 Y12 Y16 θ12.7

41 47 Yield: 360, ultimate: 511 Yield: 350, ultimate: 504 Yield: 444, ultimate: 535 – –

41 38

51 42 Yield: Yield: Yield: Yield: Yield:

a

477, ultimate: 575 311, ultimate: 483 442, ultimate: 537 430, ultimate: 527 1605a, ultimate: 1888

Taken as 0.85fpu based on AS3600 [12] and NZS3101 [9].

Table 2 Specimen properties Specimen

WBB-I1

WBB-I2

WBB-I3

bw × hb bc × hc Beam-width to column-width ratio Beam reinforcement anchored in corea

1200 × 200mm2 250 × 250mm2 4.8 20% 25% 0.86 16.7 20.8 – 1.7 2.5 –

20% 25% 1.31 16.7 20.8 – 2.0 3.1 1.5 MPa

9%b 25%

Column-to-beam moment capacity ratio (Mr) Reinforcement bond parameter (h/db)

Joint shear stress ratio (g) Effective prestress level a b c d

Top Bottom 0.86 Column bars Beam bars Beam strands NZS3101 joint shear areac ACI-318 joint shear aread –

12.5 20.8 20.0 1.65 2.5

Active reinforcement assumed to be on the top, as would be the case near the joint. Active reinforcement based on reinforcement yield strength ratio: Ap(equivalent) ⫽ (fpy / fsy)Ap⬁. For bb ⬎ bc, effective joint shear width=smaller of bc ⫹ 0.5hc, bb, 2bc. For bb ⬎ bc, effective joint shear width=bc.

vided by ACI-318 [6] and NZS3101 [9]. It should be noted that the limit provided by ACI-318 is 1.66 (SI) for an interior joint. 2.2. Test program 2.2.1. Test setup The wide beam specimens were tested in the setup as shown in Fig. 8. The specimens were pinned at the beam

and column ends to simulate points of inflection under lateral loading. The pins at the column ends were provided using mild steel rods through holes on the test frame (at the top of the column) and actuator arm (at the bottom of the column). The pins at the beam-ends were created using vertical links hanging down from the test frame. These vertical links allowed free horizontal movement and no vertical movement, similar to a roller. An actuator was used to displace the bottom of the

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Fig. 8.

285

Schematic drawing of test setup.

column horizontally to simulate the effects of lateral loading due to an earthquake. The loading sequence of the test is shown in Fig. 9. It should be noted that the same loading sequence was used for all three specimens. Axial load in the columns (approximately 0.2f⬘c) was simulated by externally prestressing the columns, and this axial load remained concentric with the column throughout testing. The initial moment at the joint (due to gravity loading) was simulated in the specimens so that it would be consistent with the prototype structure. A photo of specimen WBB-I3 in the test frame is shown in Fig. 10. 2.2.2. Instrumentation and data collection Strain gauges, load cells and displacement transducers were used to monitor the behavior of the specimens during the test. All data was continuously recorded by a data acquisition system whenever the actuator was being used to displace the specimen. Strain gauges were used to measure the strain of the column and beam reinforcement at and near the connection region. There were a total of approximately 75 strain gauges in each specimen. Load cells were used to determine the load in each column-prestressing strand. Displacement transducers were used to monitor the horizontal displacement of the frame relative to the ground.

Fig. 9.

Loading sequence.

Fig. 10. Photo of test frame with specimen WBB-I3.

2.2.3. Debonding strategy The debonding strategy was developed to inhibit torsion cracking in the beam portions located at the sides of the column. (For the purpose of this paper, this region of the wide beam is termed the ‘outside beam portion’.) The torsion cracks occurred as a result of outside beam reinforcement transferring unbalanced moment to the joint through torsion. However, it should be noted that some of the beam reinforcement passing outside the column is able to transfer load to the joint through the formation of concrete struts. The regions in which unbalanced moment is transferred to the joint through torsion and concrete struts are simply demarcated as shown in Fig. 11. The angle q can be determined using strut-andtie modeling. With the debonding strategy, outside beam reinforcement that is expected to induce torsion is debonded for a particular length. With debonding, loads from that beam reinforcement are transferred to the joint through an alternative load path to that of torsion, as

Fig. 11. Transfer of unbalanced moment at wide beam connection (without debonding).

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Fig. 12.

Alternative load path with debonding strategy. Fig. 14. Load versus drift ratio behavior of specimen WBB-I2.

Fig. 13.

Load versus drift ratio behavior of specimen WBB-I1.

shown in Fig. 12. More information on the debonding strategy is available elsewhere [13].

3. Discussion of results 3.1. Overall specimen response The wide beam specimens were tested to a maximum 4% nominal drift ratio, when the maximum stroke of the actuator arm was reached. However, due to issues such as frame deformation and slipping of pins at the top and base of the columns, the recorded hysteretic responses had to be corrected. The corrected hysteretic responses of the three specimens are shown in Figs. 13–15. From Fig. 13, it can be seen that specimen WBB-I1 did not reach its capacity, (i.e. expected to fail through

Fig. 15.

Load versus drift ratio behavior of Specimen WBB-I3.

column hinging since Mr—column-to-beam moment capacity ratio, is less than 1). This is mainly due to the fact that severe torsion cracking was present in the outside beam portion, with a maximum crack width of 2 mm. At the later cycles of the test, it can be seen that the specimen were not carrying any higher loads. As for specimen WBB-I2, the capacity of the specimen (i.e., hinging of the bottom column) was just reached as can be seen from Fig. 14. This shows that the debonding strategy adopted in specimen WBB-I2 has improved the performance of the wide beam connection. For specimen WBB-I3, the column was much stronger than the beam, with a Mr value of 1.31. At the end of the test, the capacity of specimen WBB-I3 was not reached as can be seen from Fig. 15, and the specimen would still take higher load if more displacement was applied. Once again, this good behavior is attributed to the debonding strategy. The post-test condition of specimens WBB-I2 and WBB-I3 was very good with only minor damage (mainly flexural cracks) despite reaching high drift levels. The hysteretic response envelope of the three specimens is shown in Fig. 16. Comparing specimen WBBI2 (debonded RC wide beam) and specimen WBB-I3 (debonded PC wide beam), it can be seen that specimen WBB-I3 is stiffer, as expected due to prestressing. The experimentally attained actions and the design capacities based on measured material properties are presented in Table 3.

Fig. 16.

Hysteretic response envelope of the three specimens.

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attain higher displacement ductility if it had been possible to apply a higher level of displacement. Generally, the wide beam–column specimens were found to possess low energy dissipation capacity. However, with the debonding strategy adopted, the connections would be expected to experience little damage, high displacement capacity and small residual drift, as was found for specimens WBB-I2 and WBB-I3. 3.3. Strain behavior

Fig. 17.

Equivalent viscous damping ratios.

3.2. Energy dissipation capacity The energy dissipation capacity of the specimens can be determined by finding the equivalent viscous damping ratio (heq) [2]. The heq values for the specimens are presented in Fig. 17. The increase in heq was found to correlate well with initiation and widening of cracks, and yielding of reinforcement. It can be seen from Fig. 17 that specimen WBB-I1 experienced the highest heq value. This is due to the fact that more damage occurred in that specimen. Maximum heq attained in the test was approximately 9% for the three specimens. This low value of heq (as compared to normal-width beam connections—typically ~15% [2]) and the narrow loops from the hysteretic responses indicate that the concrete wide beams in general had low energy dissipation capacity. The overall system displacement ductility of the specimens was determined using m ⫽ ⌬max / ⌬y, where, ⌬max is the maximum displacement and ⌬y, the displacement at yield. For the purpose of this investigation, the displacement at first yield was defined as the displacement corresponding to a sudden change in the stiffness of the specimen. A minimum displacement ductility factor of 3 was calculated based on the maximum displacement applied during testing. It should be noted that specimens WBB-I2 and WBB-I3 would have been expected to

First yielding of reinforcement was observed at a drift ratio of 1.6 and 0.8% for specimens WBB-I2 and WBBI3, respectively. Yielding of reinforcement was not observed for specimen WBB-I1 because of the severe torsion cracks that led to premature failure of the connection. The strain distribution across the wide beam at 3.5% drift ratio due to applied loads for the three specimens is shown in Fig. 18. As can be seen, reinforcement in the central portion of the wide beam experienced greater strain than that of the outside beam portions. This is expected due to the indirect and flexible load path of the outside beam reinforcement. Despite the beam reinforcement being debonded, some of the debonded bars were able to reach yield by the end of the test. The effectiveness of the debonding strategy can be determined by comparing the strain behavior of a beam

Fig. 18. Strain distribution across wide beam at 3.5% drift ratio (top reinforcement in tension).

Table 3 Experimentally attained actions and member strengths Design parameter

Bottom column moment (kNm) Bottom column shear (kN) Top column moment (kNm) Top column shear (kN) Beam (at column face) Negative moment (kNm) Positive moment (kNm) a

WBB-I1

WBB-I2

WBB-I3

Action

Strength

Action

Strength

Action

Strength

76.5 80.5 60.5 80.5

86 188 78 174

83.5 88 66 88

86 188 78 174

95 100 75 100

139 182 137 177

77.5 39

165 29a

98.5 41.5

162 73

110 57

144 66

Yield strength of longitudinal reinforcement taken as 0.4fy due to poor bottom bar anchorage.

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Fig. 21. Torsion cracks in WBB-I1. The lines were thickened for easy identification.

3.4. Wide beam torsion cracking

Fig. 19.

Strain plots of a beam bar.

bar (or strand) at two different locations, but within the same debonded length. If the bar or strand is properly debonded, strain within the debonded length would be the same. The strain plots of a debonded beam bar in specimen WBB-I2 are shown in Fig. 19. The two plots are identical. Hence it can be concluded that debonding was effective. It should be noted that the same behavior was observed for the debonded bars and strands in specimen WBB-I3. The shear stresses in the wide beams were found to be relatively low, as can be seen from the strain gauge reading on a vertical leg of a beam stirrup (specimen WBB-I3) as shown in Fig. 20. The beam stirrup was located 100 mm (hb/2) from the column face. Moreover, the wide beams did not exhibit any inclined cracking that could be characterized as a shear crack. The low beam shear stress is typical for wide beams, as was found in other investigations on wide beam connections.

Fig. 20.

Strain plot of a vertical leg on a beam stirrup.

As mentioned previously, severe torsion cracking was observed in specimen WBB-I1. This cracking was attributed to the high torsional moment that was induced by the beam bars located away from the column side faces. Full torsion cracking (i.e., torsion cracks extending to the sides of the beam) was observed at a drift ratio of 2.17%. The maximum torsion crack width was 2 mm. The torsion cracks on specimen WBB-I1 are shown in Fig. 21. It should be noted that torsion cracking was not observed in specimen WBB-I2, as a result of the debonding strategy. This observation verified the success of the debonding strategy. As for specimen WBB-I3, local torsion cracking in the region close to the column was observed, as shown in Fig. 22. These cracks were not expected since the debonding strategy was implemented. Nevertheless, the debonding strategy has definitely inhibited severe torsion cracking in specimen WBB-I3 based on the following observations. (1) Torsional cracking at the top and bottom surfaces of the wide beam was not observed until 2.1% drift ratio, as compared to

Fig. 22. Torsion cracks in WBB-I3. The lines were thickened for easy identification.

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the occurrence of full torsion cracking in specimen WBB-I1 at 2.17% drift ratio. (2) The torsion cracks were only present close to the joint region and did not extend to the sides of the beam. (3) The maximum torsion crack width was 0.5 mm as compared to 2 mm for specimen WBB-I1. It should also be noted that the first debonded reinforcement was located 215 mm from the column side faces, and the torsion cracks only extended to a maximum of 180 mm from each column side face. The torsion cracks did not extend into the debonding region, indicating that debonding has inhibited full torsion cracking. Without debonding, full torsion cracking would be expected to occur in specimen WBB-I3 at a drift ratio of about 1%. From this investigation, it was found that prestressing has the effect of altering the torsion crack angle. The observed torsion crack angle was approximately 45° for WBB-I1 (RC) and approximately 40° for WBB-I3 (PC). With prestressing in the wide beam longitudinal direction, cracks are prevented from opening in that direction (i.e., beam longitudinal direction). Hence, the torsional crack will be inclined closer to the beam longitudinal axis, which means a reduction in the torsion crack angle with respect to the column side face. This is illustrated in Fig. 23. The occurrence of severe torsion cracking in the outside beam portion should be prevented since the beam portions at the sides of the column have been identified as critical areas which contribute to the joint shear strength [2,7] and punching shear strength [13,14]. Therefore, Eqs. (1)–(3) are proposed to predict the torsional performance of interior concrete wide beam connections under seismic actions. The circular interaction between shear and torsion suggested in (1) is consistent with the experimental observations made by Hsu [15]. It should be noted that presently there are no code requirements for the assessment of wide beam torsion cracking.

Fig. 23.

Reduction in torsion crack angle with prestressing.

289

冉 冊 冉 冊 T1 Tuc

2

⫹

V1 2 ⱕ1 Vuc

(1)

where, V1 is the total applied shear at the joint; Vuc, the punching shear strength of concrete with no moment transfer; T1, the applied torsion (assumed to be induced by reinforcement located a distance hctanq away from each column side face); Tuc, the torsion strength of a torsional strip of dimensions x and y (see Eq. (3)); q, the concrete strut angle (can be determined using strutand-tie modeling). Vuc ⫽ udom(0.34冑f⬘c ⫹ 0.3scp

(2)

where, u is the effective length of critical shear perimeter, as per AS3600 [12] (Fig. 24); dom, the effective depth of the slab averaged around the perimeter; scp, the average intensity of effective prestress in concrete in both directions. Tuc ⫽ 0.17冑f⬘cx2y

冪1 ⫹

10scp f⬘c

(3)

(for one column side face only) where, x is the maximum of hc and hb (without debonding) or minimum of Ld and hb (with debonding); y, the maximum of hc and hb (without debonding) or maximum of Ld and hb (with debonding). The values obtained from the proposed formulae were found to correlate well with experimental observations from the three specimens. The applied actions and capacities of each specimen are evaluated using the proposed formulae and are presented in Table 4. It should be noted that for specimens WBB-I1 and WBB-I3, the applied torsion (T1) provided is the value when torsion cracking was first observed in the beam top and bottom surfaces. As for specimen WBB-I2, T1 is obtained from the maximum applied torsion attained in the test, since torsion cracks were not observed. The respective applied torsions were determined from strain gauge values on the beam reinforcement. It should be noted that a linear

Fig. 24. Critical shear perimeter.

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Table 4 Action and capacity using the proposed wide beam formulae Specimen

WBB-I1

WBB-I2

WBB-I3

Angle θ adopted T1 (kNm) Tuc (kNm) T1 / Tuc V1 (kN) Vuc (kN) V1 / Vuc (T1 / Tuc)2 ⫹ (V1 /Vuc)2 (T1 / Tuc) ⫹ (V1 / Vuc)

45° 12.5 11.7 1.07 133 688 0.19 1.18 1.26

45° 23.8 27.2 0.88 133 619 0.21 0.82 1.09

40° 30.8 28.7 1.07 63 683 0.09 1.15 1.16

cracking in interior wide beam connections is provided and was found to correlate well with experimental observations. 6. With the information gathered from this investigation, the practicing engineer has more options regarding the following two issues: 앫 Debonding outside beam reinforcement if necessary. 앫 Using the wide beam system as a gravity load resisting system in a seismic region, as a secondary lateral load resisting system, or even a primary lateral load resisting system, depending on the expected level of seismicity.

interaction between shear and torsion is adopted in AS3600 [12], while deriving the punching shear capacity in the presence of an unbalanced moment. A similar relationship for wide beams is presented in (4). The values obtained using (4) are presented in Table 4. T1 V1 ⫹ ⱕ1 Tuc Vuc

(4)

Further details of the experimental method, results, and developed theory can be found in theses related to this work [16,17].

4. Conclusions Based on the test results described in this paper, the following conclusions can be made. 1. Severe torsion cracking in the outside beam portion located at the sides of the column has been found to be a problem for wide beam connections subjected to high lateral displacement. 2. Debonding of the outside beam reinforcement was found to improve the seismic performance and displacement capacity of a wide beam connection by inhibiting torsion cracking at the outside beam portions. The debonded connections were able to withstand up to 3.5% drift ratio and they did not experience any strength degradation. It should be noted that a debonded length of 650 mm (or hc+2hb) provided this satisfactory result. 3. The debonded interior concrete wide beam connections generally have low energy absorption capacity. However, they have high displacement capacity and low residual drift, and experience minor damage. 4. There is generally no major difference in the seismic performance of debonded interior RC and PC wide beam connections except for the fact that the PC wide beam connections are stiffer and prestressing reduces the torsion crack angle, with respect to the column side face. 5. A method to determine the likelihood of torsion

Acknowledgements The authors would like to acknowledge the financial support received from the Australian Research Council under grant numbers 896018878 and S8008742, and the support of all the people who have contributed to this investigation. Special thanks are given to Structural Systems Ltd. for the post-tensioning work that was performed as an in-kind contribution to the project.

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