INHERENT DUCTILITY OF REINFORCED CONCRETE SHEAR WALLS WITH NON-SEISMIC DETAILING

INHERENT DUCTILITY OF REINFORCED CONCRETE SHEAR WALLS WITH NON-SEISMIC DETAILING J. S. Kuang*, Hong Kong University of Science and Technology, Hong Kong Y. B. Ho, Hong Kong University of Science and Technology, Hong Kong 3st Conference on OUR WORLD IN CONCRETE & STRUCTURES: 6-7 August 26, Singapore Article Online Id: 33 The online version of this article can be found at: http://cipremier.com/33 This article is brought to you with the support of Singapore Concrete Institute www.scinst.org.sg All Rights reserved for CI Premier PTE LTD You are not Allowed to re distribute or re sale the article in any format without written approval of CI Premier PTE LTD Visit Our Website for more information www.cipremier.com

3 st Conference on OUR WORLD IN CONCRETE & STRUCTURES: 6 7 August 26, Singapore INHERENT DUCTILITY OF REINFORCED CONCRETE SHEAR WALLS WITH NON-SEISMIC DETAILING J. S. Kuang*, Hong Kong University of Science and Technology, Hong Kong Y. B. Ho, Hong Kong University of Science and Technology, Hong Kong ABSTRACT Large-scale reinforced concrete low-rise shear walls designed without seismic consideration, as practised in low probability of seismic occurrence regions, are tested under reversed cyclic loading. The seismic behaviour and inherent ductility of shear walls are experimentally established by testing the non-seismically designed shear-wall specimens in the case where the reinforcement detailing is consistent with that typically adopted for shear-wall buildings in which the design code of practice has not stipulated any requirement for seismic compliance in steel details. The experimental study reveals that the inherent displacement ductility factors of 2.5 to 3 may generally achieved for the shear walls with non-seismic reinforcement detailing. By correlating the available ductility with the required ductility demands, it is indicated that ordinary low-rise shear walls with non-seismic design and detailing may not possess the ductility to adequately respond to an unexpected low-to-moderate earthquake, then not sufficiently satisfy the ductility demand for shear-wall building structures in regions of low to moderate seismicity, including Hong Kong, Singapore, the UK, Central America and many other regions in the world. Modification in reinforcement detailing specifications for non-seismic design should be made in order for improving the seismic performance and enhancing the ductility and energy dissipation capacity of non-seismically designed and detailed reinforced concrete shear walls in regions of low to moderate seismicity. Further experimental studies on improving the ductility of non-seismically designed shear-wall structures are being carried out for moderate seismicity. Keywords: Reinforced concrete shear wall, ductility, low to moderate seismicity, seismic design. Introduction The medium earthquake (M = 5.6) occurred in Newcastle, Australia in 99, which caused about 2.5 billion US dollars of damage [3], in fact revives our attention to the potential hazard of non-seismically designed and detailed RC buildings in regions of moderate seismicity. The consequence of this earthquake reveals that although the seismic intensity of the earthquake is not high, it may cause a significant loss of life and economy in a region of low to moderate seismicity, where there is no urban earthquake disaster management programme at all. In low to moderate seismicity regions, such as Hong Kong and Singapore, which has a peak ground acceleration of about.g-.5g with a % probability of exceedance in 5 years, and many other parts of Asia, Europe and North America, engineers normally do not include seismic considerations in buildings design and detailing. A structure which is designed and detailed based

only on gravity load and/or other incidental horizontal loads would have to rely on its inherent ductility to respond acceptably to unexpected seismic excitations [4, 2]. Experimental studies on the seismic response evaluation of existing frame buildings designed for only gravity load [8, 9] have indicated that the displacement ductility demand on a framed structure under low to moderate earthquakes could range between 2. and 3.5. An investigation on the ductility demands on reinforced concrete moment-resisting frames for moderate seismicity has been reported [7]. The seismic responses of these as-built structures are considered when subjected to an unexpected low to moderate seismic excitation. It is theoretically established that the displacement ductility demand could vary between 2. to 4.. Modification in the reinforcement detailing specifications for non-seismic design should be made in order for improving the seismic behaviour and enhancing the ductility of these structures [6]. There are few research efforts on the potential seismic hazard of reinforced concrete shear walls in regions where earthquake is not considered as a major problem; hence, studies from which such structure s ductility demand sufficiency can be extrapolated are not readily available. In this paper, laboratory tests of large-scale reinforced concrete low-rise shear wall specimens under reversed cyclic loading, which are designed and detailed without seismic consideration as practised in low probability of seismic occurrence regions, are reported. The primary objective of this study is to verify the available ductility of reinforced concrete walls designed and detailed without explicitly considering seismic design requirements. 2. Experimental Programme In the experimental programme, six large-scaled reinforced concrete low-rise shear wall specimens were fabricated and tested under reversed cyclic loading and a constant axial compression. The reinforcement of the specimens was detailed in accordance with the detailing practice for buildings defined in BS 8 [2], in which only gravity and lateral wind loads have been considered. This detailing typifies that of a non-seismic detailing technique for reinforced concrete. The variables investigated include the effects of longitudinal steel distribution, wall aspect ratio and availability of boundary confinement. A summary of the test specimens and their properties is presented in Table. Geometry and steel detailing is shown in Figure. In the first part of the specimen label of Table, the first letter U stands for unconfined boundaries (wall ends), which can be found in the conventional detailing, and C for confined boundaries. The second letter D stands for uniformly distributed longitudinal reinforcement and C for concentrated longitudinal reinforcement at the boundaries of the specimens. The number followed represents the aspect ratio of the specimens which is defined as the ratio of height to width of the wall panels. 2.. Test Specimens All six specimens have a rectangular cross section of 2 mm mm with the wall panel height of 2 mm and 8 mm, corresponding to the aspect ratios of. and.5 respectively. The concrete cube strength of specimens is obtained from the mean value of 5-mm control cube specimens, and the adopted yield strength of the high yield steel is 52 MPa. In the specimens, all longitudinal reinforcing bars are fully anchored in a 5 mm thick base girder that is bolted to the strong laboratory floor in the test. A 3 mm 3 mm beam is cast with the wall panel at the top of the specimen. Specimens UD-. and UD-.5 are non-seismically detailed. Two layers of longitudinal steel are placed in a uniformly distributed form with spacing of 8 mm; while the spacing of transverse steel s v is 5 mm and two layers of transverse bars are fully lapped with U-shaped stirrups at boundaries (ends) of the wall panels. Specimens UC-. and UC-.5 are detailed with the similar detailing technique to that for specimens UD-. and UD-.5. Only the difference is made in the distribution of longitudinal reinforcement, where the steel bars are generally concentrated at the boundaries of the wall panels; while extra longitudinal reinforcement is added near the centre of the specimens for the crack control purpose. The similar detailing technique used in specimens UD-. and UD-.5 is adopted for specimens CD-. and CD-.5, where additional boundary hoops are added to provide better concrete confinement for the wall panels and to simulate shear walls connected with an ordinary moment resisting frame. The design of the confining hoops is referenced to the American Concrete Institute Building Code [] with a great realest on both the required steel ratio and spacing of the confining hoops. The spacing of the confining hoops is 5 mm.

Table Properties of Test Specimens Specimen Aspect Concrete Steel ratio f cu f y (MPa) (MPa) Longitudinal bar ρ v ρ l Distribution (%) (%) UD-.. 38. 52.92 Uniform.5 No UD-.5.5 43.6 52.92 Uniform.5 No UC-.. 44. 52.5 Concentrated.5 No UC-.5.5 42.8 52.5 Concentrated.5 No CD-.. 39. 52.92 Uniform.5 Yes CD-.5.5 42.2 52.92 Uniform.5 Yes Boundary confinement () 4 6 2 38 38 35 3 2 38 5 3 25 () 35 5 3 3 T--8 T-2-5 2/8 5 3 T- T-2-5 2/8 5 6 6 66 66 ANCHORAGE BOLT HOLE ANCHORAGE BOLT HOLE (a) Specimens UD-. and.5 (b) Specimens UC-. and.5 () 2 4 6 38 38 35 3 3 2 () 26 75 T--8 T-2-5 5 3 T-3-5(9 X-hoops) 2/8 6 66 ANCHORAGE BOLT HOLE (c) Specimens CD-. and.5 Figure Geometry and Steel Detailing 6 ALL UNITS ARE IN MILLIMETERS. ANCHORAGE BEND RADIUS = 3 X BAR DIAMETER (d) Cross-section

2.2. Test Setup and Procedure The test setup and loading system are shown in Figure 2. It is seen that the specimen is mounted on the strong floor of the laboratory. A vertical load of 5 kn, which is about 2% of the ultimate axial strength of the specimen, is first applied by two hydraulic jets that are connected with a pair of loading frames, which can be moved together with the specimen during the experiment. The vertical load is distributed at the top beam of the specimen. Lateral load reversal is then applied through the top beam by a servo actuator which was supported by the strong reaction wall in the laboratory. Steel plate was attached to the actuator s swivel head and connected to the steel plates on both ends of the top beam by four 4-mm-diameter high tensile bars along the longitudinal axis of the top beam. In the test, both load and displacement controls are adopted at different loading stages. The load-control method is used at the early loading stages; one cycle of horizontal loading of ±.5P i and then ±.75P i are applied, where the load P i is the cyclic applied load at the top of the specimen when the beam reaches its ultimate flexural strength M u. The value of M u is determined using BS 8 simplified stress block for concrete at ultimate limit state without partial factors of safety. Figure 4 shows a general reversed cyclic load-deflection relationship of reinforced concrete structures. It is well recognised [] that the yield displacement Δ y can be defined and calculated based on the stiffness when the lateral load is ±.75P i, which is then extrapolated linearly to ±P i, as shown in Figure 3. Thus, 2 y () 2 where Δ and Δ 2 are horizontal displacements corresponding to P i and P i, respectively. The reversed cyclic loading arrangement is then switched to the displacement control, in which the test specimens are subjected to two complete cycles of reversed loading gradually to achieve μ = ±, ±2, ±3,, where μ is the displacement ductility factor defined as y Each test is continued until the specimen experiences a significant loss of capacity, where it is assumed that failure occurs when the ratio of the restoring force at the maximum displacement to the maximum applied lateral load reaches.8. The loading history of the test is shown in Figure 4. Figure 2 Experimental Setup

Horizontal Load P i.75 P i.5 P i Δ Δ.75 Δ.75 Horizontal Displacement Δ 2.5 P i.75 P i P i Figure 3 Definition of Yield Displacement 6 4 Ductility Ratio 2-2 -4-6 3 4 5 6 7 8 9 2 3 Cycle No. (a) Load-control cycles Figure 4 Loading History (b) Displacement-control cycles 3. Test Results 3.. General Observation All specimens failed with reinforcement yielding at boundary and concrete crushing of the compressive zone. Table 2 summarises the calculated and maximum test strengths and failure behaviour of the specimens, where the shear capacity V u of specimens are calculated using the Softened Strut-and-Tie Model [5]. Flexural cracks are observed to form after two cycles with displacement ductility ratio of.. Extensive cracking is commonly observed when the applied load reaches 75% of the calculated strength of the specimens. Flexural cracks generally propagate further towards the centre line of the wall panels with uniformly distributed longitudinal reinforcement; while those in specimens UC-. and UC-.5 are only found at the boundary regions. Diagonal shear cracks generally extending from flexural cracks form in the earlier cycles; while those in specimens UC-. and UC-.5 are extended across the diagonals of the wall panels. In general, it is observed that shear distortion contributes about 2% to 5% of the total deformation of the specimens. Typical observed crack patterns after failure of test specimens are shown in Figure 5. 3.2. Hysteretic Behaviour and Energy Dissipation Capacity Hysteresis loops of the ductility factor (defined as top displacement divided by yield displacement) of the specimens versus the corresponding normalised horizontal load (defined as the applied load divided by calculated capacity of the specimen) are plotted in Figure 6. The degradation

of strength and stiffness for specimens with a lower aspect ratio (. in the tests) tends to be steadier comparing to that with a higher aspect ratio (.5 in the tests). A small degree of pinching is observed in specimens UC-., UC-.5, CD-. and CD-.5. For all six specimens, pinching of hysteretic loops corresponding to shear distortion is not observed. The area inside the hysteretic loops can be considered as an indirect measure of the energy that is dissipated by the plastic behaviour of structural elements during seismic events, which is also an important parameter for assessing seismic damage to building structures. Values of the energy dissipation are obtained by integrating the hysteretic loops before the failure cycle. Because of the differences in strength and yield displacement of each specimen, for better comparison it would be convenient to normalise the energy dissipated with the idealised elastic strain energy defined as U P (3) 2 e The yield displacement, ductility factor and energy dissipation capacity of the specimens are given in Table 3. It is seen that the energy dissipation capacity of the wall panels without boundary confinement reinforcement varies generally from about 7 to 7 times the idealised elastic strain energy. Large difference in the energy dissipation capacity is observed between specimens UC-. and UC-.5 that have only difference in the aspect ratios. The normalised energy dissipation of specimen UC-.5 is less than half of that of specimen UC-.. y i Table 2 Calculated and Test Strengths Specimen Calculated Strength Maximum Test Load P i (kn) V u (kn) (kn) UD-. 32 737 36 UD-.5 246 689 277 UC-. 43 78 455 UC-.5 28 68 34 CD-. 343 74 45 CD-.5 232 684 28 Failure Behaviour Steel reinforcement yielding and concrete crushing at boundary (a) UD-. Figure 5 Typical Crack Patterns and Failure of Specimens (b) UC-.5

Normalized load.5.5-6 -4-2 2 4 6 -.5 - -.5 (a) UD-. Normalized load.5.5-6 -4-2 2 4 6 -.5 - Normalized load -.5 (c) UC-..5.5-6 -4-2 2 4 6 -.5 - -.5 (e) CD-. Figure 6 Hysteretic Loops of Specimens Normalized load.5.5-6 -4-2 -.5 2 4 6 - -.5 (b) UD-.5 Normalized load.5.5-6 -4-2 2 4 6 -.5 - Normalized Load -.5 (d) UC-.5.5.5-6 -4-2 2 4 6 -.5 - -.5 (f) CD-.5 Table 3 Yield displacement, Ductility and Energy Dissipation Capacity Specimen Yield displacement Δ y (mm) Displacement ductility factor μ Normalised energy dissipation UD-. 3.8 3.6 7.24 UD-.5 5.2 2.78.2 UC-. 4.75 3.7 5.6 UC-.5 6.7 2.64 7.43 CD-. 4.5 2.99 7.28 CD-.5 4.78 2.97 5.74 3. Discussions 3.. Effect of Wall Panel Aspect Ratio The walls with a higher aspect ratio are shown to have lower displacement ductility as well as the energy dissipation capacity. The walls with a lower aspect ratio show much better energy dissipating performance. Because all specimens are designed for flexural failure, flexural cracks are commonly first observed. For walls with a higher aspect ratio, most of cracks are found within the lower 2/3 portion of the wall panels; while for walls with a lower aspect ratio, cracks spread evenly over the whole panels. In addition, more shear cracks are observed on the walls with a lower aspect ratio. These differences result in a comparatively more uniform strain distribution for the walls with a lower aspect ratio in the post-yielding cycles which may enhance the displacement ductility. 3.2. Effect of Reinforcement Distribution More diagonal shear cracks are observed on the walls with the end concentrated reinforcement than those with uniformly distributed bars. This may be caused by the difference between shear and flexural capacity, together with a more observable shear dominated zone []. Strain hardening effect is more observable for the walls with uniformly distributed reinforcement. The walls with concentrated

reinforcing bars tend to exhibit more pinched hysteretic response, and it is likely that shear walls with this design may not perform as well as those having uniformly distributed bars. 3.3. Effect of Confinement Reinforcement Both specimens CD-. and CD-.5 achieve a ductility factor of 3 in the tests. The wall panels with confining hoops are found to be not effective in providing higher ductility in this study. With limited amount of boundary confinement reinforcement, the enhancement of energy dissipation capacity can be observed for the walls with an aspect ratio of.5. As the hoops attain only 2% to 5% of the ultimate capacity at failure of the specimens, closer spacing of confining reinforcement, even with smaller diameter and weaker material strength, may be taken into consideration to recognise the use of confining hoops to improve the overall seismic performance of the structure. 4. Conclusion This experimental investigation on large-scale, non-seismically designed shear walls reveals that the inherent displacement ductility factor of 2.5 to 3 may generally be achieved for the shear walls with non-seismic reinforcement detailing. The arrangement of uniformly distributed longitudinal reinforcement in wall panels seems to be more beneficial for the walls with a higher aspect ratio in term of obtaining better displacement ductility and energy dissipation capacity, compared to the concentrated arrangement of longitudinal reinforcement at the boundaries (ends) of wall panels. It is concluded that ordinary shear walls designed and detailed without seismic consideration may not sufficiently satisfy the ductility demand for shear-wall building structures in low-to-moderate earthquake regions. Modification in reinforcement detailing specifications for non-seismic design should be made in order for improving the seismic performance and enhancing the ductility and energy dissipation capacity of non-seismically designed and detailed reinforced concrete shear walls in regions of low to moderate seismicity. Further experimental studies should be conducted in this area. References: [] ACI Committee 38 (999). Building Code Requirements for Reinforced Concrete (ACI 38-99) and Commentary (38 R-99). American Concrete Institute. [2] British Standards Institution (997). Structural Use of Concrete - Part : Code of Practice for Design and Construction, BS 8. London. [3] EEFIT (99). The Newcastle, Australia Earthquake. Earthquake Engineering Field Investigation Team, Institution of Structural Engineers, UK. [4] European Committee for Standardisation TC 25 (995). Eurocode 8: Earthquake Resistant Design of Structures - Part : General Rules and Rules for Buildings. Brussels. [5] Hwang, S.J., Fang W.H., Lee H.J., and Yu H.W.. Analytical model for predicting shear strength of squat walls. Journal of Structural Engineering, ASCE, Vol. 27, No., pp. 43-5. [6] Kuang, J.S., and Atanda, A.I. (25). Enhancing ductility of non-seismically designed reinforced concrete frame buildings. Journal of Structures and Buildings, ICE, Vol. 58, Issue SB4, pp. 253-265. [7] Kuang, J.S., and Atanda, A.I. (25). Predicting ductility demands on reinforced concrete moment-resisting frames for moderate seismicity. Journal of the Structural Design of Tall and Special Buildings, Vol. 4, pp. 369-378. [8] Kunnath, S.K., Hoffmann, G., Reinhorn, A.M., and Mander, J.B. (995). Gravity-load-designed reinforced concrete buildings Part : Seismic evaluation of existing construction. ACI Structural Journal, Vol. 92, No. 3. pp. 343-354. [9] Kunnath, S.K., Hoffmann, G., Reinhorn, A.M., and Mander, J.B. (995). Gravity-load-designed reinforced concrete buildings Part 2: Seismic evaluation of detailing enhancements. ACI Structural Journal, Vol. 92, No. 4. pp. 47-478. [] Muttoni A., Schwartz J., and Thurlimann B. (997). Design of concrete structures with stress fields, Basel, Boston. [] Park, R. (989). Evaluation of ductility of structures and structural assemblages from laboratory testing, Bulletin of the New Zealand National Society for Earthquake Engineering, Vol. 22, No.3, pp. 55-66. [2] Park, R. (998). Design procedures for achieving ductile behaviour of reinforced concrete buildings. Proceedings of the International Workshop on Earthquake Engineering for Regions of Moderate Seismicity, Hong Kong.