Cyclic behaviour of a full scale RC structural wall

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1 Engineering Structures 25 (2003) Cyclic behaviour of a full scale RC structural wall P. Riva, A. Meda, E. Giuriani Civil Engineering Department, University of Brescia, Via Branze, 38, I Brescia, Italy Received 8 August 2002; received in revised form 7 January 2003; accepted 7 January 2003 Abstract The results of an experimental test on a full scale RC structural wall subjected to cyclic loading are herein presented. The tested specimen is representative of a wall in a four storey building with one underground floor, designed for moderate seismic actions (PGA=0.20 g) adopting the European Seismic Code (Eurocode 8, EC8). The experimental specimen is 15.5 m long and has a transverse section of mm. The boundary conditions consist of simple supports at the foundation and ground floor levels. The wall behaviour has been studied both under service conditions, up to yielding, defined as the point to which first yield of the outermost rebars corresponds, and ultimate conditions, up to collapse Elsevier Science Ltd. All rights reserved. Keywords: RC structural wall; Seismic design; Cyclic loading; Full scale test 1. Introduction In seismic zones, building resistance to earthquakes is often ensured by adopting structural systems where seismic actions are assigned to structural walls, designed for horizontal forces and gravity loads, while columns and beams are designed only for gravity loads [1,2]. These systems, being stiffer than earthquake resisting frames, allow a better displacement control, limiting damage in internal partition walls and non structural elements. On the contrary, frame structures generally exhibit greater ductility, at the expense of large displacements and interaction problems between structural and non-structural elements. Extensive experimental results concerning the behaviour of walls of different slenderness ratio subjected to various loading conditions are available in the literature (e.g. [3 6]). These tests are generally limited to small scale specimens, typically from 1:2 to 1:3 scale, while experimental evidence on the behaviour of full-scale walls is presently scarce. The results have shown that the inelastic response of slender walls, characterized by height-over-width ratios larger or equal to 2, is con- Corresponding author. Tel.: ; fax: address: riva@ing.unibs.it (P. Riva). trolled by flexural deformations in a plastic hinge at the base of the wall. To achieve adequate ductility, an essential role is played by confining steel, placed at the edge of the section in order to control concrete crushing and longitudinal reinforcement buckling. Shear strength is provided by distributed vertical and horizontal reinforcement on both wall faces. Inclined reinforcement is sometimes needed for protection against sliding shear. The scope of the present research is to partially fill in the gap concerning full scale tests on slender shear walls, by analysing the behaviour of a full size structural wall under cyclic loads, with particular attention to its ductility, dissipated energy, damage progression, and resisting mechanisms. The tested specimen is representative of a shear wall of a four storey building, with one underground storey and a box foundation system (Fig. 1). In box foundation systems (Fig. 2), horizontal forces are resisted by the diaphragm actions of ground and basement level slabs, leading to a considerable reduction in the wall foundation dimensions [7]. In this case, the critical section of the wall is located at the ground level, where bending actions are predominant, whereas the underground part of the wall exhibits a typical shear panel behaviour. Due to expected high shear forces in the underground part of the wall, its thickness is increased. The building, hence the experimental wall, was /03/$ - see front matter 2003 Elsevier Science Ltd. All rights reserved. doi: /s (03)

2 836 P. Riva et al. / Engineering Structures 25 (2003) Nomenclature A gt F PGA q R cm R e R m M V d g Rd g c, g s steel elongation at 0.99 R m after the peak, measured over five bar diameter; net applied force at each jack, equal to the difference between the total jack force and the force necessary to obtain a bending moment equal to 0 at the critical section; peak ground acceleration structural coefficient according to EC 8 (=3 for class M structural walls); mean cube concrete strength; reinforcing steel yield strength; reinforcing steel ultimate strength; bending moment at the critical section; shear force at the critical section; displacement at the wall end; overstrength coefficient according to EC8 (=1.15 for class M structural walls); partial safety factors for concrete and steel, respectively; Subscripts n Sd Rd u y yt nominal ultimate values; acting design values; resisting design values; ultimate values at collapse; experimental yield values; values corresponding to the theoretical first yield of the critical section. designed according to Eurocode 8 (EC8) [8 10], assuming Medium ductility class (structural coefficient q=3), a peak ground acceleration PGA=0.20 g, typical for medium seismicity zones, and a Soil type B (equivalent to Soil type C in the latest EC8 version [13]). 2. Structural wall description and test set-up The structural wall was designed according to EC8 [8 10], assuming Medium ductility class (structural coefficient q=3) and a peak ground acceleration PGA=0.20 g, typical for medium seismicity zones. A higher ground acceleration could not be adopted, due to limitations in the testing loading frame available. Verification of sectional strength was carried out according to Eurocode 8 and Eurocode 2 [8 11]. Fig. 3 illustrates the wall dimensions and steel reinforcement detailing. The wall dimensions were: section mm outside of the supports, mm between the supports, length outside the supports 12.5 m, and 16 m total length. At the ground and basement levels, two ribs were inserted to simulate the floor diaphragms. As prescribed by EC8, the main flexural reinforcement was concentrated in two chords at the edges of the wall, where the reinforcement was heavily confined. Design shear force V Sd at the critical section was determined based on the overstrength prescribed by EC 8 [10] as: V Sd e V Sd q g Rd MRd 2 q M Sd 0.1 S 2 e(t C ) V S e (T 1 ) Sd 1.50 V Sd, where: q=3 is the structural coefficient, g Rd =1.15 is the overstrength factor, M Rd and M Sd are the resisting and design bending moment, respectively, S e (T C )/S e (T 1 ) is the ratio between the elastic response spectrum ordinates at the end of the constant acceleration branch and at the fundamental period, respectively, while V Sd is the shear force derived from the analysis. Shear reinforcement was limited to vertical and horizontal bars. No inclined reinforcement was inserted, as the theoretical strength evaluated according to the code was in excess of the design shear force. The governing shear resisting mechanism was found to be sliding shear at the critical section, expressed as: V Rd,s V dd V id V fd, where V dd is due to dowel action of the web reinforcement across the critical section, V id is due to inclined reinforcement (=0 in the experimental specimen), V fd is due to friction effects in the compression chord, which is the predominant resisting term.

3 P. Riva et al. / Engineering Structures 25 (2003) Table 1 shows the design bending moment and shear strengths of the critical section computed considering the material safety factors (g c =1.5, g s =1.15) and the design strength of concrete (C30/37) and steel (B500B) (M Rd, V Rd ), and the nominal strength values determined by means of experimental tests on the materials (M n, V n ). The experimental bending moment and shear force at structural yield (M y, V y ) and at collapse (M u, V u ) are also reported in the same table. It is observed that due to the test set-up, no axial force is present in the wall. As a consequence, the yield and ultimate moment as well as the shear strength of the specimen are smaller than those in a real structural wall. The average material characteristics as determined on concrete cube specimens ( mm) and steel bars were: Concrete average cube strength: R cm =40.7 MPa; Reinforcing steel yield strength: R e =560 MPa; Reinforcing steel ultimate strength: R m =640 MPa; Reinforcing steel elongation at 0.99 R m after the peak: A gt =8.4%. Fig. 1. Four-storey building adopted for the structural wall design. Fig. 2. Foundation box system for the wall. Being the reaction structure available, a prestressed concrete underground caisson which allows testing structures of span up to 40 m, the wall had to be placed horizontally, keeping the axis of maximum inertia vertical, as shown in Fig. 4. The wall was placed on two RC supports, (a) in Fig. 4, aligned with the ribs simulating the ground and basement floor diaphragms, (b) in Fig. 4, and fixed to the caisson by adopting post tensioned 0.6 strands and high strength 32 bars, (c) in Fig. 4. Strands and bars post tension was such that no decompression of the support would occur during testing. Two steel frames, (d) in Fig. 4, were placed near the loading positions in order to avoid lateral instability. The safety of the system was improved by inserting a supplementary frame between the two supports, (e) in Fig. 4, which would intervene whenever a lack in post-tension would induce a support decompression. The loads were applied at two points by means of hydraulic jacks. The position of the jacks was defined to obtain the same bending moment and shear force around the critical section as the one resulting from the analysis of the four storey building (Fig. 5). Moreover, the load position was such that the same force could be applied, greatly simplifying load control. In order to apply cyclic reverse loads, four jacks were adopted, two acting upward, (a) in Fig. 6, and two downward, (b) in Fig. 6. The jacks acting upward were placed between the wall and the loading bench, while those acting downward were placed in two windows opened in the wall and connected to the caisson with two high strength 32 bars. The position of the opening was such that the jack

4 838 P. Riva et al. / Engineering Structures 25 (2003) Fig. 3. Wall dimension and steel reinforcement. Table 1 Theoretical and experimental bending moment and shear strength values: design (M Rd, V Rd,s ); nominal (M n, V n ); structural yield (M y, V y ); collapse (M u, V u ) M Rd =4015 kn m M n =5910 kn m M y =5300 kn m M u =6200 kn m V Rd,s =715 kn V n =1165 kn V y =615 kn V u =720 kn would act along the wall neutral axis, thus limiting their horizontal displacement. The applied load was measured by means of a full bridge resistive pressure transducer placed on the pump manifold. The displacements were measured using 17 potentiometric transducers as shown in Fig. 6: two wire transducers (16, 17) measured the vertical displacement of the wall; 11 linear transducers (1 8, 14, 15) measured the displacements in the upper and lower chords close to the critical section; two linear transducers (11, 12) Fig. 4. Test set-up: wall supports (a); ribs simulating the ground and basement floor diaphragms (b); post tensioned strands and bars (c); steel frame to avoid lateral instability (d); additional frames for improving safety in the test set-up (e).

5 P. Riva et al. / Engineering Structures 25 (2003) Fig. 5. Action diagram: comparison between true scheme and the two jacks scheme. Fig. 7. Loading history up to the theoretical yield load: M/M yt versus cycle number. Fig. 6. Jack positions (upward jacks (a) and downward jacks (b)), and measurement devices. were used for monitoring the deformation of the panel between the supports; two linear transducers (9, 10) measured the displacement between the wall and the caisson at the supports in order to monitor any potential support decompression. All the signals were conditioned by adopting a data acquisition system (Mod. UPM 100 by HBM) and recorded on a PC. At first, the test was performed by applying cyclic loads with increasing amplitude (Fig. 7), until the theoretical yield load F yt (i.e. theoretical load for which the first yield (R e =560 MPa) of the outermost rebar occurs at the critical section), corresponding to an experimental maximum displacement at the wall end (d yt ) equal to 90 mm, was reached. After three cycles with maximum displacement equal to d yt, the load was further increased in both directions until the displacements corresponding to the structural yield in both directions were detected (+d y 120 mm and d y 120 mm). These displacements were determined by intersecting two lines tangent to the load displacement experimental curve in the II stage (after cracking) and in the III stage (after yielding). The following loading history was defined by imposing cycles of increasing displacement amplitude until failure was reached (Fig. 8). The collapse occurred during the second cycle at a displacement equal to three times the structural yield displacement, also equal to four times the theoretical first yield displacement (3d y 4d yt 360 mm).

6 840 P. Riva et al. / Engineering Structures 25 (2003) δ/ δ y cycles Fig. 8. Loading history after the structural yield load: d/d y versus cycle number. 3. Experimental results Fig. 9 shows the load versus end displacement curves (F d) for the whole loading history, where F is the net transverse load, equal to the difference between the total jack load and the jack load necessary to equilibrate the wall weight, i.e. the load which annuls the bending moment in the critical section. The results are discussed in the following, by analysing separately the load cycles before yielding, typical of service conditions, and after yielding, typical of design seismic loads. The results of the experimental test not reported and discussed herein may be found in [12] Behaviour under service loads The behaviour in service conditions was analysed by applying load cycles lower or equal to the first theoretical yield load, defined as the load for which the stress in the external reinforcement at the critical section is equal to R e =560 MPa. In the present case, the theoretical yield bending moment at the critical section and applied Fig. 10. Force F versus end displacement d for the cycles up to d yt. load are equal to M yt =4851 knm and F yt =268 kn, respectively, while the corresponding experimental end displacement is equal to d yt 90 mm. Fig. 10 shows the F d curve for the cycles up to d yt. From this figure, the following observations are made: due to cracking development, the wall stiffness decreases for increasing displacement amplitudes; cycles with constant amplitude show an almost constant dissipated energy, with reduced accumulated damage after each cycle (Table 2); the wall cyclic behaviour is nearly linear elastic, as Table 2 Dissipated energy for cycles up to d yt Cycle d + d F + F Dissipated energy per cycle mm kn J Fig. 9. Force F versus end displacement d curve for all the cycles

7 P. Riva et al. / Engineering Structures 25 (2003) confirmed by comparing the experimental response with the theoretical elastic behaviour of a cracked concrete wall (heavy line in Fig. 10). The main cracks in the zone close to the critical section (Fig. 11) have limited inclination, proving that the behaviour is governed by bending. The crack distance is close to the stirrup spacing in the external chords, where longitudinal reinforcement is concentrated, while in the middle part of the critical section the cracks tend to merge in a lower number of cracks, characterised by larger opening and greater inclination, thus showing the influence of shear stress. In any case, crack opening is limited ( 0.8 mm for the cycle at d yt ) and an overall good service behaviour of the wall is observed. As for the zone between the supports, a diffused crack pattern was detected, with a crack inclination of about 45, typical of panels loaded by pure shear, and a reduced crack opening (lower than 0.1 mm). It is remarked once more that in order to improve the shear strength in a zone where shear is particularly severe, both the wall thickness (400 mm) and the stirrups ( 12@200 mm) were increased in this area. Fig. 12. Force F versus end displacement d for the cycles after d y Behaviour under ultimate loads The experimental behaviour following the theoretical yield is presented in Fig. 12 (load versus end displacement F d) and in Fig. 13 (critical section bending moment versus end displacement M d). These figures clearly show the structural yield point, corresponding to a wall end displacement approximately equal to d y = 120 mm, with bending moment M y and shear force V y given in Table 1. Furthermore, it is possible to observe that due to the different position of the jacks in the uplift (Fig. 6(a)) and lowering phases (Fig. 6(b)), the load displacement response (F d) is non-symmetrical, while the moment-displacement one (M d) is symmetrical. The experimental curves show a sudden decrease Fig. 13. Bending moment at the critical section M versus end displacement d for the cycles up to 3 d y. upon load reversal, due to the sudden pressure release in the single effect jacks when the retaining valve is opened. The test results allow the following observations: Cyclic response is stable up to failure and no relevant pinching is observed in the cycles; The dissipated energy for each cycle, reported in Table 3 and Fig. 14, increases almost linearly with Table 3 Dissipated energy for cycles beyond d y Cycle d + d F + F Dissipated energy per cycle mm kn J Fig. 11. marked). Crack pattern at the theoretical yielding (principal cracks are

8 842 P. Riva et al. / Engineering Structures 25 (2003) Fig. 14. Dissipated energy per cycle. the displacement amplitude, while it remains almost constant for repeated cycles at constant amplitude (cycles 23, 24, 25), thus showing that damage progresses in a stable manner. Fig. 15 illustrates the average strain in the chords during the cycles. This figure shows that bending reinforcement progressively yields with increasing cycle amplitude. At collapse, all of the reinforcement within the instrumented zone (up to 1600 mm from the base section) may be considered as yielded. Fig. 16 shows the development of the crack pattern in the critical section for increasing cycle amplitude. Image 16(a), concerning the last cycle at yield (cycle 25), illustrates that the main cracks are evident but their opening is still small. More evident are the cracks in the cycle at 2 d y (Fig. 16(b) and (c), cycle 27). In detail, it is possible to observe a vertical crack close to the ground floor diaphragm (section close to the support) and a series of wide opened and curved cracks, with the largest starting approximately 400 mm from the support and reaching the diaphragm section at middle wall depth. At 2 d y, a distributed crack pattern with limited opening is observed in the chords. The cracks tend to merge in the middle part of the wall in the aforementioned curved cracks. Starting from the cycle at 2.5 d y (Fig. 16(d) and (e), cycle 28) also the cracks in the chords exhibit a wider opening, while the main curved crack, which eventually will lead to the wall failure, shows a considerably increased opening. In the cycles at 3 d y (Fig. 16(f) and (g), cycles 29 and 30) the damage is relevant, especially in the zone between the chords. By comparing the crack pattern corresponding to the uplift (Fig. 16(f)) and lowering (Fig. 16(g)) phases, it is possible to note that in the latter case the dead load, increasing the bending moment gradient, induces a higher crack localisation with wider crack opening. Fig. 16(h) shows the wall after the uplift phase of the second cycle at 3 d y, just before failure, which occurred in the following unloading phase. The main crack exhibits a maximum opening of approximately 50 mm towards the wall mid-depth and 10 mm at the chords. In the chords, concrete spalling, due both to compression forces and rebar bending, is observed. As shown in Fig. 16(i), the large crack opening in the middle part of the wall leads to a marked strain localisation in the longitudinal shear reinforcement, resulting in its tensile failure with necking. Concerning shear resistance, the wide crack opening observed is not compatible with any aggregate interlock effect. Furthermore, the subsequent failure of the longitudinal panel reinforcement led to a considerable reduction in the shear strength for dowel action. As a consequence a shear failure occurred during the unloading phase, when the beneficial effect of compression in the upper chord, which enables the shear strength contribution due to friction, ceased to exist (Fig. 16(l)). Fig. 16(m) shows that the main rebars are bent due to dowel action and consequent vertical wall displacement. It is important to observe that with the wall being placed horizontally on the loading bench, the beneficial effect of the axial force due to gravity loads is not present in the tested wall, thus leading to an anticipated shear collapse. Finally, little damage is present in the zone between the two supports (Fig. 16(n)), where the observed crack pattern confirms a shear panel behaviour. The wall ductility anyway ensures adequate behaviour under the design earthquake actions. In fact, the maximum obtained ductility is equal to 4d yt, larger than the assumed design ductility, approximately equal to qd yt =3d yt. Furthermore, it is observed that a very large ultimate displacement was obtained, equal to 360 mm or l/35, l=12.50 m being the wall height. Nonetheless, it is important to note that collapse did not occur in bending, but was due to shear as a consequence of a lack of longitudinal reinforcement in the wall web (between the chords). In fact, the amount of web reinforcement provided in accordance to EC8 [10] was not enough to limit the observed crack opening, and friction contribution to shear strength resulted in being much smaller than expected. Accordingly, a higher amount of reinforcement should be adopted in the web to ensure a greater contribution of dowel action and of inclined bars, whenever present. This could be obtained by either reducing the friction contribution, or increasing the overstrength factor when determining the design shear force Conclusions This paper presents the results of an experimental test conducted on a full scale RC structural wall subjected to cyclic reverse loading with amplitude increasing up to collapse. The results allow the following considerations:

9 P. Riva et al. / Engineering Structures 25 (2003) Fig. 15. Average strain in the chords along the critical zone. service behaviour was found to be mostly linear elastic, with reduced damage and small dissipated energy; the behaviour after yielding showed a progressive damage of the wall with increasing imposed displacement cycle amplitude. The damage is mainly localised at the critical section, where cracking progressively developed into large, wide open cracks, and concrete crushing was observed; no significant strength and/or stiffness degradation was observed during cycles following yield; the collapse mechanism was governed by shear, with formation of a single large crack near the base section, almost parallel to the ground floor diaphragm, leading to the tensile (necking) failure of the longitudinal wall reinforcement; a considerable ultimate displacement was obtained, equal to 360 mm or l/35, l=12.50 m being the wall height; regardless of the early shear failure, the ductility coefficient with respect to the theoretical first yield, defined as the instant when the outermost rebars reach the yielding stress, was equal to d u /d yt =4, while the ductility coefficient with respect to the observed structural yield was equal to d u /d y =3; the stability of the wall response up to collapse leads to the conclusion that a considerable ductility margin was still available with respect to bending failure; the shear reinforcement provided at the critical section proved to be insufficient to avoid an early shear failure; the latest version of EC8 [13] prescribes a minimum overstrength for shear design equal to 1.50 for duc-

10 844 P. Riva et al. / Engineering Structures 25 (2003) Fig. 16. Cracking development around critical section from d y to collapse. tility class M walls, while for class H structures the same overstrength factor adopted by the 1994 EC 8 version [10] is given. Concerning sliding shear failure, the same formulation adopted in [10] is proposed, with a slight increase of the friction contribution for normal strength concrete. Based on the results herein presented, the adopted approach might lead to nonconservative design for sliding shear effects, as shear friction resistance might be overestimated; in order to prevent sliding shear failure, particularly when axial force in the wall is negligible, the web reinforcement amount should be increased. This could be obtained by either reducing the friction contribution, or increasing the overstrength factor when determining the design shear force. Acknowledgements The research was co-financed by MURST (Italian Ministry of University, Scientific and Technological Research) within the program COFIN-99 Safety of RC structures under seismic actions with reference to design criteria of ultimate strength and damage limitation given by EC8. The contribution of UNIECO s.c.r.l., Calcestruzzi s.p.a., Ferriera Valsabbia s.p.a., Italcables s.p.a.

11 P. Riva et al. / Engineering Structures 25 (2003) towards the construction of the experimental specimen is kindly acknowledged. Finally, the authors are grateful to Mr GianPaolo Beccari, Mr Giovanni Grazioli, Mr Maurizio Lancini, Mr Denny Rivetti, Mr Marco Sandrini, Mr Gabriele Tosi, who carried out both design and experimental test of the wall under their respective Master Thesis projects. References [1] Paulay T, Priestley MJN. Seismic design of reinforced concrete and masonry buildings. New York: J.Wiley & Sons, [2] Paulay T. Earthquake-resisting shearwalls New Zealand design trends. ACI Journal 1980;77(3): [3] Bertero VV, Popov EP, Wang TY, Vallenas J. Seismic design implications of hysteretic behavior of RC structural walls. In: Proceedings of the 6th World Conference on Earthquake Engineering, New Delhi, vol p [4] Paulay T, Priestley MJN. Stability of ductile structural walls. ACI Structural Journal 1993;77(4): [5] Pilakoutas K, Elnashai AS. Cyclic behaviour of RC cantilever walls, Part I: Experimental results. ACI Structural Journal 1995;92(3): [6] Tasnimi AA. Strength and deformation of mid-rise shear walls under load reversal. Engineering Structures 2000;22: [7] Giuriani E, Gubana A, Chiari G, Contessi R. Box foundation for structural walls under seismic actions (in Italian). Technical Report, University of Brescia, Civil Engineering department, [8] Eurocode 8, Design provisions for earthquake resistance of structures Part 1 1: General rules, seismic actions and general requirements for structures, ENV , [9] Eurocode 8, Design provisions for earthquake resistance of structures Part 1 2: General rules, general rules for buildings, ENV , [10] Eurocode 8, Design provisions for earthquake resistance of structures Part 1 3: General rules, specific rules for various materials and elements, ENV , [11] EUROCODE 2, Design of concrete structures Part 1 1: General rules and rules for buildings, ENV , [12] Riva P, Meda A, Giuriani E, et al. Cyclic behaviour of a full scale RC structural wall (in Italian). Technical Report N.18/2002, Dip. di Ingegneria Civile, Università di Brescia, November [13] EUROCODE 8: Design of structures for earthquake resistance Part 1: General rules, seismic actions and rules for buildings PrEN , DRAFT No. 5, Revised Final Project Team Draft (prestage 49), Doc CEN/TC250/SC8/N317, May 2002.