EXPERIMENTAL RESPONSE OF BOUNDARY ELEMENTS OF CODE- COMPLIANT REINFORCED CONCRETE SHEAR WALLS

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1 10NCEE Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering July 21-25, 2014 Anchorage, Alaska EXPERIMENTAL RESPONSE OF BOUNDARY ELEMENTS OF CODE- COMPLIANT REINFORCED CONCRETE SHEAR WALLS C. A. Arteta 1, D.V. To 2, and J. P. Moehle 3 ABSTRACT This paper presents laboratory test results of the response of seven thin boundary element test specimens subjected to pure compression. The tested elements have transverse reinforcement detailing that matches or exceeds current code requirements for special boundary elements. Main test variables were the amount and spacing (both vertical and horizontal) of the hoop and crosstie reinforcement. The specimens were tested under monotonic loading with uniform strain until failure. Measured load-displacement relations did not exhibit an acceptable ductile behavior suggesting that current building code requirements for special boundary elements do not necessarily achieve effective confinement to be protected against brittle axial failure. 1 Auxiliary Professor, Dept. of Civil and Environmental Engineering, Universidad del Norte, Barranquilla, Colombia. carteta@uninorte.edu.co. 2 PhD. Student, Dept. of Civil and Environmental Engineering, University of California, Berkeley, USA. tovuduy@berkeley.edu. 3 Professor, Dept. of Civil and Environmental Engineering, University of California, Berkeley, USA. moehle@berkeley.edu. Arteta CA, To DV & Moehle JP. Experimental response of boundary elements of code-compliant reinforced concrete shear walls. Proceedings of the 10 th National Conference in Earthquake Engineering, Earthquake Engineering Research Institute, Anchorage, AK, 2014.

2 EXPERIMENTAL RESPONSE OF BOUNDARY ELEMENTS OF CODE-COMPLIANT REINFORCED CONCRETE SHEAR WALLS C. A. Arteta 1, D. V. To 2 and J. P. Moehle 3 ABSTRACT This paper presents laboratory test results of the response of seven thin boundary element test specimens subjected to pure compression. The tested elements have transverse reinforcement detailing that matches or exceeds current code requirements for special boundary elements. Main test variables were the amount and spacing (both vertical and horizontal) of the hoop and crosstie reinforcement. The specimens were tested under monotonic loading with uniform strain until failure. Measured load-displacement relations did not exhibit an acceptable ductile behavior suggesting that current building code requirements for special boundary elements do not necessarily achieve effective confinement to be protected against brittle axial failure. Introduction Reinforced concrete shear walls are widely used in earthquake-resistant building construction because, in addition to being economical, they are capable of providing required stiffness, strength, and ductility capacity to protect building systems against strong earthquake shaking. Where earthquake shaking will result in inelastic response including compressive yielding of the boundary elements of the wall, special boundary elements are required. Such boundary elements generally have closely spaced transverse reinforcement to confine the core concrete and provide lateral support for longitudinal reinforcement, thereby providing a capability for stable compressive yielding. Building codes around the world prescribe reinforcing steel detailing requirements whose purpose is to enable the boundary element to achieve the anticipated performance. As wall boundary element geometries have evolved over past decades, however, questions have arisen about the effectiveness of some of these building codes to achieve the desired performance. This paper presents laboratory test results of the response of thin boundary element test specimens subjected to pure compression. The tested elements have transverse reinforcement detailing that matches or exceeds current code requirements for special boundary elements. The 1 Auxiliary Professor, Dept. of Civil and Environmental Engineering, Universidad del Norte, Barranquilla, Colombia. carteta@uninorte.edu.co. 2 PhD. Student, Dept. of Civil and Environmental Engineering, University of California, Berkeley, USA. tovuduy@berkeley.edu. 3 Professor, Dept. of Civil and Environmental Engineering, University of California, Berkeley, USA. moehle@berkeley.edu. Arteta CA, To DV & Moehle JP. Experimental response of boundary elements of code-compliant reinforced concrete shear walls. Proceedings of the 10 th National Conference in Earthquake Engineering, Earthquake Engineering Research Institute, Anchorage, AK, 2014.

3 main test variables are the amount and spacing (both vertical and horizontal) of the hoop and crosstie reinforcement. The effects of these variations are reported, with an aim of better understanding the detailing requirements necessary to achieve ductile response. Outcomes of this investigation suggest that current building code requirements for special boundary elements do not necessarily achieve effective confinement to protect their sections against brittle axial failure. Field observations and motivation The 2010 Chile earthquake resulted in crushed boundary elements in several multi-story shear wall buildings. Figure 1 shows examples of observed localization of damage in the boundary zones of thin reinforced concrete walls. Though these failures could be explained by the fact that transverse reinforcement is widely spaced and inadequately anchored, a question that still needs to be answered is whether thin walls with improved boundary element detailing would behave in a more ductile manner. The research reported here is designed to answer this question. Figure 1. Localization of damage in boundary zones of thin shear walls during the 2010 Chile earthquake (photo on the left by Patricio Bonelli). Geometry and reinforcement characteristics of tested specimens Specimens were detailed following ACI [1] provisions for special boundary elements. Design nominal concrete strength was 4,000 psi at 28 days and reinforcing steel yield strength was 60,000 psi. Figure 2 depicts general geometric characteristics of two sets of walls tested: the first set comprised two short walls with cross section 8 by 24 and height of 48, and the second set comprised five thicker walls with cross section 12 by 36 with height of 72.

4 Figure 2. General geometry of tested specimens (left) and test setup for specimens W1 (center) and W5 (right). Transverse reinforcement of all specimens complies with Section c of ACI Table 1 describes the transverse reinforcement characteristics of the tested specimens. Equations 1 and 2, corresponding to equations 21-5 and 21-4 of ACI respectively, define the amount of transverse reinforcement in the two orthogonal directions of a confined column cross section. Only Equation 1 is required by ACI for special wall boundary elements... (1) (2) where s [in.] is the center-to-center spacing of transverse reinforcement; b c [in.] is the crosssectional dimension of the core measured to the outside edges of the transverse reinforcement; A sh [in. 2 ] is the total provided cross-sectional area of transverse reinforcement within spacing s, and perpendicular to dimension b c ; A g is the gross area of the concrete section; A ch is the crosssectional area of the core measured to the outside edge of the perimeter hoop; f c is the specified unconfined concrete compressive strength; and f yt is the specified yield strength of the transverse reinforcement. Since it was of interest to validate whether compliance with either of the above equations have positive repercussions in the ductile behavior of boundary elements under pure compression, the design of Walls 1 through 7 complied with Equation 1 and, Walls 2 and 4 also complied with Equation 2. Columns a through e in Table 1 summarize the characteristics mentioned above for the seven specimens.

5 Table 1. Geometry and transverse reinforcement of tested walls. Eq. 1 / (s b c ):. Eq. 2 / (s b c ):. ( ) % ( ) % ( ) W W % ( ) As designed + : As tested + : (0.55) (0.56) % ( ) 1.66 (1.52) 1.66* (1.54)*. ( ) ( ) W (0.47) 0.91 (0.71) W (0.55) 0.91 (0.83)* W (0.56) 0.91 (0.85) W (0.58) 0.91 (0.89) W () 0.91 (0.91) h x : center-to-center horizontal spacing of tied bars in the long direction of the section. * complies with ACI Eq For columns d and e, As designed values are based on nominal concrete and steel strength: f c = 4 ksi and fyt = 60 ksi respectively. Values in parenthesis, As tested, are calculated accounting for actual material properties. It was hypothesized that the ratio / in some designs is too large to achieve effective confinement, hence the impact of the confinement window size, as shown in Figure

6 3, was studied by varying the area which is bounded by the horizontal spacing between tied bars and the vertical spacing between adjacent layers of transverse reinforcement. To test this, specimens W6 and W7 have the same gross geometry and similar area of longitudinal reinforcement as W3 and W5, but the distance between tied bars is 25% shorter. To test the relative effectiveness of 90-degree hooks and 135-degree hooks to prevent longitudinal bar buckling, two pairs of the seven specimens were constructed with identical longitudinal and transverse reinforcement ratios, differing only in the detailing of the cross-ties. The two pairs are specimens {W3, W5} and {W6, W7}. As shown in Table 1, the first specimen of each pair (W3 and W6) had alternating ties with 90-degree hooks at one end and 135-degree hooks at the other, while the others (W5 and W7) had two 135-degree hooks anchored into the core at both ends. Figure 3. Geometry of the confinement window" bounded by hx and s. Test program Figure 4 shows the setup and instrumentation used for testing. Axial shortening of the walls was recorded using linear displacement potentiometric transducers. Embedded strain gauges measured strains in the longitudinal steel of the corner bars and those of the transverse steel were measured in three different layers along the height of the wall in the through-thickness direction. Concrete strain gages adhered to the surface of the concrete allowed to monitor onset of spalling at five locations along the height of the specimens. Axial load was applied monotonically at a rate of 60,000 to 85,000 lb/min for W1 and W2 and 150,000 to 200,000 lb/min for the rest of the test specimens. Tests were terminated when the applied load dropped below approximately 50% of the maximum applied force. Figure 2 shows photographs of the test setup for specimens W1 and W5 as representative of the two sizes of wall boundary elements studied. Test results Failures of the specimens occurred in the wall segments between the bottom and top heads, except that of specimen W4 for which the bottom head started cracking at early stages of loading. This specimen reached the expected maximum load but all its plastic deformation occurred at or below the interface of the wall segment and the base. Although the latter failure

7 invalidates objective comparison of the results of this specimen, some of its outcomes are presented because the trend of its behavior coincides with the rest of the walls. General observed behavior Figure 4. Instrumentation setup. Figure 5 shows typical load-displacement relations along with evolution of strains in the concrete cover and damage in the specimens. All specimens, except W4 and W7, experienced a sudden drop of load-carrying capacity after very small plastic deformation was sustained past the peak applied force. This reduction of force occurred after the specimens achieved average strains on the order of 0.3%, after which loss of cover occurred. Figure 6 shows pictures of specimens W1, W3, W5, and W7 after each test was concluded. It was observed that for all specimens, damage concentrated over a length of around two to three wall thicknesses along the height of the wall specimen. Normalized load versus average strain relations for all specimens are presented in Figure 7. It is apparent that the behavior of all specimens was very similar and differences are attributed mainly to aleatory variability and, to some extent, to variations in the transverse reinforcement detailing. In accordance with what was stated before, specimen W7, with the smallest confinement window, showed the least pronounced slope of the plastic deformation portion after attaining the maximum load. Use of 135-degree seismic hook yielded marginal benefits only.

8 Figure 5. Response evolution of Wall 7: load-axial shortening relations (top); cover concrete axial strains evolution (center); damage evolution (bottom). Figure 6. Localization of damage.

9 Figure 7. Normalized force versus average strain relations for specimens W1 to W7. Reinforcing steel buckling and loss of confinement Longitudinal steel buckling was observed in the damage zones where plastic deformation concentrated. For the large specimens, unrestrained bars buckled over approximately three hoop sets. Restrained bars also buckled but over smaller distances. The crosstie hook detail (90-degree versus 135degree bend) did not have an apparent effect. Corner bars buckled in higher modes apparently because of restraint provided by the hoops. Figure 8 presents close-up views of specimens W1 and W7 where localization of damage took place. It is hypothesized that the lack of capability for sustaining load with increasing deformation is due in part to initiation of buckling of the non-tied longitudinal bars which is also associated to the onset of concrete cover separation from the core. This behavior introduces a weakening mechanism on the exposed core section, forcing all the deformation into a small region, resulting in low overall displacement capacity. Figure 8. Buckling of longitudinal steel on specimens W1 and W7.

10 One of the requisites for a reinforced concrete element in pure compression to behave in a ductile manner is that, after spalling, the confined core axial strength is larger than that of the unconfined gross cross-section. Under this condition, strain-hardening of the core after initial spalling produces a spread of plasticity in which undamaged sections lose their cover progressively, thereby avoiding the concentration of deformation in a single location. Equations 3 and 4 define the axial load capacity of a reinforced concrete section in terms of stresses and forces acting on the unconfined gross cross-section (at the onset of spalling) and on the confined core (after spalling) with anticipation of longitudinal reinforcement buckling: = + (3) = ( ) + + ( ) (4) where is the axial strength of the section at the onset of spalling; is the post-spalling strength accounting for partial reinforcement buckling for strains just beyond the spalling strain; is in-situ confined concrete strength; is the total longitudinal bar area; and are the areas of restrained (by a tie or hoop leg) and buckled longitudinal bars in the critical section, respectively; is a factor between zero and one to account for smaller stress capacity in the buckled portion of the reinforcement. The condition to achieve ductile behavior is >. Assuming the reduction in concrete area due to embedded reinforcement is negligible (i.e., and ) and that the area of restrained longitudinal steel is equal to the area of the non-tied ones ( = = / ), manipulation of Equations 3 and 4 leads to the following expression to ensure ductile behavior: > + (5) where is the ratio of area of steel to gross area of concrete cross section. As an example for the geometries of the seven walls presented here, for a hypothetical limiting case where the stress capacity of the buckled longitudinal bars is 40% of the nonbuckled one ( = 0.40), which is a lower bound for non-restrained bars under pure compression [2], satisfying Equation 5 would have required the confined concrete strength to achieve values in the range of 1.7 to 2.0 times that of the unconfined concrete. In theory, these properties of the confined concrete are achievable with the addition of proper transverse reinforcement as demonstrated by [3], where values of / in the range of 1.35 to 2.15 are reported. For the walls reported here, confined concrete stress-strain relations were calculated for the portions of the walls where damage took place. The force attained by the core was calculated by subtracting the force on the cover and the reinforcing steel (from a model that included buckling of the unrestraint rebar) from the total measured load applied to the specimens. This force was then divided by the area of the core ( ) to calculate the achieved confined strength of the core. From the latter, confined-to-unconfined-strength ratio range was. /.. Even obtaining higher confined concrete strengths may be insufficient for

11 ductile behavior because the buckling of a longitudinal bar typically results in loss of additional concrete from the core, thereby further reducing the effective core area. Conclusions Results from the tests suggest that ductile behavior of thin boundary elements of special structural walls under pure compression is not achievable by only complying with ACI detailing provisions. Enhanced detailing (increasing the volumetric ratio of confinement reinforcement and decreasing its horizontal spacing) improved behavior but did not produce a ductile response. It is hypothesized that buckling of the longitudinal rebar both (a) reduced postspalling axial capacity of the longitudinal bars and (b) reduced post-spalling axial capacity of the confined core. These two effects led to post-spalling strength that was substantially less than the spalling load, such that plasticity did not spread along the test specimen height. Acknowledgments The research presented in this paper has been supported in part by the National Institute of Standards and Technology through the ATC 94 project and by the National Science Foundation under award number Opinions, findings, conclusions and recommendations in this paper are those of the writers and do not necessarily represent those of the sponsor. The authors would like to thank Departamento Administrativo de Ciencia, Tecnología e Inovación, Colciencas and Universidad del Norte, for providing necessary financial assistance to Carlos Arteta to pursue his PhD in the U.S. The authors would also like to thank ERICO (LENTON) and CMC Rebar Fabricators for providing part of the reinforcing steel used in the laboratory specimens. References 1. ACI Committee 318. Building Code Requirements for Structural Concrete and Commentary (ACI ). American Concrete Institute: Farmington Hills, MI, 2011; 503 pp. 2. Monti G, Nuti C. Nonlinear cyclic behavior of reinforcing bars including buckling. Journal of Structural Engineering 1992, 118:12, , ASCE. 3. Mander JB, Priestley MJN, Park R. Observed stress-strain behavior of confined concrete. Journal of Structural Engineering 1988, 114:8, , ASCE.