damage and implications for seismic design of rc structural wall building

Damage and Implications for SeismicDesign of RC Structural Wall Buildings

John W. Wallace,a) M.EERI, Leonardo M. Massone,b) Patricio Bonelli,c)

Jeff Dragovich,d) M.EERI, Ren Lagos,e) Carl Lders,f)

and Jack Moehle,g) M.EERI

In 1996, Chile adopted NCh433.Of96, which includes seismic designapproaches similar to those used in ASCE 7-10 (2010) and a concrete codebased on ACI 318-95 (1995). Since reinforced concrete buildings are thepredominant form of construction in Chile for buildings over four stories, the27 February 2010 earthquake provides an excellent opportunity to assess the per-formance of reinforced concrete buildings designed using modern codes similarto those used in the United States. A description of observed damage is providedand correlated with a number of factors, including relatively high levels of wallaxial load, the lack of well-detailed wall boundaries, and the common usage offlanged walls. Based on a detailed assessment of these issues, potential updates toU.S. codes and recommendations are suggested related to design and detailing ofspecial reinforced concrete shear walls. [DOI: 10.1193/1.4000047]

INTRODUCTION

The area impacted by the MW 8.8 2010 earthquake is the most densely populated regionof Chile and includes the cities of Via del Mar, Santiago and Concepcin. Between 1985and 2009, a total of 1,939 construction permits were issued for reinforced concrete buildingswith nine or more floors (CChC 2010). Damage observed following the earthquake wasgenerally concentrated in newer and taller buildings, with one complete collapse (AltoRo building), several partial collapses, and about 40 severely damaged buildings thatwere either repaired, or in rare cases, demolished. Severely damaged reinforced concretebuildings correspond to about 2% (40/1,939) of the newer building stock with nine ormore floors in south-central Chile.

As summarized by Massone et al. (2012), typical buildings in Chile include a large num-ber of structural walls compared with U.S. construction, such that the ratio of total wall area

Earthquake Spectra, Volume 28, No. S1, pages S281S299, June 2012; 2012, Earthquake Engineering Research Institute

a) Professor, University of California, Los Angeles, Department of Civil & Environmental Engineering, 5731Boelter Hall, Los Angeles, California, 90095-1593

b) Assistant professor, University of Chile, Department of Civil Engineering, Blanco Encalada 2002, Santiago,Chile

c) Universidad Federico Santa Mara, Departamento de Obras Civiles, Av. Espaa 1680, Valparaso, Chiled) Research Structural Engineer, National Institute of Standards and Technology, 100 Bureau Drive, Gaithersburg,

Maryland, 20899e) Ren Lagos & Asociados, Magdalena 140, Santiago, Chilef) Professor Emeritus, Catholic University of Chile, Department of Civil & Geotechnical Engineering, Av. Vicua

Mackenna 4860, Santiago, Chileg) Professor, University of California, Berkeley, Department of Civil & Environmental Engineering, 775 Davis

Hall, Berkeley, California, 94720-1710

S281

to floor plan area is on average roughly 3% in each principal direction of a building. Walls ineach principal direction are often connected to form T-shaped or L-shaped cross sections.Buildings constructed since 1996 tend to use a common layout, with longitudinal walls oneach side of a central corridor that form a spine and perpendicular (transverse) walls, or ribs,spaced at roughly 5m along the central corridor. Wall thickness of 15 cm to 20 cm is com-mon. Seismic design approaches used in Chile are similar to those used in the United States,e.g., modal response spectrum analysis; however, code design spectra in the 1996 Chile codewere strongly influenced by the M7.8 March 3, 1985 earthquake that occurred along thesubduction zone at the interface of the South American and Nazca plates.

Design requirements for RC shear wall buildings, the predominant form of constructionin Chile for buildings over four stories, were updated via reference from NCh433.Of96 toACI 318-95, with only minor exceptions. An important exception, based on the good per-formance of roughly 400 RC wall buildings in Via del Mar in the 1985 earthquake, is thatACI 318-95 requirements for special transverse reinforcement at wall boundaries to confinethe concrete and restrain rebar bucking were eliminated.

Given the large number of modern RC buildings that exist in the impacted region, the27 February 2010 MW 8.8 Chile earthquake offers a excellent opportunity to assess the per-formance of reinforced concrete buildings designed using modern codes similar to those usedin the United States. In the following, we discuss damage observations and identify issuesthat may not be adequately addressed or may need to be updated in U.S. codes and recom-mendations as well as in other countries that experience strong ground shaking.

OBSERVED DAMAGE

Widespread and significant damage to RC structural walls (Figure 1), particularly at ornear the ground level, was observed in buildings in Santiago, Via del Mar, and Concepcin,the major population centers in Chile, following the 2010 earthquake. In general, crushingand spalling of concrete and buckling of vertical reinforcement were observed, often over theentire wall length. Typically, the damage was concentrated over a short height equal to one tothree times the wall thickness, apparently because buckling of vertical bars led to damageconcentration. Closer inspection of wall boundary regions (Figure 2) revealed relatively large

Figure 1. Typical wall damage in the 2010 Chile earthquake.

S282 WALLACE ET AL.

spacing of hoops (20 cm) and horizontal web reinforcement (20 cm), as well as 90-degreehooks (Massone et al. 2012). Because walls are thin, typically 15 cm to 20 cm thick, anyspalling of cover concrete (about 2 cm on each side) results in a 20% or 27% reduction in thewall thickness. Once cover concrete spalls, the 90-degree hooks used on transverse reinforce-ment at wall boundaries become ineffective (they open). The large spacing of transversereinforcement and the opening of 90-degree hooks after concrete spalling likely contributedto buckling of vertical reinforcement.

Large cyclic strain demands on reinforcement, combined with relatively large axial stress,may have produced abrupt buckling in some walls without the need for significant or anyconcrete cover spalling. Buckling of longitudinal reinforcement in walls also was combinedin some cases with fracture of reinforcement, again, likely due to large cyclic strain demandsand the long duration motions. Lateral instability (buckling), primarily at web boundaries ofT- or L-shaped wall cross sections, was observed (Figure 3); prior to the Chile and

Figure 2. Boundary wall reinforcement: (a) Fracture, (b) buckling, and (c) detailing.

Figure 3. (a) Wall lateral instability and (b) wall discontinuities/instability.

DAMAGE AND IMPLICATIONS FOR SEISMIC DESIGNOF RC STRUCTURALWALL BUILDINGS S283

New Zealand earthquakes, this lateral instability failure had been primarily observed inlaboratory tests (e.g., Thomsen and Wallace 2004). Detailed surveys conducted as part ofATC-94 (2011) indicate that lateral wall instability was not likely driven by prior yieldingin tension (as had originally been suspected based on past research, e.g., Corley et al. 1981;Paulay and Priestley, 1993; Chai and Elayer, 1999) but instead was the result of lateralinstability of previously crushed, thin wall boundary zones. This failure mode has notbeen adequately studied.

Damage to thin walls was commonly observed in first floor or the first basement level,either at the base of the wall or near the top of the wall (Figure 1). The location of damagemight have been influenced by concrete quality at the top of walls (e.g., over- consolidationproducing a weak layer at the top of the wall); however, in many cases the location of damagewas likely due to stress/strain concentrations from abrupt changes in the wall length toaccommodate parking spaces, either at ground or basement levels (Figure 3b). Damageto vertical and horizontal wall segments, created by window and door openings on buildingfacades, was frequently observed (Figure 4a). Damage to these wall segments, typically withaspect ratio close to one, generally consisted of wide diagonal cracks (wider than 1cm inmany cases) and concrete spalling. Similar damage was observed in coupling beams(Figure 4b), as well as in wall segments across corridors, either near the roof level ornear basement levels (Figure 5, also see Figure 9). In some cases, coupling beams andwalls were not placed in the same plane (Figure 4b), creating eccentric connections thatled to damage. Significant concrete spalling and large diagonal cracks (width of several cen-timeters) also were observed at discontinuities, such as shown in Figure 5, and within floordiaphragms, where struts formed to redistribute forces from damaged walls to adjacent walls.

In some older buildings (pre-1990), damage was observed in lightly reinforced (bothlongitudinal and transverse reinforcement, by U.S. standards) lintel beams along interior

Figure 4. (a) Pier and spandrel failure. (b) Coupling beam spandrel failure.

S284 WALLACE ET AL.

corridor walls (Figure 6a), with beam damage sometimes extending over the entire buildingheight (in all stories). In other cases, poor anchorage of beam reinforcement within wallswas noted (Figure 6b). In newer construction, lintel beams over doorways were replacedwith nonstructural material; therefore, only slab coupling exists along and across corridorwalls. Damage to corridor slabs, which usually span 2 m to 3 m and are 15 cm to 20 cmthick, was observed in numerous buildings (Figure 7a), as the slabs are subjected to largerotation demands due to wall deformations. However, in some cases, severe slab damagewas observed to be due to, or exacerbated by, large vertical deformations imposed on theslab due to wall shortening at lower levels due to concrete crushing/spalling and rebarbuckling. In some taller buildings, large (cap) beams were used to couple walls at ornear the roof level. Large shear demands on cap beams, or similar cap beams nearthe building base where walls were extended across the corridor (Figure 5), tended to pro-duce significant diagonal cracking and concrete spalling. It is noted that use of diagonalreinforcement in wall piers, coupling beams, and cap beams, is not common in Chile.

In typical Chilean buildings, shear walls are used to divide apartment units; within units,partition walls are common. Severe damage was observed to these interior partition walls insome buildings (Figure 7b). Evidence of foundation rotation (soil heaving) was observed at

Figure 5. Damage at various discontinuities.

Figure 6. (a) Lintel beam failure and (b) wall-beam connection damage.


one building in Concepcin and one building in San Pedro with mat foundations near grade;however, foundation performance was generally good for taller buildings.

Although severe damage was observed in taller buildings in Concepcin, only one mod-ern RC building collapsed, the 15-story Alto Ro building constructed in 2007. The buildingincluded features similar to those in other damaged buildings, such as T-shaped wall sec-tions and vertical irregularities (setbacks) in transverse walls and discontinuous perimeterwalls just below the top of the first floor along axis I to accommodate parking (Figure 8), aswell as discontinuities below the first floor where transverse walls were extended across thecentral corridor at axes 5, 8, 11, and 13 (e.g, Figures 8b, 9a, 9b). Damage at transverse wallelevations at Axes 8 and 13 are provided in Figure 9a, 9b (IDIEM 2010), which indicateconcrete spalling and reinforcement buckling at discontinuities. A schematic of a possiblecollapse scenario is shown in Figure 9c. Based on observed wall damage in a large numberof buildings, as well as the wall configuration for the Alto Ro building, it appears quitelikely that flexural compression failure (concrete crushing, rebar buckling) occurred atthe location where the transverse wall lengths were reduced on the side of the buildingwith parking (at the top of the first story, indicated as 1 on Figure 9c). Damage mayhave initiated at axes with T-walls (8, 13, and 20), or alternatively at the discontinuityin the L-shaped to rectangular transverse wall (11, 17, 24) at the top of the first story.In either case, once crushing/buckling initiated in one of these walls, axial load wouldbe redistributed, making the other wall more susceptible to failure. High shear stressesin the first story corridor wall below the open corridor above are likely to produce damageat this location (Fig, 9a,b), indicted as 2 on Figure 9c. Once these transverse walls weredamaged at the perimeter and at the central corridor, it is plausible that splice failureoccurred along the remaining portion of the wall, indicated as 3 on Figure 9c. Splicefailures of web vertical bars and boundary vertical bars at the intersection of Axis 8

Figure 7. (a) Corridor slab damage and (b) nonstructural elements failure.

S286 WALLACE ET AL.

and Axis A were observed (Figure 10a), whereas smaller (10 mm) diameter vertical bars inthe short perimeter wall segments at Axis 8 along Axis A fractured (Figure 10a). It isunclear whether splice failure and fracture of vertical bars led to overturning, or were aresult of overturning (Figure 10b) due to damage observed between axes D and I (Figure 9).It also is noted that response spectra for a ground motions recorded about 1 km away from

Figure 8. (a) Alto Ro typical floor plan (above first level) and (b) Alto Ro first floor plan.

Figure 9. Alto Ro partial elevations with observed damage: (a) Axis 8, (b) Axis 13, (c) collapsescenario.


the building (Massone et al. 2012) indicate generally larger spectral demands near the AltoRo building, likely due to the impact of soil. More detailed studies are needed to assess thisand other possible collapse scenarios.

DAMAGE ASSESSMENT

Damage observed to RC buildings in Chile raises a number of issues, e.g., what role diddrift demand, axial stress, wall configuration, lap splices, and the lack of well detailed regionsat the wall boundaries play in the degree of damage observed? An initial assessment of theseissues as well as potential code approaches that could be used to address these issues arepresented.

DRIFT DEMANDS

Linear and nonlinear displacement spectra for various ground motions recorded in the1985 and 2010 earthquakes are presented in Figure 11a, along with simple bilinear spectra formotions recorded at various locations (simplified SIII) and in downtown Concepcin about1 km from the Alto Rio building on soft soils (simplified SIV). The simplified (bilinear)spectra are used to estimate building drift demands, which are plotted in Figure 11b for build-ings greater than 5 stories using fundamental building periods of T1 N20

2

p, where

N20 represents the fundamental period for low-amplitude vibrations (Wood et al. 1987,Midorikawa 1990) and

2

paccounts for concrete cracking for larger amplitude vibrations.

Roof displacement is estimated from the spectra by multiplying the spectral displacement(of the single-degree-of-freedom oscillator) by 1.25 for five-story buildings, 1.5 for buildingsgreater than ten stories, and by using linear interpolation between five and ten stories (similarto the approach used in ASCE-41). A story height of 2.75 m is assumed, and spectral dis-placements for both SIII and SIV simplified spectra are multiplied by 1.25 to estimate spectrafor 2% damping, which provide a better estimate of inelastic drift (Shibata and Sozen 1976).

Figure 10. (a) Wall damage at ground line, Axes A and 8; (b) overall view of collapsed building.

S288 WALLACE ET AL.

The resulting roof drift ratios range from about 0.8% drift for five-story buildings up to a peakvalue of about 1.0% for buildings between 10 and 30 stories for the simplified SIV spectrum.

The drift at failure for compression-controlled walls is estimated to identify buildingheights with vulnerable walls for both SIII and SIV simplified spectra as: uhw 1140maxhw Nhs, where the constant 11/40 is based on a linear increasing distributionof lateral forces over the height of the wall (Wallace and Moehle 1992), max is estimatedequal to 0.004lw for compression-controlled walls, lw is the wall length, hs is the storyheight, and N is the number of stories. In general, compression-controlled walls in 10- to15-story buildings (SIII) and 10- to 20-story (SIV) buildings would be vulnerable to com-pression failure. Walls in buildings below ten stories typically have axial load P 0.10Ag f 0cand thus are unlikely to be compression-controlled.

THE ROLE OF AXIAL STRESS

Neither the Chilean (NCh433.Of96; INN 1996) nor the ACI code (ACI 318-08 2008) puta limit on the level of axial stress allowed for gravity load or combined gravity and lateralloads, although a limit of Pu < 0.35P0 was incorporated into UBC-94. As noted by Massoneet al. (2012), median axial wall stress has increased from about 0.1Ag f 0c for pre-1965 con-struction and 0.2Ag f 0c for post-1980 construction. For taller buildings (15 to 25 stories), thin-ner walls, and walls with larger tributary areas, axial load ratios of 0.3Ag f 0c to 0.4Ag f 0c arepossible (Massone et al. 2012).

To assess the impact of axial stress on wall deformation capacity, moment-curvaturerelations (Figure 12a) were calculated for a typical wall web in a 12-story building inSantiago. The wall cross section is 7m 0.15m with 8-25 mm diameter bars at eachwall boundary and 8 mm vertical web bars at 20 cm spacing. Two levels of axial stressare considered, 0.2 and 0.3Agf

0c; ratios of clw, or neutral axis depth to wall length, also

are plotted (Figure 12b). The relations plotted reveal negative slope (likely to producedamage concentration) beyond the yield curvature with eventual concrete crushing and

Figure 11. (a) Displacement spectra and (b) estimated drift demands.


relatively low curvature capacity (it is noted that the buckling of vertical bars is not consid-ered). The plot of clw indicates that this ratio is never less than approximately 0.45 forP 0.3Agf 0c. It is noted that ACI 318-99, and subsequent editions, require Special BoundaryElements for clw > 0.24 (Wallace and Orakcal 2002).

ACI 318-08 (2008) requires that transverse reinforcement (Ash) at wall boundaries satisfyEquation 21-5 Ash 0.09sbc f 0cf yt, where s is vertical spacing of transverse reinforcement,bc is the dimension of the confined core, f 0c is the concrete compressive strength, and f yt isthe yield stress of the transverse reinforcement. ACI 318-08 Equation 21-4, Ash 0.3sbcf 0cf ytAgAch 1, where AgAch is the ratio of the gross area to the confined core atthe wall boundary, is based on equating the pre- and post-spalling axial load for columnsusing a simple model that accounts for the stress increase due to confinement of the columncore. This equation does not need to be satisfied at wall boundaries, although it was requiredprior to ACI 318-99. For thin walls, the ratio of concrete cover to wall thickness is large, oftenin the range of 0.2 to 0.3. In such cases, spalling of concrete cover results in substantial loss ofaxial load capacity at a wall boundary, possibly overstressing wall boundary vertical rein-forcement (and some web vertical reinforcement), resulting in abrupt strength loss dueto buckling of reinforcement. These observations suggest studies are needed to considerwhether Equation 21-4 should be reinstated, or whether relatively thin walls require evenmore stringent detailing to adequately confine the core concrete and to restrain bucklingof reinforcement (i.e., the transverse reinforcement required by ACI 318-08 21-4 mightnot be adequate). Alternative means to address this issue might be to limit the ratio ofcover to wall thickness or to specify a minimum wall thickness.

Wall lateral stability failures were observed for walls with apparent high axial stress(Figure 3a, 3b) suggesting that it might be prudent to incorporate a minimum wall thicknessas a function of the unsupported wall height, e.g., tw hs:, where hs = unsupported wall(story) height. The value used for might depend on the level of axial stress, neutral axisdepth, or expected maximum extreme fiber compressive strain; commonly suggested valuesfor range from 1/10 (Moehle et al. 2011) to 1/16 (UBC 1997).

Figure 12. (a) Moment - curvature relations and (b) neutral axis depths.

S290 WALLACE ET AL.

THE ROLE OF WALL CONFIGURATION

Common floor plans used for buildings in Chile, with long corridor walls in one direc-tion and perpendicular walls in the transverse direction (sometimes referred to as a backboneand rib pattern, or a fishbone pattern), results in buildings where lateral force resistanceis provided by walls with T- and L-shaped cross sections. Prior research on walls withT-shaped cross-sections has revealed that these walls behave substantially different thanrectangular walls with symmetrically placed longitudinal (vertical) reinforcement (Wallace1994, 1996; Thomsen and Wallace 2004). For a wall with a T-shaped cross section, the webboundary is subjected to both large tensile and compressive strains (e.g., 0.025 in tensionand 0.01 in compression for TW2 at 1.5% lateral drift; Orakcal and Wallace 2006); there-fore, this region typically must be well confined to avoid concrete crushing and reinforce-ment buckling.

Load versus displacement results are presented for similarly detailed web boundaries forwalls TW1 and RW2 (Figure 13) tested by Thomsen and Wallace (2004). Significant loss inlateral-load capacity was observed for wall TW1 at a little greater than 1% lateral drift whenall eight vertical boundary bars and some vertical web bars buckled. In contrast, wall TW2(Figure 13, photo) reached lateral drift ratios in excess of 2.5% before lateral strength degra-dation initiated. For wall TW2, strength loss resulted due to lateral instability (buckling) ofthe well-confined web boundary, after concrete cover spalled. Lateral instability failures andfracture of wall web boundary vertical reinforcement (Figure 14) suggest that the lack oftransverse reinforcement at web boundaries of T- and L-shaped walls could have played a

Figure 13. Load-displacement relations for TW1 and RW2.


significant role in the level of damage observed to walls in Chile and existing U.S. coderequirements should be carefully reviewed.

DISPLACEMENT-BASED DESIGN PROVISIONS AND MODEL ASSUMPTIONS

Sections 21.6.2 of ACI 318-99 and 21.9.6.2 of ACI 318-08 include provisions that arederived from a displacement-based approach (Moehle 1992, Wallace and Orakcal 2002).In the model used to develop ACI 318-99 provisions, the design displacement u is relatedto local plastic hinge rotation (p) and extreme fiber compressive strain (cu) as:

EQ-TARGET;temp:intralink-;e1;50;296p uhw

u

cuc

lp

lw2

(1a)

EQ-TARGET;temp:intralink-;e2;50;246 cu 2uhw

clw

(1b)

If the compressive strain exceeds a limiting value, taken as 0.003 in ACI 318-08, thenspecial boundary elements are required. In ACI 318-08, this approach is modified to define alimiting neutral axis depth (instead of a limiting concrete compressive strain) as:

EQ-TARGET;temp:intralink-;e3;50;166cACI;limit 0.003lw2uhw

lw667uhw

lw

600uhw(2)

In this approach, it is obvious that the result is sensitive to the values used for designdisplacement u and plastic hinge length lp, where it is assumed that yielding spreads out over

Figure 14. Rebar fracture at web boundaries of T-shaped walls in Concepcin: (a) Centro Mayorbuilding and (b) Alto Ro Building @ 13-A (Figure 8).

S292 WALLACE ET AL.

a height (plastic hinge length) of lp lw2. Despite fairly low drift demands (Figure 11),significant wall damage was observed (Figure 1) and it was evident that inelastic deforma-tions did not spread out in poorly detailed, highly compressed walls commonly used in Chile;therefore, it is important to reassess the ACI (and the ASCE 7) provisions in light of theseobservations.

The approach presented by Wallace and Orackal (2002) is modified here to assess thepotential impact of variation of the plastic hinge length on the need for Special BoundaryElements (SBEs), that is, the regions at wall edges where closely-spaced transverse reinfor-cement is needed to provide nonlinear deformation capacity by ensuring a stable compressionzone (and thus, adequate spread of plasticity over the wall height). Given that the ACI 318-08relation assumes a specified spread of plasticity lp lw2, the impact of concentratingdamage over shorter height was investigated by modifying the relationship presented byWallace and Orakcal (2002) to use plastic hinge length equal to a multiple of the wall thick-ness lp tw:

EQ-TARGET;temp:intralink-;e4;71;472

uhw

cutwlw

lwc

1

2

twhw

ylw

11

40

hwlw

twlw

2

2

twhw

twlw

(3)

Where tw is the wall thickness, c is the neutral axis depth, hw is the wall height, lw is thewall length, and y is the yield curvature of the section. In this study, the yield curvatureis estimated as y sylw c; alternatively, yield curvature can be estimated as1.5 to 2.0sylw, where sy is the tensile reinforcement yield strain. The constant 11/40 isbased on a linear increasing distribution of lateral forces over the height of the wall (Wallaceand Moehle, 1992). For this preliminary study, wall aspect ratio hwlw is set to 10 (for a 20story building; hwlw 50m5m) and the ratio of lwtw is set to 25 for Chilean buildings(5m0.2m). Concrete compressive strain is set to 0.003 (the value that defines when SBEsare required by ACI 318-08). Results are presented in Figure 15a for a plastic hinge length oflp tw, with set equal to 2, 6, and 12. Results for the ACI 318-08 model also are shown; itis noted that the ACI model results are different than those produced with Equation 3 becausethe ACI model neglects elastic deformations. For lp 12tw, if the drift ratio is about 0.015,

Figure 15. Impact of plastic hinge length on SBE variable (a) clw, (b) hwlw.


the neutral axis must exceed 0.15lw before SBE are required. However, for the same neutralaxis depth, 0.15lw, if inelastic deformations are concentrated over lp 2tw, a drift ratio ofonly about 0.008 can be tolerated before SBEs are required. In the limit, if approaches zero(a compression-controlled wall, where the extreme fiber concrete compressive stress reaches0.003 prior to yield of tensile reinforcement at the opposite wall boundary, i.e., no spread ofplasticity), then only the elastic drift is reached, assuming rebar buckling does not occur priorto reaching the yield displacement. For equal to zero and yield curvature of 2.0sylw, theresulting elastic drift is: uhw 0.01hwlw 0.0005N, for hw N2.5m and lw 5m.Elastic drift ratios are plotted on Figure 11b and suggest that compression controlled walls inbuildings between roughly 10 and 20 stories are more susceptible to damage. This resultis consistent with observed damage. Walls in buildings less than ten stories are likely tohave lower levels of axial stress < 0.10Ag f 0c and are less likely to be compression-controlled, which is consistent with the performance of buildings in Via del Mar in the1985 earthquake.

Variation of the wall aspect ratio hwlw is shown in Figure 15b for walls whereplasticity spreads out over the ACI assumed plastic hinge length of lp lw2. Resultsfor lp 2tw (not shown) are similar to that for well-detailed walls with large c/lw values(large axial load). Figure 15b includes aspect ratios ranging from 5 to 20, or 10 to 40 stories,respectively. Results indicate that buildings over 30 stories are less prone to damage even forrelatively large clw values, for example, due to large axial load and/or large flexural com-pression force, since drift levels greater than about 2% are required in most cases to requireSBEs and anticipated drift demands are generally less than 1.0% (Figure 11b). Figure 15balso shows that well-detailed walls with small ratios of clw values can easily achieve1% roof drift ratios. Poorly-detailed walls or walls with large axial load (large clw values)for relatively short buildings (e.g., ten stories or less) require lateral drift ratios of about0.5% to 1% before SBEs are required per ACI 318; however, displacement demands forthese shorter, stiffer buildings (fundamental period of approximately 0.5 sec) are onlyabout one-half of that needed to require SBEs. These observations again help explainthe higher concentration of damage in buildings with poorly detailed walls in the rangeof 15 to 20 stories.

As noted previously, the design displacement for the displacement-based designapproach for shear walls in ACI 318 is obtained using ASCE 7 provisions as:

EQ-TARGET;temp:intralink-;e5;50;247uACI xASCE CdeI (4)

where e is the elastic displacement for cracked section properties reduced by the responsemodification coefficient R (equal to 5 and 6 for bearing wall systems and building framesystems, respectively), Cd is the deflection amplification factor (equal to 5 for both systems),and I is the Importance Factor. In current U.S. codes, the intent is to provide 90% confidenceof non-collapse for MCE shaking. In contrast, the current ACI confinement trigger is basedon 50% confidence of not exceeding the concrete crushing limit in the Design Basis Earth-quake, which is much lower shaking intensity than the MCE.

It is necessary to adjust ACI 318-08 Equation 21-8 to be more consistent with thebuilding code performance intent. Three factors need to be considered (neglecting the

S294 WALLACE ET AL.

role of Cd): (1) MCE exceeds DBE; (2) there is dispersion about the median response; (3)damping is likely to be lower than the 5% value assumed in the ACI provisions (e.g., on theorder of 2% to 3%; ATC-72 2010). To address these issues, the displacement value used inthe denominator of Equation 21-8 in ACI 318-11 should be increased by a factor of approxi-mately 1.5 to adjust to MCE level shaking and to consider dispersion, and by approximately1.2 to 1.3 to account for potential lower damping ratios; therefore, for Equation 21-8, either amultiplier of two should be applied to the ASCE 7-05 displacement, or the coefficient of 600in the denominator should be approximately doubled to 1,200.

BUCKLING AND FRACTURE OF REINFORCEMENT

Poorly detailed and/or compression-controlled walls, that is, walls that lack closelyspaced transverse reinforcement to sustain a stable compression zone and ensure spreadof plasticity by confining core concrete and suppressing rebar buckling, exhibited poor beha-vior (Figure 2). The longitudinal boundary bars, typically of 18 mm to 25 mm diameter inChilean buildings, are typically enclosed within 8 mm diameter horizontal web reinforcementspaced at s 20 cm on center with 90 hooks at the wall boundary (see Figure 2c); therefore,sdb ratios are typically in the range of 8 to 11. Such large sdb ratios are likely to resultin buckling of vertical reinforcement following even modest tensile strain excursions, i.e.,around 0.01 (Rodriguez et al. 1999). Damage appeared to initiate at the wall boundary (andextend over a height of approximately two to three wall thicknesses), and then propagatetowards the interior of the wall, or towards the wall flange in the case of T-shaped crosssection walls.

Reinforcement at wall boundaries is subjected to large variations in tension and compres-sive strains when subjected to earthquake (reversed cyclic) loading. Compressive strains arelarge, especially for large axial stress and for web boundaries of walls with T-shaped crosssections (Figure 16), potentially leading to compression failure. Alternatively, if reinforce-ment is subjected to large tensile strain demands (exceeding yield), cracks open, and uponreverse loading, all compression must be resisted by reinforcement, potentially leadingto buckling failure. As noted previously, due to larger variation in the tension and compres-sive strain demands (Figure 16), web boundary longitudinal reinforcement for walls withT-shaped cross sections are much more susceptible to buckling than are boundary longitu-dinal reinforcement in walls with rectangular cross section (for the same axial stress ratio;e.g., see Figure 13, Thomsen and Wallace 2004).

Given the typical wall configurations used in Chile (corridor and transverse walls), aswell as the lack of closely spaced transverse reinforcement at wall boundaries, it is not sur-prising that damage at web boundaries of walls with T-shaped cross sections was common(Figure 16). However, the quantity of transverse reinforcement required by ACI 318-08 doesnot ensure that the post-spalling axial strength of the core concrete is sufficient to sustain theaxial load demand. In addition, relatively few tests have been conducted on T-shaped walls,or walls with flanges, and in the tests where ductile behavior was observed, e.g., TW2(Figure 13), the sdb ratio was substantially less than the limiting value (3.33db versus6db). The findings suggest that more stringent detailing may be needed to ensure ductilebehavior at T-shaped walls.


CONCLUSIONS

The MW 8.8 earthquake that struck Chile on 27 February 2010 provides a excellentopportunity to study the performance of reinforced concrete buildings designed using amodern seismic code provisions and concrete design based on ACI 318-95. Based on recon-naissance efforts and subsequent preliminary studies, a number of issues have been identifiedto help us understand the observed damage, particularly the damage to shear walls, and toidentify areas where changes to ACI 318-08 may be warranted. Based on this preliminarywork, the following observations are noted:

Pre-1985 buildings in Chile typically performed well because of the large ratios ofAWAf (stiff buildings), relatively low wall axial stress (typically < 0.10Agf 0c), andrelatively low displacement demands.

Post-1985 buildings, and particularly buildings constructed since 2000, shear wallstend to have larger axial stress ratios and larger roof drift ratios, particularly on softsoils (SIV). For compression-controlled walls, where the concrete compressivestrain reaches 0.003 prior to yield of tension reinforcement, simplified estimatesof drift demands and drift capacities indicated that compression-controlled wallsin 10-15 and 10-20 story buildings were susceptible to failures for SIII and SIVsoils, respectively. These findings are generally consistent with observed damage.

Post-2000 buildings in Chile commonly include central corridor walls and multipletransverse walls, creating T-shaped walls, which are susceptible to web boundarydamage due to the large reinforcement tensile strains accompanied by large concrete

Figure 16. Web boundary strain demands for T-shaped wall.

S296 WALLACE ET AL.

compressive strains that develop at web boundary. For poorly-detailed web bound-aries, sudden loss of lateral strength due to buckling of vertical boundary and webbars has been observed in tests and likely was a significant factor in the degree ofdamage observed following the 2010 earthquake (along with axial stress).

The displacement-based design approach used in ACI 318-08 assumes a plastichinge length of lp 0.5lw to determine the need for special boundary elements;however, damage was concentrate over a much shorter distance, typically over aheight equal to two to three times the wall thickness. Sufficient transverse reinforce-ment must be provided to ensure a stable compression zone at wall boundaries toachieve the spread of plasticity assumed.

In current U.S. codes the intent is to provide 90% confidence of non-collapse forMCE shaking. In contrast, the current ACI 318-08 Equation 21-8 confinement trig-ger is based on 50% confidence of not exceeding the concrete crushing limit in theDesign Basis Earthquake (which is much lower shaking intensity than the MCE andis based on an assumed 5% damping ratio). To address these factors, the displace-ment used in Equation 21-8 should be increased; a factor of two is suggested.

Limits on story height to wall thickness hstw, such as the limit of 16 used in the1997 Uniform Building Code, should be considered to reduce the likelihood of walllateral instability failures. Use of a lower ratio might be appropriate for web bound-aries of flanged walls, where large cyclic tension and compressive strains occur.

ACKNOWLEDGEMENTS

The authors would like to individuals from three Chile universities who contributed gen-erously to the EERI team reconnaissance include:

From the Federico Santa Maria Technical University: Carlos Aguirre and ArturoMilln

From the Pontificia Universidad Catlica de Chile: Professors Juan Carlos De LaLlera, and Matias Hube, with Vicente Arizta, Juan Jose Besa, Alvaro Carboni,Claudio Frings, Juan Pablo Herranz, Cristbal Moena, Rodrigo Oviedo, VictorSandoval, Nicols Santa Cruz, Csar Seplveda, Benjamin Westenenk

From the Universidad de Chile: Professors Mara Ofelia Moroni, and RodolfoSaragoni

All of the members listed provided tremendous assistance, especially the students fromPontificia Universidad Catlica de Chile and Federico Santa Maria Technical University.

Travel funds were provided by the EERI Learning From Earthquakes program (NSFCMMI-0758529) and by NEEScomm (NSF CMMI-0927178). Opinions, findings, conclu-sions, and recommendations in this paper are those of the authors, and do not necessarilyrepresent those of the sponsor or others mentioned.

REFERENCES

American Concrete Institute (ACI 318-99), 1999. Building Code Requirements for StructuralConcrete (ACI 318-99) and Commentary (318R-99), Farmington Hills, MI, 391 pp.




American Society of Civil Engineers (ASCE 7-10), 2010. ASCE/SEI 7-10: Minimum DesignLoads for Buildings and Other Structures, Reston, VA, 650 pp.

Applied Technology Council (ATC 72-1), 2010. Modeling and Acceptance Criteria for SeismicDesign and Analysis of Tall Buildings, Report No. PEER/ATC-72-1, Applied TechnologyCouncil, Redwood City, CA, 242 pp.

Brown, J., and Kunnath, S. K., 2004. Low-cycle fatigue failure of reinforcing steel bars, ACIMaterials Journal 101, 457466.

Caldern, J. A., 2007. Update on Structural System Characteristics used in RC Building Con-struction in Chile, Civil Engineering Thesis, University of Chile, 76 pp. (in Spanish).

Comit Inmobiliario (CChC), 2010. Communication based on INE data, Instituto Nacional deEstadsticas (National Institute of Statistics), http://www.ine.cl/.

Centro de Investigacin, Desarrllo e Innovacin de Estructuras y Materiales (IDIEM), 2010. Peri-taje Estructural Edificio Alto Ro, Ciudad de Concepcin, Informe Final, Descripcin deCaida y Factores Asociados al Colapso, Revisin 1, Informe No 644.424-00, Universityof Chile (in Spanish).

Instituto Nacional de Normalizacin (INN), 2008. NCh430.Of2008: Reinforced Concrete Design and Calculation Requirements, Santiago, Chile, 17 pp. (in Spanish).

Instituto Nacional de Normalizacin (INN), 1996. NCh433.Of96: Earthquake Resistant Design ofBuildings, Santiago, Chile, 43 pp. (in Spanish).

Massone, L. M, Bonelli, P., Lagos, R., Lder, C., Moehle, J., and Wallace, J. W., 2012. Seismicdesign and construction practices for reinforced concrete structural wall buildings, EarthquakeSpectra 28, this issue.

Midorikawa, S., 1990. Ambient Vibration Tests of Buildings in Santiago and Via del Mar, DIENo.90-1, Departamento de Ingenieria Estructural, Pontificia Universidad Catlica de Chile,169 pp.

Moehle, J. P., Displacement-based design of RC structures subjected to earthquakes, EarthquakeSpectra 8, 403428.

Rodriguez, M. E., Botero, J. C., and Villa, J., 1999. Cyclic stress-strain behavior of reinforcingsteel including effect of Bbuckling, J. Struct. Eng., ASCE, 125, 605612.

Shibata, A., and Sozen, M. A., 1976. Substitute structure method for seismic design in R/C,Journal of the Structural Division, ASCE, 102, 118.

Thomsen, J. H., and Wallace, J. W., 2004. Displacement-based Design of slender RC structuralwalls experimental verification, J. Struct. Eng., ASCE, 130, 618630.

Uniform Building Code (UBC-97), 1997. Uniform Building Code, International Conference ofBuilding Officials, Whittier, California, 492 pp.

Wallace, J. W., 1995. Seismic design of RC shear walls; Part I: New code format, J. Struct. Eng.,ASCE, 121, 7587.

Wallace, J. W., 1996. Evaluation of UBC-94 provisions for seismic design of RC structural walls,Earthquake Spectra 12, 327348.

Wallace, J. W., and Moehle, J. P., 1989. The 3 March 1985 Chile Earthquake: StructuralRequirements for Bearing Wall Buildings, EERC Report No. UCB/EERC-89/5, EarthquakeEngineering Research Center, University of California, Berkeley.

S298 WALLACE ET AL.

Wallace, J. W., and Moehle, J. P., 1992. Ductility and detailing requirements of bearing wallbuildings, J. Struct. Eng., ASCE, 118, 16251644.

Wallace, J. W., and Moehle, J. P., 1993. An evaluation of ductility and detailing requirementsof bearing wall buildings using data from the 3 March 1985 Chile earthquake, EarthquakeSpectra 9, 137156.

Wallace, J. W., and Orakcal, K., 2002. ACI 318-99 provisions for seismic design of structuralwalls, ACI Structural Journal 99, 499508.

Wallace, J. W., and Thomsen, J. H., 1995. Seismic design of RC shear walls; Part II: Applica-tions, J. Struct. Eng., ASCE, 121, 88101.

Wood, S. L., Wight, J. K., and Moehle, J. P., 1987. The 1985 Chile Earthquake, Observations onEarthquake Resistant Construction in Via del Mar, Civil Engineering Studies, StructuralResearch Series No. 532, University of Illinois, Urbana, 176 pp.

(Received 26 July 2011; accepted 14 April 2012)


damage and implications for seismic design of rc structural wall building

Documents