Section C6.7.1 of ASCE/SEI 41-06,16 “Seismic Rehabili-tation of Existing Buildings,” categorizes the response of RC walls based on aspect ratio. According to ASCE/SEI 41-06, walls with an aspect ratio of less than 1.5 are considered to be squat and their response is assumed to be controlled by shear. Conversely, walls with an aspect ratio of greater than 3 are assumed to be slender and their response is consid-ered to be controlled by flexure. The response of walls with intermediate aspect ratios is considered to be controlled by both flexure and shear. As the walls presented herein had
aspect ratios ranging between 1.35 and 1.85, the response of the walls was supposed to be controlled by shear and by both shear and flexure, respectively, according to ASCE/SEI 41-06.
Wall strength
ASCE/SEI 41-06 refers to Chapter 21 of ACI 318-11 to determine the shear strength of existing walls. The nominal shear strength of walls is given in ACI 318-11, Eq. (21-7), as
Vn =Acv(ac fc′ +ρt ytf ) (1) Fig. 6—Lateral-force, top-displacement responses of test specimens.
This nominal shear strength is limited to 0.083Acv√fc′ (MPa) (Acv√fc′ [psi]). The coefficient αc is 0.25 for hw/lw ≤ 1.5, is 0.17 for hw/lw ≥ 2.0, and varies linearly between 0.25 and 0.17 for intermediate hw/lw values. When determining the yield and nominal shear strengths, ASCE/SEI 41-06 limits the strength of the transverse reinforcing bars fyt to the specified yield strength.
The nominal shear-friction strength of walls across a sliding plane perpendicular to shear-friction reinforcing bars is given by ACI 318-11, Eq. (11-25), as
Vn = (Avffy + N*)μ (2) The nominal shear-friction strength given in Eq. (2) is limited to the smaller of 0.2Acfc′ (N) (0.2Acfc′ [lb]) and 5.5Ac
(N) (800Ac [lb]).
For the flexural strength of existing walls, ASCE/SEI 41-06 refers to the basic principles outlined in Chapter 10 of ACI 318-11, but requires the use of expected yield strengths of the longitudinal reinforcing bars instead of specified minimum yield strengths. When determining flexural yield strengths of walls with no boundary members, ASCE SEI 41-06 requires considering only the longitudinal reinforcing bars within the outer 25% of the wall cross section, but when determining the nominal flexural strengths, the contributions of all longi-tudinal reinforcing bars within the wall component cross section need to be considered.
As shown in Table 3, the lateral-force-carrying capacity of the test specimens was limited by their flexural strength and, thus, categorizing the response of RC walls by ASCE/SEI 41-06 based on aspect ratio only is found to be misleading.
The results also show that the “plane sections remain plane”
hypothesis, which was used during the calculation of the flexural strengths, provided strengths that agree well with those determined experimentally, even for squat walls with an aspect ratio of 1.35. In addition, prior to the peak strength of the test specimens being attained, the strain gauge read-ings from longitudinal reinforcement of most test specimens generally varied linearly across the section.
Required length of splices
Current provisions for the required length of splices are based on studies conducted to determine bond-slip relation-ships between isolated reinforcing bars and the surrounding concrete. Transfer of force between starter and main rein-forcing bars over the provided splice length involves a different force transfer mechanism than that occurring between an isolated reinforcing bar and the surrounding concrete, but it is widely accepted that the required length of splices is the same as the required development lengths of single embedded reinforcing bars.24 Accordingly, ASCE/
SEI 41-06 refers to the provisions for required length of tension splices of ACI 318-11, which are based on require-ment for tension developrequire-ment length. ASCE/SEI 41-06 specifies that the required splice length of plain round rein-forcing bars to be taken as twice that required for deformed reinforcing bars. According to ACI 318-11, Eq. (12-1), the required length of Class B deformed reinforcing bar splices
subjected to tension, which is required to be at least 300 mm (12 in.), is
The length of deformed reinforcing bar splices subjected to compression can be determined from Section 12.16.1 of ACI 318-11 as
This development length is required to be at least 300 mm (12 in.) and is required to be increased by one-third for a concrete compressive strength of less than 21 MPa (3.0 ksi).
When the splice length of reinforcing bars within an existing wall is found to be inadequate, ASCE/SEI 41-06 stipulates the maximum stress that can be developed within the spliced bars to be determined as follows
fs = (lb/ld)fy ≤ fy (5) Ratios of the maximum stresses, which can be developed by spliced plain round bars according to Eq. (5) to the corre-sponding yield strengths of the spliced bars, are presented in Table 4. Although the provided splice lengths were signifi-cantly less than those required by Eq. (5) to develop the yield strength of the spliced bars, the walls were able to develop 97 to 102% of their computed flexural strength (refer to Table 3), which was determined assuming that the lap splices would develop the yield strength of the spliced bars. Peak strengths reached by the test specimens with lap-spliced reinforcing bars were underestimated by an average of 41%
(refer to Table 5) by predictions made using Eq. (5). Similar findings were previously reported in studies24,25 under-taken to investigate the behavior of columns with short lap splices. The lateral force capacity of columns investigated by Cho and Pincheira24 was underestimated using Eq. (5) by an average of 28%. Similarly, columns tested by Melek and Wallace25 were reported to have achieved 97 to 103%
Table 4—Ratios of maximum stress to yield stress for spliced bars
Test specimen
φ10 reinforcing bars φ12 reinforcing bars fs/fy* fs/fy† fs/fy* fs/fy†
WPS5 0.51 0.80 0.44 0.73
WPS6 0.51 0.80 0.44 0.73
WPS7 0.51 0.80 0.77 1.00
WPS8 0.51 0.80 0.77 1.00
*Ratio according to Eq. (5).
†Ratio according to Eq. (6).
of their yield strengths, which were calculated assuming that the lap splices would develop the yield strengths of the spliced bars, but the provided splice lengths were approxi-mately 67% of that required by ASCE/SEI 41-06.
Because Eq. (5) was found to be excessively conserva-tive, the provided splice lengths used in the study reported herein were also compared with those implied by a proposed supplement26 to ASCE/SEI 41-06. The proposed equation, which is a modified version of Eq. (5) and based on the work of Cho and Pincheira,24 is
f l
l f f
s
b d
y y
=
≤
1 25
0 67
.
.
(6)
In most cases, the provided splice lengths were shorter than required by Eq. (6) (refer to Table 4) to develop the full strength of the spliced bars. Predictions made using Eq. (6) underestimated the peak strengths by an average of 21%.
As discussed previously, both Eq. (5) and Eq. (6) predict slip to occur at force levels less than the yield strength of the spliced bars. However, during testing, all of the monitored splices developed tensile stresses that were greater than the experimentally determined yield strength of the spliced bars.
The maximum stresses that were measured during testing were converted to maximum bond stresses as follows
u f d l
s b d
= 4 (7)
Using Eq. (3) and (7), the ASCE/SEI 41-06 implied maximum bond stress that was expected to develop between a plain round bar subjected to tension and the surrounding concrete was determined as
u f c K
d
c
t e s
b tr
b
= ′ +
1
2 1 4
1 1.
ψ ψ ψ l (8)
Maximum bond stresses developed between the spliced bars and the surrounding concrete are compared in Fig. 7, with maximum bond stress values implied by ASCE/SEI 41-06 and the proposed supplement to ASCE/SEI 41-06.
The ASCE/SEI 41-06 implied average maximum bond stress of 0.29√fc′ (MPa) (3.49√fc′ [psi]) is significantly less than the measured average bond stress of 0.57√fc′ (MPa) (7.95√fc′ [psi]). The average maximum bond strength implied by the proposed supplement to ASCE/SEI 41-06, Table 5—Measured and calculated strengths of test specimens with spliced longitudinal reinforcement
Test specimen VTest, kN Vn,F-A/S*, kN Vn,F-PS†, kN Vn,F‡, kN VTest/Vn,F-A/S VTest/Vn,F-PS VTest/Vn,F
WPS5 199 91 129 195 2.14 1.51 1.02
WPS6 194 91 144 197 2.16 1.37 0.98
WPS7 231 181 208 233 1.29 1.12 0.99
WPS8 271 234 262 278 1.19 1.06 0.97
Average 1.70 1.27 0.99
Standard deviation 0.46 0.18 0.02
†Nominal flexural capacity according to ASCE/SEI 41-06.
‡Nominal flexural capacity according to proposed supplement26 to ASCE 41-06.
*Nominal flexural capacity according to ACI 318-11 and assuming splices would develop yield strength of spliced bars.
Note: 1 kN = 0.225 kip.
Fig. 7—Normalized maximum bond stresses that developed between spliced bars and surrounding concrete.
which was dependent on the provided splice lengths, was 0.46√fc′ (MPa) (5.54√fc′ [psi]).
The performance of the splices reported herein was better than had been expected, principally because of the exces-sively conservative requirement of current design codes19,20 and assessment recommendations16 for the required splice length of plain round bars. For example, ACI 318 requires the length of splices to be increased by one-third if all of the longitudinal bars are spliced at the same location. This requirement is not based on strength criteria,27 but is primarily intended to encourage designers to stagger bar splices. This requirement alone results in underestimating the capacity of spliced reinforcing bars in existing structures by 23%.
In some cases encountered herein, the compression splice length requirements of ACI 318 governed the required length of splices. The compression splice length require-ments of ACI 318 have remained essentially the same since the 1963 version of the code and appear to be conservative.
Required tension splice lengths are expected to govern the required length of splices, as the formation of transverse tension cracks around a splice that is subjected to tension reduces bond strength and, thus, increases the length that is required to allow the splice to transfer the desired magnitude of stress. Chun et al.28 have recently reported a better perfor-mance of compression splices than tension splices, which was principally attributed to end bearing of reinforcing bars when subjected to compression. Similarly, Cairns29 has found the equations contained in ACI 318-11 to be conserva-tive, and has proposed that the length of compression splice for deformed bars be taken as 30% shorter than that required by the equations.
In addition to the aforementioned excessively conservative requirements of ACI 318 for deformed bars, ASCE 41-06 recommends taking the required splice length of plain round bars as twice that required for deformed bars, which results in significantly underestimating the capacity of plain round splices located in existing structures.
The good performance of the splices reported herein was not considered to be influenced by the thick concrete cover that the splices were provided with. ACI 318 employs parameters for concrete cover/reinforcement spacing cb
and transverse reinforcement index Ktr to account for the confinement term (cb + Ktr)/db, and in ACI 318, it is assumed that an increase in the value of this confinement term, above the maximum allowed 2.5, is not likely to increase anchorage capacity and is not likely to prevent a pullout failure, which is a failure mode typically exhibited by plain round bars. Similarly, Eligehausen et al.30 reported that an increase in concrete cover, reinforcement spacing, or transverse confinement can prevent concrete splitting fail-ures only, which is a failure mode typically sustained when using spliced deformed bars. The maximum allowed value of (cb + Ktr)/db = 2.5 was employed when calculating the ASCE/SEI 41-06 implied bond strength of the plain round bars discussed herein.
CONCLUSIONS
Twelve test specimens replicated from wall segments of an existing building were experimentally tested to assess the
seismic behavior of RC walls constructed before the intro-duction of seismic design requirements in the New Zealand Standard Model Building By-law.14 The results of the exper-imental tests were evaluated and compared with current assessment provisions.
Based on the study presented herein, the following conclu-sions are drawn:
1. The lateral-force-carrying capacity of lightly reinforced existing walls is limited by their flexural strength. Owing to the low quantity of reinforcing bars provided, yielding of the longitudinal reinforcing bars dictates the strength of this wall type.
2. The strength and the stiffness of this type of wall degrade rapidly and significantly. The walls have limited energy dissipation capacity, principally due to the provision of few longitudinal reinforcing bars. In addition, the walls suffer from a lack of transverse confinement reinforcement to contain concrete in compression zones and to prevent longitudinal reinforcing bars from buckling.
3. During testing, the wall pier specimens having no axial load exhibited cracks that were wide during low-level drift cycles, but the cracks, which were located near the supports, closed up and appeared inconspicuous after the tests were completed, with longitudinal reinforcement that had previ-ously yielded in tension hidden in the cracks. This type of crack could easily be overlooked and the walls may appear intact during post-earthquake inspections, even if the stiff-ness and the strength of the walls deteriorated significantly.
4. From peak stresses measured during testing, it was shown that the provisions contained in ASCE/SEI 41-06 significantly underestimate the maximum stresses that can be developed by plain round reinforcing bar lap splices. The relatively recent recommendations of a proposed supple-ment to ASCE/SEI 41-06 also underestimate the maximum stresses that can be developed, but to a lesser extent.
5. During testing, the lap splices were able to develop bond stresses that were significantly higher than the maximum possible bond stresses implied by ASCE/SEI 41-06. Further research is recommended, as the provisions of ASCE/SEI 41-06 for required splice lengths of plain round reinforcing bars are based on studies conducted for deformed bars. The provisions are excessively conservative and potentially lead to unnecessary or expensive seismic retrofitting solutions.
The provisions also lead to the potential for overlooking the danger of existing walls failing in shear during an earthquake, before the actual strength of the tension splices is exceeded.
ACKNOWLEDGMENTS
The authors would like to gratefully acknowledge the financial support provided by the New Zealand Foundation for Research, Science and Tech-nology (FRST) through Grant UOAX0411.
AUTHOR BIOS
Adane Gebreyohaness is a Structural Engineer at Beca Limited in New Zealand. His research interests include the seismic assessment, strength-ening, and design of reinforced concrete and steel structures.
Charles Clifton is an Associate Professor at the University of Auckland, Auckland, New Zealand. His research interests include the performance of steel and composite steel/concrete buildings in severe earthquakes and severe fires.
John Butterworth is an Associate Professor at the University of Auckland.
His research interests include nonlinear structural and solid mechanics;
stability; buckling behavior of thin-walled sections; structural dynamics;
earthquake engineering; base isolation using rolling, sliding, and rocking mechanisms; passive control of structure response (especially by energy dissipating joints); experimental dynamics; pounding of bridges and build-ings; and assessment and retrofit of steel structures.
ACI member Jason Ingham is an Associate Professor and Deputy Head (Research) of the Department of Civil and Environmental Engineering at the University of Auckland. His research interests include seismic assess-ment, retrofit and design of reinforced and prestressed concrete structures, and sustainable concrete technology.
NOTATION
Ab = area of boundary reinforcement
Ac = area of concrete section resisting shear transfer Acv = shear area
Al = area of longitudinal reinforcing bar At = area of transverse reinforcing bar Avf = area of shear friction reinforcement
cb = smaller of: (a) distance from center of bar or wire to nearest concrete surface; and (b) one-half the center-to-center spacing of bars or wires being developed
db = nominal bar diameter
fc′ = specified compressive strength of concrete fs = splice strength
fult = ultimate strength of reinforcement fy = yield strength of reinforcement fyt = yield strength of shear reinforcement hw = height of wall
Ktr = transverse reinforcement index l = length of splice
lb = provided splice length
ld = required splice length according to ACI 318-11 lw = horizontal length of wall
M/Vlw = shear span-to-depth ratio N* = design axial load u = bond stress
Vn = nominal shear strength VTest = peak strength of test specimen Vu = factored shear force
λ = modification factor related to density of concrete μ = coefficient of friction
ρl = ratio of area of distributed longitudinal reinforcement to gross concrete area perpendicular to that reinforcement
ρt = ratio of area of distributed transverse reinforcement to gross concrete area perpendicular to that reinforcement
ψe = modification factor based on reinforcement coating ψs = modification factor based on reinforcement size ψt = modification factor based on reinforcement location
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