DEFORMATION CAPACITY OF OLDER RC SHEAR WALLS: … · EC8-3 (CEN, 2005; CEN, 2010) in Annex A...

1

DEFORMATION CAPACITY OF OLDER RC SHEAR WALLS:

EXPERIMENTAL ASSESSMENT AND COMPARISON WITH

EUROCODE 8 - PART 3 PROVISIONS

Konstantinos CHRISTIDIS1, Emmanouil VOUGIOUKAS

2 and Konstantinos TREZOS

3

ABSTRACT

A series of six reinforced concrete shear walls was tested under cycling static loading and is presented

within this paper. The specimens represent shear walls designed according to older seismic codes,

characterized by the absence of confined (column like) boundary elements and by the low ratio of

shear reinforcement. The experimental behaviour of these specimens was evaluated in chord rotation

and shear strength terms and it was compared with the values given by the application of Eurocode 8 -

Part 3 (EC8-3) provisions, which refer to the assessment of reinforced concrete members.

INTRODUCTION

It is widely accepted that reinforced concrete shear walls play an important role in modern seismic

design, as they constitute a bracing system capable of contributing to the lateral seismic resistance of

the structure. However, it is well-known that in a lot of countries with high seismic exposure, there is a

significant number of existing buildings designed according to older seismic codes, which include

shear walls non-compliant with modern seismic provisions. These shear walls do not include confined

(column like) boundary elements and they are characterized by low ratios of shear reinforcement.

Theoretically, the lacking of confining stirrups leads to lower values of ultimate concrete strains,

therefore lowering ductility values, and in addition the sparse shear reinforcement does not ensure that

the shear strength exceeds the flexural one, leading to a prior shear (brittle) failure, often even before

flexural yielding. In addition, the absence of stirrups makes the walls vulnerable to buckling of the

compressive longitudinal rebars.

A lot of scientific researches have been conducted over the last years regarding the assessment

of existing reinforced concrete buildings. Several models have been developed, trying to predict the

deformation capacity and the degradation of the shear strength with the inelastic cyclic displacements

of existing reinforced concrete members (e.g. Priestley et al., 1994; Kowalsky and Priestley, 2000;

Panagiotakos and Fardis, 2001; Biskinis et al., 2004; Biskinis and Fardis, 2010a; Biskinis and Fardis,

2010b; Krolicki et al., 2011). Some of these models are adopted, with some modifications, from

modern seismic codes, such as EC8-3 (CEN, 2005; CEN, 2010) which includes a series of expressions

which focus on the assessment of existing reinforced concrete members. The main aim of the present

paper is to compare the experimental values which characterize the behaviour of the specimens with

the values given by the application of EC8-3 (CEN, 2005; CEN, 2010) provisions.

1 Ph.D. Candidate, National Technical University of Athens, Athens, [email protected]

2 Lecturer, National Technical University of Athens, Athens, [email protected]

3 Assistant Professor, National Technical University of Athens, Athens, [email protected]

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EUROCODE 8 - PART 3 PROVISIONS

EC8-3 (CEN, 2005; CEN, 2010) in Annex A proposes models for the performance assessment of reinforced concrete members that refer to their deformation capacity and their shear strength.

The deformation capacity of reinforced concrete members is defined in terms of the chord rotation, θ, i.e. the angle between the tangent to the axis at the yielding end and the chord connecting that end with the end of the shear span (LV = M/V = moment/shear at the end section). The chord rotation is, simultaneously, equal to the element drift ratio, i.e. the deflection at the end of the shear span with respect to the tangent to the axis at the yielding end divided by the shear span. EC8-3 (CEN, 2005; CEN, 2010) defines three different limit states, indicating the state of damage of the concrete member – Damage Limitation (DL), Significant Damage (SD) and Near Collapse (NC).

The value of the chord rotation, θy, in the limit state of Damage Limitation (DL) for rectangular walls is given in Eq.(1) and it corresponds to the yielding moment capacity, My.

θy=φyLV+αVz

3+0.0013+φy

dbLfy8 fc

(1)

where: φy is the yield curvature of the end section, αVz is the tension shift of the bending moment diagram (Eurocode 2 - Part 1-1 (CEN, 2004) with,

z the length of internal lever arm, taken equal to 0.8h in walls with rectangular section (h is the depth of cross-section), αV=1 if shear cracking is expected to precede flexural yielding at the end section (i.e. if My>LVVR,c ); otherwise (i.e. if My

K. Christidis, E. Vougioukas and K. Trezos 3

transverse deflection of the shear span or of the chord rotation at member end: μΔpl= μΔ-1. This ductility can be calculated as the ratio of the plastic part of the chord rotation, θumpl=θum-θy, normalized to the chord rotation at yielding, θy.

The shear strength as controlled by the stirrups, VR, can be derived by Eq.(3) (with units: MN and meters).

VR=1γel

h-x2LV

min N;0.55Acfc + 1-0.05min 5;μΔpl ∗

∗ 0.16max 0.5;100ρtot 1-0.16min 5;LVh

fcAc+Vw (3) where: γel is equal to 1.15 for primary seismic elements and 1.00 for secondary seismic elements, x is the compression zone depth, N is the compressive axial force (positive, taken as being zero for tension), Ac is the cross-section area, equal to bwd for a cross-section with a rectangular web of width (thickness) bw and structural depth d ρtot is the total longitudinal reinforcement ratio, Vw is the contribution of transverse reinforcement to shear resistance, equal to: Vw = ρwbwzfyw (ρw is the transverse reinforcement ratio)

The shear strength as controlled by the web crushing, VR,max, can be derived by Eq.(4) (with units: MN and meters).

VR,max=0.85 1-0.06min 5;μΔ

pl

γel1+1.8min 0.15;

NAcfc

∗

∗ 1+0.25max 1.75;100ρtot 1-0.2min 2;LVh

fcbwz (4) where: γel is equal to 1.15 for primary seismic elements and 1.0 for secondary seismic elements and the other variables as previously defined.

According to the above provisions the final deformation capacity could be predicted using a common diagram which includes both flexural and shear strength capacities expressed in Load- Rotation terms (P-θ) (Fig.1).

Figure 1. Definition of failure point according to EC8-3 (CEN, 2005; CEN, 2010)

θumθy

(NC)

Flexural behaviour

Shear behaviour(DL)

(SD)

θu,SDθfail

Failure point

Py

Pfail

Pu,SDPum

Chord rotation

Load

4

CHARACTERISTICS OF SPECIMENS

The experimental process included the testing of a series of six older reinforced concrete shear walls. Four of these specimens were previously presented in Christidis et al. (2013), while the last two are first presented within this paper. The dimensions and the reinforcement configuration of all specimens are shown in Fig.2 and Fig.3 and they are summarized in Table.1.

Figure 2. Dimensions of specimens (in m)

It is noted that, as it was mentioned before, in older reinforced concrete shear walls one of the most important causes of the loss of their bearing capacity is the buckling of the compressive rebars, which reduce the bearing capacity of the reinforcement not only in compression but in tension too. In addition, the loss of the bearing capacity due to buckling often occurs quite prematurely, making the wall unable to develop its plastic characteristics. Thus, many times, even in walls with low shear strength, the buckling phenomenon appears to be the cause of the loss of their capacity. However, the specific series of tests focuses, mainly, on the influence that shear reinforcement has at the behaviour of each specimen. For this purpose two layers of Glass Fiber Reinforced Polymers (GFRP) were placed in specimens W2-FRP, and W4-FRP while in specimens W5,W6,W9,W11 a configuration of open stirrups was added. With these two different interventions the buckling phenomenon of compressive reinforcement rebars was avoided.

Table 1. Reinforcement ratios and material properties of specimens

Concrete compression

strength fc (MPa)

Reinforcement yield/failure fy/fu (MPa)

Longitudinal reinforcement ratio

(uniform) ρtot (‰)

Transverse (shear) reinforcement ratio

ρw (‰) Longitudinal Reinforcement Stirrups

Wall W2-FRP 32.12 610/670 622/733 15.28 3.35 Wall W4-FRP 30.12 610/670 622/733 15.28 3.35

Wall W5 33.55 586/699 505/568 15.28 1.88 Wall W6 26.35 575/694 516/569 15.28 1.88 Wall W9 31.12 580/670 588/681 12.06 2.01 Wall W11 31.12 580/670 568/654 12.06 1.13

ρtot=Area sum of longitudinal reinforcement/cross section area=ΣAs,L/(bw*h) ρw=Area sum of stirrup/(wall width*stirrups distance)=ΣΑs,w/(bw*s)

L1

L6

L3

bw

Lw

L2

L5

Lw

L2

L5

L1 L2 L3 L4 L5 L6 h bw LwW2-FRP

W4-FRP

W5-FRPW6-FRPW9W11

1.40

1.40

1.401.401.401.40

0.10

0.10

0.100.100.1250.125

0.74

0.74

0.740.740.750.75

1.20

1.20

1.201.201.201.20

0.20

0.20

0.200.200.200.20

0.30

0.30

0.500.500.300.30

1.40

1.40

1.501.501.501.50

0.60

0.60

0.600.600.600.60

0.50

0.50

0.500.500.500.50

L4

h


Figure 3. Reinforcement configuration of specimens (in mm)

APPLICATION OF EUROCODE 8 - PART 3 PROVISIONS

The first step of this paper was to apply EC8-3 (CEN, 2005; CEN, 2010) provisions in order to predict the deformation capacity of each specimen. The application of Eq.(1) requires the calculation of the yield curvature, φy. However, the fact that there are several layers of reinforcement does not lead to a clear yield point. Thus, and since EC8-3 (CEN, 2005; CEN, 2010) does not give specific instructions for that case, some assumptions had to be made in order to proceed the application. Finally, it was decided to search for a conventional yield point coming from the bilinearization of the Moment-Curvature diagram, M-φ, using an equivalent bilinear curve (elastic-plastic with hardening) where the elastic stiffness is defined from the first yield point (Py1, φy1). The conventional yield point (Py, φy) is

Wall W11Wall W9

Walls W2-FRP,W4-FRP Walls W5,W6

D6/300 stirrupsD8/300 openstirrups

3x2D12

2D12 2D12

D8/300 stirrups

3x2D12

Section A-A'

2D12

GFRP400

D8/400 stirrups

D8/100 openstirrups

3x2D12

2D12 2D12

D6/400 stirrups

D8/100 openstirrups

3x2D12

2D12 2D12

2D12

A'

A A' A A'

A A'

D12

D8/300

4x175

D8/400D8/100

D6/300D8/300

D12

D12

4x175

D12

D12

D6/400D8/100

Section A-A'

4x173 4x173Section A-A' Section A-A'

A

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

D12

6

derived by equating the area (energy) of the two curves, as shown in Fig.4. The most characteristic values of the diagrams shown in Fig.4 are summarized in Table.2.

(a) Wall W2-FRP (b) Wall W4-FRP

(c) Wall W5 (d) Wall W6

(e) Wall W9 (f) Wall W11

Figure 4. Analytical Moment-Curvature curve and bilinearization of specimens

Table 2. Theoretical chord rotation ductility according to EC8-3 (CEN, 2005; CEN, 2010)

1st Yield Conventional Yield Ultimate

My1 (kNm)

Py1=My1/LV (kN)

φy1 (1/m)

My (kNm)

Py=My/LV (kN)

φy (1/m)

Mu (kNm)

Pu=Mu/LV (kN)

φu (1/m)

Wall W2-FRP 152.76 101.84 0.0056 191.03 127.35 0.0071 218.54 145.69 0.0259 Wall W4-FRP 151.89 101.26 0.0057 189.25 126.17 0.0071 217.49 144.99 0.0252

Wall W5 147.45 98.30 0.0054 185.90 123.93 0.0068 214.46 142.98 0.0268 Wall W6 173.75 115.83 0.0059 200.53 133.69 0.0068 235.08 156.72 0.0192 Wall W9 148.79 99.19 0.0053 188.66 125.77 0.0067 217.29 144.86 0.0288 Wall W11 148.79 99.19 0.0053 188.66 125.77 0.0067 217.29 144.86 0.0288

The diagrams of Fig.5 are derived from the application of Eq.(1), Eq.(2) and Eq.(3) where the

safety factor γel is taken equal to 1. It is noted that in these diagrams only the shear strength as controlled by the stirrups, VR, is included as the shear strength as controlled by the web crushing, VR,max, is in all cases higher than the flexural strength. As shown in this figure for walls W5, W6 and W11, EC8-3 (CEN, 2005; CEN, 2010) predicts a prior brittle failure which, in one case (W11), occurs even before the yielding of the wall. The ductility values calculated according to the previous methodology are included in Table.3.

0

50

100

150

200

250

0.000 0.005 0.010 0.015 0.020 0.025 0.030

Mom

ent(

kNm

)

Curvature (1/m)

Analytical Moment-Curvature curveBilinearization

0

50

100

150

200

250

0.000 0.005 0.010 0.015 0.020 0.025 0.030

Mom

ent(

kNm

)

Curvature (1/m)


0

50

100

150

200

250

0.000 0.005 0.010 0.015 0.020 0.025 0.030

Mom

ent(

kNm

)

Curvature (1/m)


0

50

100

150

200

250

0.000 0.005 0.010 0.015 0.020 0.025 0.030

Mom

ent(

kNm

)

Curvature (1/m)


0

50

100

150

200

250

0.000 0.005 0.010 0.015 0.020 0.025 0.030

Mom

ent(

kNm

)

Curvature (1/m)


0

50

100

150

200

250

0.000 0.005 0.010 0.015 0.020 0.025 0.030

Mom

ent(

kNm

)

Curvature (1/m)






Figure 5. Deformation capacity of specimens according to EC8-3 (CEN, 2005; CEN, 2010)

Table 3. Theoretical chord rotation ductility according to EC8-3 (CEN, 2005; CEN, 2010)

Conventional Yield Ultimate (failure) Ductility

μ=θum/θy Py (kN) θy Pu (or Pfail)

(kN) θum (or θfail)

Wall W2-FRP 127.35 0.0074 145.69 0.0158 2.14 Wall W4-FRP 126.17 0.0074 144.99 0.0156 2.11

Wall W5 123.93 0.0071 132.36 0.0110 1.55 Wall W6 133.69 0.0072 148.52 0.0115 1.60 Wall W9 125.77 0.0070 144.86 0.0156 2.23 Wall W11 125.77 0.0070 117.82 0.0066 1.00

Note that in Walls W5, W6 and W11 the failure point (Pfail, θfail) is derived from the intersection of curves of flexural and shear capacity and it is prior to ultimate point (Pu, θu). In addition, in wall W11 where the failure rotation, θfail, is lower than the yielding chord rotation, θy, the ductility value is taken equal to 1.

EXPERIMENTAL PROCESS

The specimens were tested as cantilevers, under static cyclic loading. During the testing procedure, the displacement control method was adopted, including an initial displacement of ±10 mm with steps of

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0.000 0.005 0.010 0.015 0.020

Loa

d(k

N)

Chord Rotation

Flexural capacityShear capacity

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0.000 0.005 0.010 0.015 0.020

Load

(kN

)

Chord Rotation

Flexural capacity

Shear capacity

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0.000 0.005 0.010 0.015 0.020

Load

(kN

)

Chord Rotation

Flexural capacity

Shear capacity

Failure Point

θfail=0.01100

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0.000 0.005 0.010 0.015 0.020Lo

ad(k

N)

Chord Rotation

Flexural capacity

Shear capacity

Failure Point

θfail=0.0115

ν=0.070

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Load

(kN

)

Chord Rotation

Flexural capacity

Shear capacity0

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0.000 0.005 0.010 0.015 0.020

Loa

d(k

N)

Chord Rotation

Flexural capacityShear capacity

Failure Point

θfail=0.0066

8

10 mm until failure. Three cycles of each displacement group were applied. Time - history of the loading is shown in Fig.6. Axial load was also applied to wall W6. The axial load had an initial value of ν=0.07 and during the experiment had a variation as pointed in Fig.7d.

Figure 6. Time history of loading of specimens

EXPERIMENTAL RESULTS

The experimental results are presented in Fig.7 as a Load-Displacement diagram, P-δtop, where δtop is the displacement at the top of each specimen.




Figure 7. Lateral load - Top displacement curves of specimens

LV

time

+10+20+30+40+50

-10-20-30-40-50

DisplacementD

ispl

acem

ent (

mm

)

-200

-150

-100

-50

0

50

100

150

200

-60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60

Loa

d (k

N)

Displacement (mm)-200

-150

-100

-50

0

50

100

150

200

-60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60

Loa

d(k

N)

Displacement(mm)

-200

-150

-100

-50

0

50

100

150

200

-60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60

Loa

d (k

N)


-150

-100

-50

0

50

100

150

200

-60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60

Load

(kN

)

Displacement (mm)

ν=0.100

ν=0.070

ν=0.085

ν=0.060

ν=0.100ν=0.085

ν=0.060

-200

-150

-100

-50

0

50

100

150

200

-60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60

Loa

d (k

N)


-150

-100

-50

0

50

100

150

200

-60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60

Load

(kN)

Displacement (mm)


In Fig.8 the experimental envelopes of all specimens are presented in a common Load-Chord rotation (or drift) diagram, P-θ, where θ is defined as the experimental top displacement normalized to the shear span (θ=δtop/ LV).

Figure 8. Experimental envelopes, P-θ, of specimens

In order to compare experimental and theoretical values, the experimental envelopes had first to be edited so as to define an experimental conventional yield and failure point. The experimental ductility, in chord rotation terms, is derived from the bilinearization of the experimental envelope of each specimen (P-θ), assuming an equivalent bilinear curve (elastic-plastic with hardening). The initial stiffness of the bilinear curve is obtained from the point where the theoretical first yield load, Py1, (Table.2) intersects the experimental curve and the ultimate rotation θum,exp corresponds to 0.80Pmax,exp, where it is assumed that failure occurs. Finally, the conventional experimental yielding (Py,exp, θy,exp) point is arrived at by equating the area (energy) of the two curves (Fig.9).

Figure 9. Bilinearization of experimental envelope, P-θ

Applying the above procedure to the experimental envelopes of all specimens the curves of Fig.10 are derived. The theoretical values and experimental values obtained from Fig.10 are compared in Table.4.

0

50

100

150

200

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035

Load

(kN

)

Chord rotation

W11W9W2-FRP

W4-FRPW5

W6

P-θ envelope

Equivalent

bilinear curve

θy' θy,exp θum,exp

Py1

Py,exp 0.80Pmax,exp

Chord rotation

Load

Pmax,exp

10



(e) Wall W9 (f) Wall W11 Figure 10. Experimental envelope and bilinearization of specimens

Table 4. Experimental and theoretical chord rotation ductility

Chord Rotation Load

Experimental EC8-3 (from Table.3) Experimental EC8-3

(from Table.3)

θy,exp θum,exp Ductility μexp

θy θum (or θfail) Ductility

μ Pmax,exp (kN)

Pu (or Pfail) (kN)

Wall W2-FRP 0.0074 0.0270 3.64 0.0074 0.0158 2.14 165.97 145.69 Wall W4-FRP 0.0070 0.0320 4.57 0.0074 0.0156 2.11 171.20 144.99

Wall W5 0.0068 0.0197 2.90 0.0071 0.0110 1.55 153.48 132.36 Wall W6 0.0053 0.0144 2.72 0.0072 0.0115 1.60 158.92 148.52 Wall W9 0.0087 0.0297 3.41 0.0070 0.0156 2.23 177.02 144.86 Wall W11 0.0060 0.0300 5.00 0.0070 0.0066 1.00 173.31 117.82

DISCUSSION OF RESULTS

As concluded from the experimental results, the different ratios of shear reinforcement do not seem to affect the load capacity of shear walls. This was clearly demonstrated by the fact that all specimens exhibited a level of lateral load capacity close to their flexural strength, not affected by the amount of shear reinforcement. As for the post yield behaviour, namely the deformation capacity, the results of walls W5 to W11 lead to conflicting conclusions. Walls W5 and W6 exhibited a post yield behaviour

0

50

100

150

200

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035

Loa

d(k

N)

Chord Rotation

Experimental envelopeBilinearization

0

50

100

150

200

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035

Load

(kN

)

Chord Rotation


0

50

100

150

200

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035

Loa

d(k

N)

Chord Rotation


0

50

100

150

200

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035Lo

ad(k

N)

Chord Rotation

Experimental envelope

Bilinearization

0

50

100

150

200

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035

Load

(kN

)

Chord Rotation


0

50

100

150

200

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035

Loa

d(k

N)

Chord Rotation



with a descending branch, whilst walls W9 and W11, with the same or even less shear reinforcement, exhibited a clear flexural behaviour. The different behaviour of walls W5 and W6 compared to walls W9 and W11 may be due to the difference in wall thickness (10 cm and 12.5 cm respectively) and the influence of accidental eccentricities. However, all four specimens (wall W5 to W11) exhibited a similar cracking mode, being characterized by the formation of significant inclined cracks (in contrast to W2-FRP and W4-FRP with higher amount of shear reinforcement - Table.1) which were, in one case (Wall W11), followed by the failure of one of the stirrups located approximately at the middle height of the wall web. The deformation capacity of shear walls is directly connected to the shear strength. As concluded from the comparison of theoretical and experimental results, EC8-3 (CEN, 2005; CEN, 2010) provisions (Eq. (3)) seem to underestimate the shear strength of walls as in two cases (walls W5 and W6) the calculated shear strength intersects the flexural strength curve before the ultimate chord rotation, θum, and in one case (wall W11) the calculated shear strength is lower even than the yield flexural strength, predicting values of ductility lower than the experimental ones (μexp is 87%, 70% and 400%, respectively, higher). At this point, some comments have to be made. The first one concerns the method of bilineazisation of the M-φ diagram. Within this paper the calculations were made assuming an equivalent elastic-plastic with hardening diagram. However, the use of an elastic-perfectly plastic diagram may lead, in some cases, to different theoretical values. For example, assuming an elastic-perfectly plastic M-φ diagram for wall W6 (Fig.11a) a theoretical ductility value μ=1.81 is calculated (13% higher than μ=1.60 with elastic-plastic with hardening diagram). The second comment concerns the value of z=0.8h which in EC8-3 (CEN, 2005; CEN, 2010) is defined as the lever arm of internal forces. This value is quite accurate for beams, but not for walls, especially when the longitudinal reinforcement is distributed uniformly. Applying the value z=0.57h as derived from the cross-section analysis in wall W9 (Fig.11b) a theoretical ductility μ=1.29 is calculated (73% lower than μ=2.23 with z=0.8h).

(a) Wall W6 (b) Wall W9 Figure 11. Deformation capacity of specimens (a) W6 assuming an equivalent elastic- perfectly plastic M-φ curve

and (b) W9 for z=0.57h from cross-section analysis, according to EC8-3 (CEN, 2005; CEN, 2010)

Finally, in walls W2-FRP, W4-FRP and W9 where the calculated shear strength is always higher than the flexural one, the experimental ductility is 70%, 117% and 53%, respectively, higher than the theoretical one. As shown in Table.4, while there is a good convergence between experimental and theoretical values of θy, the experimental values of θum are approximately twice the theoretical ones. This remark leads to the conclusion that Eq.(2) underestimates the ultimate chord rotation, θum.

CONCLUSIONS

Six reinforced concrete shear walls, representing shear walls of existing buildings designed according to older codes, were tested under static cyclic loading. As concluded from the experimental results, low ratios of shear reinforcement seem to affect neither the bearing capacity nor the deformation capacity of walls. All the specimens reached their flexural bearing capacity, while in most cases they exhibited a flexural post yield behaviour followed by significant values of ductility. Low ratios of

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Load

(kN

)

Chord Rotation

Flexural capacity

Shear capacity0

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Load

(kN

)

Chord Rotation

Flexural capacity

Shear capacity

Failure Point

θfail=0.0085

12

shear reinforcement seem to determine the crack mode of walls which was characterized by the formation of significant inclined cracks. However, it is noted that these cracks were not followed by loss of bearing capacity. The experimental results were compared with the provisions included in EC8-3 (CEN, 2005; CEN, 2010), which refer to the assessment of reinforced concrete members. EC8-3 (CEN, 2005; CEN, 2010) seems, in some cases, to underestimate the shear strength of the walls indicating a prior brittle failure mode, therefore poor deformation capacity, not confirmed by the experimental results obtained in the present research. In addition, EC8-3 (CEN, 2005; CEN, 2010) underestimates the ultimate chord rotation, θum. However, in general, all the above provisions underestimate the behaviour of shear walls, thus they are on the safe side.

REFERENCES

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DEFORMATION CAPACITY OF OLDER RC SHEAR WALLS: … · EC8-3 (CEN, 2005; CEN, 2010) in Annex A...

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