13th World Conference on Earthquake Engineering Vancouver, B.C., Canada

August 1-6, 2004 Paper No. 1597

EXPERIMENTAL STUDY FOR DEVELOPING HIGHER SEISMIC PERFORMANCE OF BRICK MASONRY WALLS

Koji Yoshimura1, Kikuchi Kikuchi1, Masayuki Kuroki2, Hideko Nonaka2, Kyong Tae Kim3, Reezang Wangdi3 and Ayasa Oshikata3

SUMMARY

To investigate the effective seismic strengthening methods for masonry walls in developing countries, a total of twenty eight unreinforced masonry (URM) and confined masonry (CM) walls were constructed and tested. The specimens include two-dimensional (2D) and three-dimensional (3D) masonry walls with and without wall reinforcing bars or U-shaped connecting bars, which were tested under constant gravity load and repeated lateral forces. The heights of inflection point considered were 0.67 and 1.1 times the height of the wall measured from the top of foundation beam. The constant vertical axial stresses applied were 0.48 and 0.84MPa. Test results obtained for each specimen include cracking patterns, load-deflection data, and strains in reinforcement and walls in critical locations. As observed from the obtained test results, it may be concluded that the CM walls with connecting bars at the vertical wall-to-column connections and the horizontal wall reinforcement are quite effective to develop the higher seismic performance than the ordinary URM walls.

BACKGROUND AND OBJECTIVE

In the event of strong motion earthquakes, a large number of severe structural damage has been observed in the unreinforced masonry (URM) wall buildings, which are widely constructed in the developing countries. Fig. 1 shows one of the typical examples of the extensive structural damage to URM walls caused by the 1999 Quindio earthquake in the Republic of Colombia, where the URM walls failed due to separation along their vertical wall-to-wall connections. The walls also failed in their own plane and turned

1 Professor, Dr. Eng., Department of Architectural Engineering, Oita University, Oita, 870-1192, Japan. 2 Research Associate, Department of Architectural Engineering, Oita University, Oita, 870-1192, Japan. 3 Graduate Student, Department of Architectural Engineering, Oita University, Oita, 870-1192, Japan.

Fig. 1 Damage to URM walls due to 1999 Colombia earthquake

over in their out-of-plane directions [1]. Similar structural damage to the URM walls was widely observed during the 1976 Tanshang earthquake in the People’s Republic (P.R.) of China, where more than 240,000 people were killed by this earthquake. In order to prevent this kind of wall separation damage during big earthquakes, confined masonry (CM) wall system was developed and has been constructed in many earthquake countries. Some of the CM walls used in the P. R. China and El Salvador are shown in Figs. 2 and 3. However, during the 1999 Colombia earthquake, extensive structural damage occurred in some of the newly constructed CM walls, in which adjacent URM walls were separated from the reinforced concrete (R/C) confining columns, as shown in Fig. 4. To prevent this kind of masonry wall-separation from the R/C confining columns, connecting bars recommended in National Standards of P.R. China [2] shown in Fig. 5 are provided between R/C columns and attached masonry walls in China. The main objective of the present study is to investigate the effective seismic strengthening methods for the masonry walls against severe earthquakes in the developing countries. Therefore, the wall specimens with different reinforcing details were designed, constructed and tested, as discussed hereafter in detail.

Fig. 2 Confined masonry walls under construction (Jimo, P.R. China, 1999)

Fig. 3 Confined masonry walls under construction (El Salvador, 2001)

Fig. 4 Damage to R/C column in CM wall due to 1999 Colombia

Fig. 5 Recommended details in masonry wall-to-R/C column connection in P.R. China

WALL SPECIMENS AND MATERIAL PROPERTIES

A total of twenty-eight unconfined and confined masonry wall specimens with different wall-to-wall connection details, listed in Table 1 and typical specimens shown in Fig.6, were designed and constructed. These specimens were tested under the repeated lateral forces, and constant axial stresses of 0.48MPa and 0.84MPa respectively. The thickness (t) of the masonry walls is 105mm, and the cross-section of the reinforced confining columns (bxD) in the extreme edges of the wall is 105mm x 105mm. However, the wall thickness and size of confining columns of the specimens with relatively higher compressive prism strength are t=100mm, b=100mm and D=100mm respectively. These specimens are classified into two test series, 2D and 3D specimens, depending on the shapes of the specimen as shown in Fig. 6(a) and 6(b), respectively. The effective length of intersecting (or flange) walls, constructed at the extreme edges of the main wall in the 3D-series, was determined based on the AIJ [3] recommendation. Each of the specimens is designated by the symbol code, such as 2D-L1-H42V0-48, 3D-L1-H42V0-48, 2D-H1-H0V0-48-CB and 3D-L1-H0V0-84-CB, 3D-L1-H0V0-84-CB-A etc. The first symbol “2D” or “3D” represents the two- or three-dimensional test specimens respectively. In the second symbol “L1” or “H1”, L or H indicates the location of the point of application

Table 1 List of test specimens and their details

Specimen ShapeInflecion

point

Longitudinal column

Re-bar

Horizontal wallRe-bar

Verticalwall Re-

bar

Axialstress

s0

Connectingbars

2D-L0-H0V0-48-1 Nil (0) 2D-L1-H0V0-48-1

2D-L1-H0V0-48-CB-1 D6@21 (CB)

2D-L1-H42V0-48-1 2-D6@42 (H42 Nil 3D-L0-H0V0-48-1 Nil (0) 3D-L1-H0V0-48-1

3D-L1-H0V0-48-CB-1 D6@21 (CB)

3D-L1-H42V0-48-1 2-D6@42 (H42 Nil 3D-L1-H0V0-48-CB-2 Nil (H0) D6@21 (CB)

3D-L1-H42V0-48-2 2-D6@42 (H42 Nil 2D-L0-H0V0-84-1 Nil (0) 2D-L1-H0V0-84-1

2D-L1-H0V0-84-CB-1 D6@21 (CB)

2D-L1-H42V0-84-1 2-D6@42 (H42 Nil 3D-L0-H0V0-84-1 Nil (0) 3D-L1-H0V0-84-1

3D-L1-H0V0-84-CB-1 D6@21 (CB)

3D-L1-H42V0-84-1 2-D6@42 (H42

3D-L1-H0V0-84-CB-2 Nil (H0) D6@21 (CB)

3D-L1-H42V0-84-2 2-D6@42 (H42 Nil 2D-H0-H0V0-48-2 Nil (0) 2D-H1-H0V0-48-2

2D-H1-H0V0-48-CB-2 D6@21 (CB)

2D-H1-H42V0-48-2 2-D6@42 (H42 Nil 2D-H0-H0V0-84-2 Nil (0) 2D-H1-H0V0-84-2

2D-H1-H0V0-84-CB-2 D6@21 (CB)

2D-H1-H42V0-84-2 2-D6@42 (H42 Nil

Nil (H0)

Nil

Nil0.84MPa(84)

0.48MPa(48)

0.84MPa(84)

Nil

Nil

0.48MPa(48)

Nil

Nil

Nil

1-D19 (1)

Nil (H0)

Nil (H0)

Nil (H0)

Nil (H0)

Nil (H0)

1-D19 (1)

1-D19 (1)

1-D19 (1)

1-D19 (1)

1-D19 (1)

1-D19 (1)

1-D19 (1)

2D

0.67

h 0 (h

0: w

all h

eigh

t mea

sure

d fr

om th

e to

p of

foun

datio

n be

am)

1.1h

0 (h

0: w

all h

eigh

tm

easu

red

from

the

top

offo

unda

tion

beam

)

2D

3D

2D

3D

Fig. 6(a) Typical 2D specimens Fig. 6(b) Typical 2D specimens

of lateral forces (or point of inflection) is “low” (taken as 0.67 times the wall height) or “high” (equal to 1.1 times the wall height) respectively and “1” represents that one longitudinal Re-bar with bar-size of D19 is provided in each of the confining R/C column section, with circular spiral hoops of D6 as shown in Fig. 6, and “0” means that no longitudinal Re-bar is provided, in other words, the masonry wall is not confined. The third symbol “H42” indicates that the horizontal Re-bars are provided at the spacing of 42cm, and “H0” means no horizontal Re-bars are used. The fourth symbol “V0” indicates that no vertical Re-bars are provided in the masonry wall. The numerals “48” or “84” represent that the constant vertical axial stress of 0.48MPa or 0.84MPa are applied to the specimens. The sixth symbol “CB” indicates that U-shaped connecting bars with D6 (or #2) were placed at every 21cm between the masonry wall-edges and the attached R/C column sections. The last numeral “1” represents the test specimens with the medium compressive prism strength (Fm=18~27MPa) and “2” represents those specimens having comparatively the higher compressive strength (Fm=38~45MPa). The compressive strengths and mechanical properties of the materials used for the specimens are shown in Table 2.

TEST SETUP, INSTRUMENTATION AND

TEST PROCEDURE

The test setup adopted in the present study is shown in Fig. 7. Test setup consisted of steel testing frame and two hydraulic actuators, fixed to the frame in order to simulate constant vertical and in plane lateral repeated loads. The vertical load was applied to the specimens by a hydraulic jack with 2,000kN capacity, connected to Servo-impulse controller system in order to keep

Table 2 Materials and prism strengths and mechanical properties of Re-bars

Fig. 7 Test setup

2D-L0-H0V0-48-1 Nil 29.0 27.0 28.3

2D-L1-H0V0-48-1 26.4 25.5 20.6 36.0

2D-L1-H0V0-48-CB-1 24.8 25.7 24.8 33.0

2D-L1-H42V0-48-1 28.2 29.7 21.6 26.6

2D-H0-H0V0-48-2 Nil 27.8 59.6 51.7

2D-H1-H0V0-48-2 36.7 26.3 53.6 63.2

2D-H1-H0V0-48-CB-2 35.0 26.3 51.6 51.7

2D-H1-H42V0-48-2 29.4 19.6 52.9 51.1

3D-L0-H0V0-48-1 Nil 28.8 24.7 34.7

3D-L1-H0V0-48-1 26.4 24.2 22.0 33.6

3D-L1-H0V0-48-CB-1 27.0 30.7 22.6 24.6 448 565 10 341 483 24

3D-L1-H42V0-48-1 28.2 30.8 19.9 33.4

3D-L1-H0V0-48-CB-2 32.5 22.8 45.2 42.6

3D-L1-H42V0-48-2 28.1 21.4 37.6 32.4

2D-L0-H0V0-84-1 Nil 29.8 20.8 22.7

2D-L1-H0V0-84-1 30.5 28.6 18.9 25.7

2D-L1-H0V0-84-CB-1 28.8 28.6 19.4 24.6 448 565 10 341 483 24

2D-L1-H42V0-84-1 27.4 29.4 18.5 21.5

2D-H0-H0V0-84-2 Nil 19.6 45.3 53.6

2D-H1-H0V0-84-2 35.5 27.8 60.6 40.3

2D-H1-H0V0-84-CB-2 32.3 29.4 51.7 52.2

2D-H1-H42V0-84-2 31.4 29.4 56.6 51.5

3D-L0-H0V0-84-1 Nil 30.7 22.1 26.9

3D-L1-H0V0-84-1 28.5 28.2 23.9 34.7

3D-L1-H0V0-CB-84-1 27.2 25.7 23.4 32.7

3D-L1-H42V0-84-1 28.8 29.4 21.2 28.8

3D-L1-H0V0-84-CB-2 31.2 23.0 43.3 43.0

3D-L1-H42V0-84-2 28.2 25.6 38.6 28.6

511 19 341 483

Nil

334

Nil

Nil Nil

JointMortar(MPa)

Prism(MPa)

489 25387 505

NilNil

Properties of Steel Re-bars

Tensilestrength(MPa)

Axialstress

σ0

(MPa)

Specimen

24 334

Column(MPa)

Concrete

Beam(MPa)

D6 D19

Elongation(%)

387 505

387

Nil

24

24

Nil

24

Nil

330

0.84

0.48

Nil

334

Nil

387

489

505

25505

489 25

334 489 25

24

Yieldstrength(MPa)

Tensilestrength(MPa)

Elongation(%)

Yieldstrength(MPa)

Nil

330 511 19 341 483 24

constant vertical load during the test. The lateral load was applied to the specimen by a double-acting hydraulic jack with 2,000kN capacity, placed laterally and fixed to the testing frame and reaction wall. The heights of the longitudinal axis of the lateral forces applied to the specimens (or height of inflection point) were approximately 0.67 and 1.1 times the wall height measured from the foundation beam. The test specimens were instrumented to monitor the applied loads, displacements and the resulting strains induced in wall and reinforcing bars. The displacement transducers attached at the center of wall top were used to monitor the top surface of the in-plane lateral displacements, etc. The transducers were also installed at the critical locations to measure other parameters such as slip, out-of-plane displacement, etc. All the measuring instruments and load cells were connected to a computerized control system and data acquisition system. The wall specimens were tested under constant gravity load and repeated lateral forces. The axial compression load was applied through the loading beam using vertical hydraulic jack. The repeated lateral forces were applied in a direction along the plane of the main wall. Two types of constant vertical loads corresponding to the axial stresses of 0.48MPa and 0.84MPa were considered. These axial stresses were determined based on the values of the vertical loads, which are supported by the first-story masonry walls in some typical medium-rise residential buildings in Japan. Each test was first conducted under lateral load control, and then changed to the lateral displacement control as the specimen became more flexible during testing. A constant rate was applied and the data were recorded at the specific regular increments. The cracks were marked as they occurred. The test was ended when the load-deflection curve monitored on computer screen or/and on the X-Y recorder showed a continuous drop in the load to about 80% of the peak lateral load as the lateral displacement increased. The load and all the instrument measurements were continuously and automatically scanned and recorded on hard disk and then analyzed in personal computer.

TEST RESULTS AND DISCUSSIONS

Crack patterns and modes of failure The crack and crack propagation during the tests were monitored and recorded by marking the cracks at the end of the half cycle of loading while the specimen was held at the maximum displacement, although crack widths were not filed. The cracks were partially closed with load reversal. The final crack patterns developed in the selected specimens are shown in Figs. 8 and 9. Almost all the CM specimens in L-series (whose infliction point is 0.67 times wall height) failed in shear mode and also sliding was recorded in some of the specimens. However, the clear separation of wall from the R/C confining column was not seen in unreinforced CM specimens although the cracks were developed along the vertical joint. On the other hand, unreinforced CM specimens in H-series (whose inflection point is 1.1 times wall height) showed a distinct separation of wall from the R/C confining column as shown in Fig. 9(b). This type of wall separation was not seen in CM specimens provided with horizontal wall reinforcement and U-shaped

c) 2D-L1-H0V0-84-CB-1

Fig. 8 Final crack patterns of 2D-L-84 series

a) 2D-L0-H0V0-84- b) 2D-L1-H0V0-84-1 d) 2D-L1-H42V0-84-1

connecting bars, though few cracks were developed. All H-series specimens failed in flexure mode at first and ultimately in either sliding or shear failure mode in some cases. Further, the unreinfoirced specimens failed in flexure, i.e. cracks developed in horizontal bed joint at the bottom course and also in other courses on the tension side

Q-R hysteresis loops The hysteresis loops of the applied lateral load (Q) versus story drift (R) response curves for constant vertical axial stress of 0.84MPa only are shown in Figs.10, 11 and 12. The story drift (R) is defined as a story displacement between top and bottom of the wall divided by the wall height (h) of the specimen measured from top of the foundation beam, and the mean shearing stress (τ ), which is obtained by dividing the applied lateral load by the gross horizontal cross-sectional area of the main wall (and main plus flange walls in case of 3D specimens), are also shown in these figures. In the Q-R curves, crack and strain information are also presented by using different symbols as shown in the figures. The dashed lines

-400

-300

-200

-100

0

100

200

300

400 Q(kN) (MPa)τ

-1.5 -1.0 -0.5 0 0.5 1.0 1.5 R(x10-2rad)

0.5

0

-0.5

-1.0

-1.5

-2.0

1.0

1.5

2.0

[F SL]

-400

-300

-200

-100

0

100

200

300

400 Q(kN) (MPa)τ

-1.5 -1.0 -0.5 0 0.5 1.0 1.5 R(x10-2rad)

0.5

0

-0.5

-1.0

-1.5

-2.0

1.0

1.5

2.0

[S]

-400

-300

-200

-100

0

100

200

300

400 Q(kN) (MPa)τ

-1.5 -1.0 -0.5 0 0.5 1.0 1.5 R(x10-2rad)

0.5

0

-0.5

-1.0

-1.5

-2.0

1.0

1.5

2.0

[S]

-400

-300

-200

-100

0

100

200

300

400 Q(kN) (MPa)τ

-1.5 -1.0 -0.5 0 0.5 1.0 1.5 R(x10-2rad)

0.5

0

-0.5

-1.0

-1.5

-2.0

1.0

1.5

2.0

[S]

a) 2D-L0-H0V0-84-1 b) 2D-L1-H0V0-84-1 c) 2D-L1-H0V0-84-CB-1 d) 2D-L1-H42V0-84-1

Fig. 10 Q-R hysteresis loops of 2D-L-84 series

Predicted ultimate flexural strength Predicted ultimate shear strength Initial flexural crack Initial shear crack Initial yield in South column bar Initial yield in North column bar Initial yield in U-shaped connecting bar or horizontal wall Re-bars

View 3 d) 2D-H1-H42V0-84-2

Fig. 9 Final crack patterns of 2D-H-84 series

a) 2D-H0-H0V0-84-2 b) 2D-H1-H0V0-84-2

c) 2D-H1-H0V0-84-CB-2

View 1 View 2

1 2

3 4

View 4

in the figures represent the theoretical values determined by the ultimate flexural capacity at the bottom of each wall (Qmu), while the dotted lines represent the ultimate lateral strengths determined in shear failure mode of the wall with flexural reinforcement in its wall edges or R/C confining columns (Qsu), and dashed-dot lines in Fig 12 represent the initial flexural crack strength (Qmc).

Load-displacement envelope curves Envelope curves of lateral load in terms of shearing stress, ( τ )- versus- story drift obtained from the Q-R hysteresis loops of the specimens are presented in Figs.13 (a) thorough (f). These curves represent the average of positive and negative loadings. In all curves shown, the plotted drift is on the basis of horizontal top displacement without any correction for slip (for the specimens which exhibited a sliding failure). The failure modes are also indicated by different symbols as shown in the figures. The shape of

-400

-300

-200

-100

0

100

200

300

400 Q(kN) (MPa)τ

-1.5 -1.0 -0.5 0 0.5 1.0 1.5 R(x10-2rad)

0.5

0

-0.5

-1.0

-1.5

-2.0

1.0

1.5

2.0

[F S]

-400

-300

-200

-100

0

100

200

300

400 Q(kN) (MPa)τ

-1.5 -1.0 -0.5 0 0.5 1.0 1.5 R(x10-2rad)

0.5

0

-0.5

-1.0

-1.5

-2.0

1.0

1.5

2.0

[S]-400

-300

-200

-100

0

100

200

300

400 Q(kN) (MPa)τ

-1.5 -1.0 -0.5 0 0.5 1.0 1.5 R(x10-2rad)

0.5

0

-0.5

-1.0

-1.5

-2.0

1.0

1.5

2.0

[S]

-400

-300

-200

-100

0

100

200

300

400 Q(kN) (MPa)τ

-1.5 -1.0 -0.5 0 0.5 1.0 1.5 R(x10-2rad)

0.5

0

-0.5

-1.0

-1.5

-2.0

1.0

1.5

2.0

[S]

a) 3D-L0-H0V0-84-1 b) 3D-L1-H0V0-84-1 c) 3D-L1-H0V0-84-CB-1 d) 3D-L1-H42V0-84-1

Fig. 11 Q-R hysteresis loops of 3D-L-84 series

Predicted ultimate flexural strength Predicted ultimate shear strength Initial flexural crack Initial shear crack Initial yield in South column bar Initial yield in North column bar Initial yield in U-shaped connecting bar or horizontal wall Re-bars

-400

-300

-200

-100

0

100

200

300

400

0

-0.5

-1.5

-2.0

2.0

1.5

1.0

0.5

-1.0

-2.0 -1.5 -1.0 -0.5 0 0.5 1.0 1.5 2.0

Q(kN) (MPa)τ

R(x10-2rad)

[F]

[F]

-400

-300

-200

-100

0

100

200

300

400

-2.0 -1.5 -1.0 -0.5 0 0.5 1.0 1.5 2.0

Q(kN) (MPa)τ

R(x10-2rad)

0

-0.5

-1.5

-2.0

2.0

1.5

1.0

0.5

-1.0

[F]

[S]

-400

-300

-200

-100

0

100

200

300

400

-2.0 -1.5 -1.0 -0.5 0 0.5 1.0 1.5 2.0

Q(kN) (MPa)τ

R(x10-2rad)

0

-0.5

-1.5

-2.0

2.0

1.5

1.0

0.5

-1.0

[F]

[F]

-400

-300

-200

-100

0

100

200

300

400

-2.0 -1.5 -1.0 -0.5 0 0.5 1.0 1.5 2.0

Q(kN) (MPa)τ

R(x10-2rad)

0

-0.5

-1.5

-2.0

2.0

1.5

1.0

0.5

-1.0

[F]

[F]

Predicted ultimate flexural strength (Qmu) Predicted ultimate shear strength (Qsu) Predicted initial flexural crack strength (Qmc) Initial flexural crack Initial shear crack Initial yield in South column bar Initial yield in North column bar Initial yield in U-shaped connecting bar or horizontal wall Re-bars

a) 2D-H0-H0V0-84-2 b) 2D-H1-H0V0-84-2 c) 2D-H1-H0V0-84-CB-2

d) 2D-H1-H42V0-84-2

Fig. 12 Q-R hysteresis loops of 2D-H-84 series

the envelope curves beyond the maximum ultimate load varies from one wall specimen to another specimen depending upon the level of axial stress and whether they are confined or unconfined. As can be understood from the envelope curves, the CM specimens with horizontal wall reinforcement and U-shaped connecting bars showed a better performance as compared to the unreinforced CM and URM specimens, in terms of lateral load carrying capacity and ductility.

0.0

0.5

1.0

1.5

2.0

: R(x 10-2 rad)Story drift

: (

MP

a)τ

Av.

she

arin

g st

ress

0 0.5 1.0 1.5

2D-H1-H0V0-84-2

2D-H1-H0V0 -84-CB-2

2D-H1-H42V0-84-2

2D-H0-H0V0-84-2

Flexure failureShear failureSliding failure

f) 2D-H-84 series

Flexure failure

Shear failureSliding failure

0.0

0.5

1.0

1.5

2.0

: R(x 10-2 rad)Story drift

: (

MP

a)τ

Av.

she

arin

g st

ress

0 0.5 1.0 1.5

2D-H1-H0V0-48-22D-H1-H0V0-48-CB-2

2D-H1-H42V0-48-2

2D-H0-H0V0-48-2

e) 2D-H-48 series

0.0

0.5

1.0

1.5

2.0

2D-L1-H0V0-48-1

2D-L1-H0V0-48-CB-12D-L1-H42V0-48-1

2D-L0-H0V0-48-1

: R(x 10-2 rad)Story drift

: (

MP

a)τ

Av.

she

arin

g st

ress

0 0.5 1.0 1.5

Flexure failureShear failureSliding failure

a) 2D-L-48 series

0.0

0.5

1.0

1.5

2.0

2D-L1-H0V0-84-12D-L1-H0V0-84-CB-1

2D-L1-H42V0-84-1

2D-L0-H0V0-84-1

: R(x 10-2 rad)Story drift

: (

MP

a)τ

Av.

she

arin

g st

ress

0 0.5 1.0 1.5

Flexure failureShear failureSliding failure

b) 2D-L-84 series

0.0

0.5

1.0

1.5

2.0

3D-L1-H0V0-48-1

3D-L1-H42V0-48-23D-L1-H0V0-48-CB-1

3D-L0-H0V0-48-1

3D-L1-H42V0-48-1

3D-L1-H0V0-48-CB-2

: R(x 10-2 rad)Story drift

: (

MP

a)τ

Av.

she

arin

g st

ress

0 0.5 1.0 1.5

Flexure failureShear failure

Sliding failure

c) 3D-L-48 series

0.0

0.5

1.0

1.5

2.0

3D-L1-H0V0-84-CB-1

3D-L1-H0V0-84-1

3D-L1-H42V0-84-23D-L1-H0V0-84-CB-2

3D-L1-H42V0-84-1

3D-L0-H0V0-84-1

: R(x 10-2 rad)Story drift

: (

MP

a)τ

Av.

she

arin

g st

ress

0 0.5 1.0 1.5

Flexure failure

Shear failureSliding failure

d) 3D-L-84 series

Fig. 13 τ -R envelope curves

Data on the predicted and observed ultimate strengths and failure modes of all the specimens, and related data are summarized in Table 3. Letters F, SL and S represent the failure modes in flexure, sliding and shear respectively. The values of ductility factors given in the table were calculated from the proposition of Sheikh et al [4]. Although the procedure suggested was to evaluate the ductility factor of R/C columns, this concept was applied to evaluate the yield- and ultimate-story drift for the present specimens. In the idealized τ -R envelope curve shown in Fig. 14, the yield drift, R1, is defined as the story drift corresponding to the intersection point of initial stiffness line and the maximum shearing stress, and the ultimate story drift, R2, is defined as the maximum story drift corresponding to about 80% of the maximum shearing stress on the τ -R curves.

Table 3 Predicted and observed flexural, shear and ultimate strengths, failure modes and ductility factors

K e R2④／①　　⑤／②

(=Q/R) R1 ⑥／③　⑧／⑦ V mc V sc V mu V su

(×102) (×10-2) (×10-2) (kN) (kN) (kN) (kN)(MN/rad) (rad) (rad)

2D-L0-H0V0-48-1 ⓐ 3.9 70 - 77 (0.42) SL 0.02 >0.80 　>40 60 215 83 118 F

2D-L1-H0V0-48-1 ⓑ 7.2 127 165 166 (0.90) 0.02 0.17 　　7 98 244 211

2D-L1-H0V0-48-CB-1 ⓒ 5.6 118 144 162 (0.88) 0.03 >0.90 >30 95 266 230

2D-L1-H42V0-48-1 ⓓ 4.6 133 135 183 (1.00) 0.04 0.89 22 89 250 288 F

3D-L0-H0V0-48-1 ⓐ 5.6 177 148 183 (1.00) F→SL 0.03 0.08 2 1.44 - 2.36 198 206 144 129

3D-L1-H0V0-48-1 ⓑ 5.0 - 139 149 (0.81) 0.03 0.28 9 0.69 0.84 0.90 240 252 218

3D-L1-H0V0-48-CB-1 ⓒ 9.2 167 186 219 (1.19) 0.02 >0.93 >47 1.64 1.29 1.35 219 255 221

3D-L1-H42V0-48-1 ⓓ 6.8 181 192 214 (1.16) 0.03 0.40 13 1.48 1.42 1.17 240 240 289 F (S)

3D-L1-H0V0-48-CB-2 ⓒ 9.7 203 195 234 (1.34) 0.03 0.30 12 1.73 1.35 1.44 244 335 296

3D-L1-H42V0-48-2 ⓓ 8.8 232 190 234 (1.34) 0.03 0.26 9 1.91 1.41 1.28 225 307 372

2D-L0-H0V0-84-1 ⓐ 3.6 122 124 124 (0.67) F→SL 0.04 0.51 15 1.60 74 208 140 116

2D-L1-H0V0-84-1 ⓑ 3.9 126 149 149 (0.81) 0.04 0.33 9 0.90 117 252 214

2D-L1-H0V0-84-CB-1 ⓒ 5.1 139 164 166 (0.90) 0.03 >0.67 >21 1.02 116 255 217

2D-L1-H42V0-84-1 ⓓ 5.9 158 158 190 (1.03) 0.03 0.60 19 1.04 113 249 290 F (S)

3D-L0-H0V0-84-1 ⓐ 7.0 201 188 224 (1.22) F→SL 0.03 0.25 8 1.94 1.52 1.81 1.23 241 213 244 125

3D-L1-H0V0-84-1 ⓑ 5.6 215 202 230 (1.25) 0.04 0.40 10 1.44 1.36 1.54 1.54 314 278 238

3D-L1-H0V0-84-CB-1 ⓒ 6.4 - 181 213 (1.16) 0.03 0.70 21 1.25 1.10 1.28 0.97 309 276 235

3D-L1-H42V0-84-1 ⓓ 7.7 - 204 228 (1.24) 0.03 >0.93 >31 1.31 1.29 1.20 1.07 301 264 297

3D-L1-H0V0-84-CB-2 ⓒ 9.1 294 298 329 (1.88) 0.04 0.43 11 1.78 1.82 1.98 1.41 313 344 301 S

3D-L1-H42V0-84-2 ⓓ 8.6 - 241 273 (1.56) 0.03 0.44 13 1.46 1.53 1.44 1.17 284 327 387 F

2D-H0-H0V0-48-2 ⓐ 4.1 39 - 48 (0.27) 0.01 >1.35 >115 43 313 48 146

2D-H1-H0V0-48-2 ⓑ 5.2 82 98 142 (0.81) 0.03 0.27 10 72 298 236

2D-H1-H0V0-48-CB-2 ⓒ 4.5 39 100 149 (0.85) 0.03 0.29 9 70 293 232

2D-H1-H42V0-48-2 ⓓ 4.4 89 110 147 (0.84) F, F→SL 0.03 0.46 14 67 297 354

2D-H0-H0V0-84-2 ⓐ 3.7 60 - 78 (0.45) F 0.02 >0.99 >47 1.63 53 261 81 150

2D-H1-H0V0-84-2 ⓑ 3.7 99 137 166 (0.95) F, S 0.04 0.46 10 1.17 88 332 261

2D-H1-H0V0-84-CB-2 ⓒ 4.1 106 140 175 (1.00) F 0.04 0.43 10 1.17 86 310 243

2D-H1-H42V0-84-2 ⓓ 4.7 89 140 175 (1.00) F, F→SL 0.04 0.76 21 1.19 81 322 377

①

②

③

Effect of cross-section(3D/2D)

②／①　　③／①⑤／④　　⑥／④

Initialstiffness

Shearcrack

strength

Ultimatestrength

SpecimensR1 R2

Story drift and ductilityfactor

Shearcrack

strength

Ultim atestrength

Flexuralcrack

strength

⑧

F

Q sc

(kN)

Q mc

(kN)

Failuremode

Q max (kN)

⑤

⑥ S

Ratio of Experimental valuesExperimental value

Initialstiffness

⑦

S

S

S

S

S④

Theoretical value

366

140

172

232

286

277

283

379

Failuremode

S

Ultimatestrength

Flexuralcrack

strength

Shearcrack

strength

Ultimateflexuralstrength

Ultimateshear

strength

Effect of verticalaxial load

(0.84/0.48)

F

S

F

S

S

F

⎟⎟⎟

⎠

⎞

⎜⎜⎜

⎝

⎛ =

(MPa)w

maxmax A

Qτ

Fig. 14 Idealized τ–R envelope curve

DISCUSSIONS

Effect of cross-section The difference in seismic behavior of 2D and 3D confined wall specimens (as indicated by ratio 3D/2D in Table 3) had been observed in each case of the constant vertical axial stress (σ0) of 0.48MPa and 0.84MPa acting on these specimens. It may be argued that the presence of intersecting (or flange) walls in 3D wall specimens might have attributed to the difference in their action against the lateral loads. The effect of the cross-sectional difference is further illustrated graphically in Figs. 15 (a), (b) and (c) with respect to initial stiffness (Ke), shear crack strength (Qsc) and maximum ultimate lateral load (Qmax) respectively. As can be seen from these figures, all the 3D specimens having flange walls, except for 3D-L1-H0V0-48, showed better results in terms of these parameters than 2D specimens without the intersecting walls. For the axial stress (σ0) of 0.48MPa and 0.84MPa, the difference in seismic behavior of 3D over 2D specimens expressed in terms of initial stiffness (Ke) is between 1.44~1.91 and 1.25~1.94 times, shear crack strength (Qsc) from 1.29~1.42 and 1.10~1.82 times, and maximum ultimate lateral load (Qmax) from 1.17~2.36 and 1.2~1.98. Particularly for the four 3D specimens with higher prism strength (3D-2) when compared to 2D specimen, the difference in Qmax was observed to be between 1.28~1.44 and 1.44~1.98 times.

Effect of axial load Depending upon the magnitude of constant axial stress (σ0) applied to the specimen, the ultimate lateral load resisted also varies accordingly which is shown in Fig. 16. For those specimens that failed in flexure mode indicated as “F” on the graph, the ultimate lateral strengths (Qmax) corresponding to the vertical axial stress of 0.84MPa presented by dark bars are higher by about 1.17~1.63 times of that carried by the

Fig. 15 (c) Effect of cross-sectional shape on ultimate lateral strength

Fig. 15 (b) Effect of cross-sectional shape on crack strength

0

50

100

150

200

250

300

350

400

L0-

H0V

0-48

-1

L1-

H0V

0-48

-1

L1-

H0V

0-48

-CB

-1

L1-

H42

V0-

48-1

L1-

H0V

0-48

-CB

-2

L1-

H42

V0-

48-2

L0-

H0V

0-84

-1

L1-

H0V

0-84

-1

L1-

H0V

0-84

-CB

-1

L1-

H42

V0-

84-1

L1-

H0V

0-84

-CB

-2

L1-

H42

V0-

84-2

Qmax

(kN)

Sp

ecim

en2D

spe

cim

en3D

spe

cim

en

2D s

peci

men

3D s

peci

men

0

50

100

150

200

250

300

350

400

L0-

H0V

0-48

-1

L1-

H0V

0-48

-1

L1-

H0V

0-48

-CB

-1

L1-

H42

V0-

48-1

L1-

H0V

0-48

-CB

-2

L1-

H42

V0-

48-2

L0-

H0V

0-84

-1

L1-

H0V

0-84

-1

L1-

H0V

0-84

-CB

-1

L1-

H42

V0-

84-1

L1-

H0V

0-84

-CB

-2

L1-

H42

V0-

84-2

Qsc

(kN)

Sp

ecim

en

Fig. 15 (a) Effect of cross-sectional shape on initial stiffness

2D s

pec

imen

s3D

spe

cim

ens

0

2

4

6

8

10

L0-

H0V

0-48

-1

L1-

H0V

0-48

-1

L1-

H0V

0-48

-CB

-1

L1-

H42

V0-

48-1

L1-

H0V

0-48

-CB

-2

L1-

H42

V0-

48-2

L0-

H0V

0-84

-1

L1-

H0V

0-84

-1

L1-

H0V

0-84

-CB

-1

L1-

H42

V0-

84-1

L1-

H0V

0-84

-CB

-2

L1-

H42

V0-

84-2

Ke(x102MN/rad)

Sp

ecim

en

specimens tested under lower vertical axial stress of 0.48MPa shown by white bars. From this result, it is understood that the specimens, which failed in flexure exhibited different lateral load carrying capacity depending upon the magnitude of an applied constant vertical axial load. Letter “S” on the graph in Fig. 16 indicates the specimens failed in shear mode. In series 2 specimens (with higher material strength) the ultimate lateral strengths (Qmax) borne by the specimens subjected to constant axial stress (σ0) of 0.84MPa, shown by dark bars, have increased by about 1.41 and 1.17 times of those specimens when the applied vertical axial stress was 0.48MPa, as shown by white bars. Here also, it is seen that with the increasing values of the vertical axial load, the lateral load carrying capacities of the specimen were also increased. However, among the series 1 specimens (with medium material strength) the ultimate lateral strengths were almost the same and in few cases the results were reversed. Effect of horizontal wall reinforcement and connecting bars The reinforcing steel i.e. horizontal wall reinforcement and U-shaped connecting bars used in the specimens mentioned in this report, can be seen as one of the critical parameters affecting wall ductility. However, it was impossible to evaluate the ductility factor of the tested wall specimens correctly because these specimens did not show the perfect elasto-plastic behavior. Therefore, the procedure explained in the forgoing section under load-displacement envelope curves has been adopted to calculate the ductility factor, which is given by the ratio R2/R1 whose values are presented in Table 3 above as well as in the chart shown in Figs. 17 (a) and (b). Fig.17 (a) shows the ductility factors of the specimens failed in shear mode and Fig. 17(b) shows the ductility factors of the specimens failed in flexure mode. In Fig.17 (a) the ductility factors (R2/R1) of the unreinforced CM specimens (L1-H0V0-1) represented by white bars were within 7~10 while the ductility factors (R2/R1) of the CM specimens provided with horizontal wall reinforcement and U-shaped connecting bars (L1-CB-1) and (L1-H42-1) represented by dark bars were ranging from 21~47 and 13~31 respectively. As indicated by these ductility factors as well as by the load-displacement curves in Fig. 13 above, it may be concluded that horizontal wall reinforcement and U-shaped connecting bars improve the deformation behavior of wall after attaining the maximum ultimate lateral load. However, the ductility factors of series 2

0

50

100

150

200

250

300

350

400

2D-L

0-H

0V0-

1

2D-L

1-H

0V0-

1

2D-L

1-H

0V0-

CB

-1

2D-L

1-H

42V

0-1

2D-H

0-H

0V0-

2

2D-H

1-H

0V0-

2

2D-H

1-H

0V0-

CB

-2

2D-H

1-H

42V

0-2

3D-L

0-H

0V0-

1

3D-L

1-H

0V0-

1

3D-L

1-H

0V0-

CB

-1

3D-L

1-H

42V

0-1

3D-L

1-H

0V0-

CB

-2

3D-L

1-H

42V

0-2

Qmax

(kN)

F

S SS

SS S

S

S

F

F F F

F

S: Shear failure

F: Flexure failure : =0.84MPa

: =0.48MPa

σ0

σ0

Sp

ecim

en

Fig. 16 Effect of vertical axial load on ultimate lateral strength

Fig. 17 Effect of horizontal wall reinforcement and U-shaped connecting bars on ductility factor

Ductility factor (R2/R1)

Ductility factor (R2/R1)

(a) (b) S

pec

imen

Sp

ecim

en

specimens with higher material strength, (3D-L1-2) were around 9~13 shown by shaded bars and these indicate that deformation capacities are lower as compared to series 1 specimens (L1-CB-1) and (L1-H42-1). The ductility factors of the specimens, which failed in flexure mode, are shown in Fig.17 (b). The values of R2/R1 for the unreinforced CM specimens (H1-H0V0) and CM specimens with connecting bars (H1-CB) were around 9~10 while CM specimens with horizontal wall reinforcement were around 14~21 and this implies that horizontal steel reinforcement contributed to the ductile behavior of the wall specimens. However, as can be seen from the envelope curves in Figs.13 (e) and (f) above, the CM specimens with connecting bars (H1-CB) showed more or less stable deformability after the story drift of about 0.5x10-2rad. Hence, it may be concluded that the CM specimens with horizontal wall reinforcement and connecting bars exhibited ductile behavior after yielding of the reinforcements. The ductility of URM specimens could not be calculated, as it was not possible to determine the ultimate drift indicated as >40, 30, etc. in Table 3 above. The reason was that the strength of specimens did not fall to or below 80% after attaining the maximum ultimate strength. Deformation capacity Load-displacement characteristics of the specimens were discussed through envelope curves in the previous section. The maximum strength was reached at a drift of approximately 0.20%. Since ductility factors discussed above may not be a representative measure of the inelastic behavior in this case, a better measure of the deformation capacity may be considered to be the ultimate drift, R2. The ultimate drift values of the reinforced CM specimens ranged between 0.26~0.76% approximately and that of the unreinforced CM specimens were in around 0.17~0.46%. These values are given as R2 in Table 3. It may be concluded that the deformation capacity of the wall specimens with horizontal and connecting bars are better than unreinforced CM specimens. As can be seen from the τ -R envelope curves, the URM specimens, after developing their maximum strengths, remained constant for successive repeated loading and the strength drop, with the exception of 3D specimens, was not observed particularly in the specimens that failed in flexure mode. Evaluation of theoretical values using existing equations The predicted theoretical values, given in Table 3, as flexural crack strength (Vmc), shear crack strength (Vsc), ultimate flexural strength (Vmu), and ultimate shear strength (Vsu) for all the masonry wall specimens were determined by the existing equations discussed below. Crack strengths The flexural crack strengths of the URM specimens (Vmc1), were calculated from the following equation recommended in the National Standards of P.R. China [5].

( )( )h

fZV tmmc ′

+⋅= 01

σ (1)

where Vmc1: flexural crack strength (N), Z: section modulus (mm3), mft: tensile strength of masonry which is given by

ztm Ff 125.0= (MPa), where Fz: compressive strength of joint mortar (MPa), and h′ : height

of inflection point (mm). The flexural crack strengths of the CM specimens (Vmc2), were calculated based on the cracking moment obtained by assuming the horizontal cross-section of the wall to remain plane after bending and also neglecting the effect of reinforcing bars in the confining columns. The strain in concrete on the tension side was calculated using the relation,

ctcc Ef=ε , where cε : strain in concrete, Ec: modulus of elasticity

of concrete taken as 21000MPa, cft: tensile strength of concrete given by Btc f σ56.0= (MPa) in AIJ

Standards [6], where Bσ : compressive strength of concrete (MPa). The strains in required points on the

horizontal cross-section were obtained from the strain distribution diagram. The strain in masonry calculated from using the property of strain diagram was checked and found to be within the values obtained from the expression,

mtmm Ef=ε , where εm: strain in masonry, Em: modulus of elasticity of

masonry taken as 5000MPa, mft: tensile strength of masonry, given byztm Ff 125.0= (MPa), where Fz:

compressive strength of joint mortar (MPa). Finally, using the stress diagram, the cracking moment (Mc) was calculated and the corresponding crack strength (Vmc2) was obtained by dividing it by the height of inflection point ( h′ ), i.e. hMV cmc

′=2.

The theoretical shear crack strengths of the CM wall specimens were calculated using the equation proposed by Matsumura [7] as given below.

( )3

0 103.02

1 ⋅⋅⋅⎭⎬⎫

⎩⎨⎧

⋅⋅++

= jtFdh

kV mcsc σα (2)

where Vsc: shear crack strength (kN), kc: reduction factor for partially grouted masonry, adopted as 1.0, h: height of the masonry wall (m), d: distance between the compression fiber and the tension bar in the confining columns (m), Fm: compressive strength of the masonry prism (MPa), α: concentration stress, adopted as 1.0, σ0: constant vertical axial stress (MPa), t: thickness of the masonry wall, and j: distance between the forces of compression and tension (m), taken as equal to 7/8d. Ultimate strengths The ultimate shear strengths of the URM specimens (Vmu1) after the development of flexural cracks were calculated by the following equation.

( )h

lwNVmu ′

⋅+=2

01

(3)

where Vmu1: ultimate lateral shear strength corresponding to the ultimate flexural moment (N), N: vertical axial load (N), w: self-weight of wall plus collar beam (N), l0: length of wall (mm) and h′ : height of inflection point (mm). The ultimate shear strengths of the CM specimens (Vmu2) corresponding to the ultimate flexural moment were calculated from the following equation recommended by AIJ Standards [8].

( ) hlNlalaV wwwywwytmu′⋅+⋅⋅+⋅= ⋅ 5.05.02 σσ (4)

where Vmu2: ultimate lateral shear strength corresponding to the ultimate flexural moment (N), at: cross-sectional area of longitudinal Re-bar in confining column (mm2), σy: yield strength of longitudinal Re-bar in confining column (MPa), lw: center to center distance between of longitudinal Re-bar in confining column (mm), aw: cross-sectional area of vertical wall reinforcing bars (mm2) and is equal to zero in the present test specimens, σwy: yield strength of vertical wall reinforcing bars (MPa), N: vertical axial load acting on the masonry wall (Newton), and h′ : height of inflection point (mm). The ultimate shear strengths of URM masonry wall specimens (Vsu1) were determined from the following equation recommended in the Chinese Standards [9]

w

v

vsu Af

fV ⋅⎟⎟⎠

⎞⎜⎜⎝

⎛+= 0

1 45.012.1

1 σ (5)

where Vsu1: ultimate lateral shear strength (N), fv: ultimate shear stress with axial load (MPa), where

zv Ff 125.0= (MPa), Fz: compressive strength of joint mortar (MPa), σ0: constant vertical axial stress

(MPa), and Aw: horizontal cross-sectional area of masonry wall (mm2). The ultimate shear strengths of CM masonry wall specimens (Vsu2) were calculated from the following equation recommended by Matsumura [7], which is actually for reinforced hollow concrete masonry wall.

3

02 102.018.0012.07.0

76.0 ⋅⋅⋅

⎪⎪⎭

⎪⎪⎬

⎫

⎪⎪⎩

⎪⎪⎨

⎧

+⋅⋅⋅+

⎥⎥⎥⎥

⎦

⎤

⎢⎢⎢⎢

⎣

⎡

+⎟⎠

⎞⎜⎝

⎛ +⋅= jtFpF

d

hkkV myhhmpusu σσδγ

(6)

where Vsu2: ultimate lateral shear strength (kN), ku: reduction factor equal to 1.0, kp: 1.16pt

0.3 (pt=at/(t.d) in %), h: height of the masonry wall (m), d: distance between the compression fiber and the tension bar in the confining columns (m), Fm: compressive strength of prism (MPa), γ: factor corresponding to the confine grout equal to 0.8, δ: factor concerning loading method equal to 1.0, ph: horizontal steel reinforcement ratio, hσy: yield strength of horizontal reinforcing steel bar (MPa), σ0: vertical axial stress (MPa), t: thickness of the masonry wall (m), and j: distance between the forces of compression and tension and is equal to 7/8d (m). Theoretical versus observed ultimate strengths The observed maximum ultimate strengths (Qmax) of the test specimens were compared to their predicted theoretical values Vsu and Vmu to investigate the variation between these values. The ratio of experimental to theoretical values, that is, Qmax /Vsu and Qmax /Vmu of all the specimens are shown graphically in Figs.18 (a) and (b). The Qmax /Vsu values of almost all the specimens which failed in shear mode shown in Fig.18 (a) is seen below 1.0 indicating that the theoretical values are slightly being overestimated. However, the Qmax /Vmu values of almost all the specimens which failed in flexure mode shown in Fig.18 (b) are more or less equal to 1.0 which implies that the theoretical value for ultimate flexural strength can be well predicted by the existing equation.

(a) Specimens that failed shear mode (b) Specimens that failed flexure mode

Fig. 18 Comparison of experimental values with calculated values

0

0.5

1.0

1.5

2.0

2D-L

1-H

0V0-

48-1

2D-L

1-H

0V0-

48-C

B-1

2D-L

1-H

42V

0-48

-1

2D-L

1-H

0V0-

84-1

2D-L

1-H

0V0-

84-C

B-1

2D-L

1-H

42V

0-84

-1

3D-L

1-H

0V0-

48-1

3D-L

1-H

0V0-

48-C

B-1

3D-L

1-H

42V

0-48

-1

3D-L

1-H

0V0-

48-C

B-2

3D-L

1-H

42V

0-48

-2

3D-L

1-H

0V0-

84-1

3D-L

1-H

0V0-

84-C

B-1

3D-L

1-H

42V

0-84

-1

3D-L

1-H

0V0-

84-C

B-2

3D-L

1-H

42V

0-84

-2

Qmax

/ Vsu

Sp

ecim

en

0

0.5

1.0

1.5

2.0

2D-L

0-H

0V0-

48-1

2D-L

0-H

0V0-

84-1

3D-L

0-H

0V0-

48-1

3D-L

0-H

0V0-

84-1

2D-H

0-H

0V0-

48-2

2D-H

1-H

0V0-

48-2

2D-H

1-H

0V0-

48-C

B-2

2D-H

1-H

42V

0-48

-2

2D-H

0-H

0V0-

84-2

2D-H

1-H

0V0-

84-2

2D-H

1-H

0V0-

84-C

B-2

2D-H

1-H

42V

0-84

-2

Qmax

/ Vmu

Sp

ecim

en

CONCLUSIONS

Based on the observations during tests and analysis of data, the following conclusions were obtained. � The confined masonry (CM) wall system is effective to improve the poor seismic performance of the

ordinary URM, by enhancing the lateral load carrying capacity. � The confined masonry (CM) wall system with connecting bars at the vertical wall-to-wall connections

as well as the horizontal wall reinforcing bars developed reasonably higher ultimate lateral strength with the increase of vertical axial load and showed better ductility as compared to the unreinforced wall specimens.

� The wall separation effect from the R/C confining columns can be avoided by providing the U-shaped connecting bars at the wall-to-wall or wall-to-column joints as recommended in China.

� The increase in axial stress tends to increase the lateral load carrying capacity of the masonry walls and the observed values showed that the ultimate flexural strength could be well predicted by the existing equation.

In brief, it can be concluded that the horizontal wall reinforcement and/or connecting bars provided between masonry walls and R/C columns play an important role to improve the poor seismic performance in the ordinary URM and CM walls, by enhancing the ductile behavior to some extent and lateral load carrying capacity.

ACKNOWLEDGEMENT

The authors deeply express sincere appreciation to Japan Society for Promotion of Science (JSPS) for financial support that made possible to carry out this experimental study. Further, thanks are extended to Mr. T. Hiramatsu, Ms. F. Kankura and all other students of the Structural Engineering Laboratory of Oita University for their assistance during the study.

REFERENCES

1. Yoshimura, K., T. Croston, H. Kagami and Y. Ishiyama, “Damage to Building Structures Caused by

the 1999 Quindio Earthquake in Colombia,” Reports of the Faculty of Engineering, Oita University, Vol.40, September 1999, Oita, Japan, pp.17-24.

2. National Standards of China, 1989: National Standards of P. R. of China, “Seismic Design Standard for Building Structures (GBJ 11-89),” 1989, p.987, in Chinese

3. Architectural Institute of Japan (AIJ) Committee for Concrete and Masonry Wall Building Structures, “Standard for Structural Design of Reinforced Concrete Boxed Wall-Buildings,” 1997, pp.127-128, in Japanese.

4. Sheikh, S.A., and S.S. Khoury, “Confined Concrete Columns with Stubs”, ACI Structural Journal, Vol. 90, No.4, July-August 1993, pp.414-431.

5. National Standards of P.R. China, “Seismic Design Standards for Building structures” (GBJ 3-88, pp.9, 1989, in Chinese.

6. AIJ Standards for Structural Calculation of Reinforced Concrete Structures - Based on Allowable Stress Concept, 1999, pp.54, in Japanese.

7 Matsumura, A., “Shear Strength of Reinforced Masonry Walls”, Proceeding of 9th World Conference on Earthquake Engineering, pp. VI-121-126, Tokyo-Kyoto, 1988.

8. AIJ, “Ultimate Strength and Deformation Capacity of Buildings in Seismic Design”, 1999, pp.592-593, in Japanese.

9. National Standards of P.R. China, “Seismic Design Standards for Building structures” (GBJ 11-89, pp.35, 1990, in Chinese.