12TH INTERNATIONAL

BRICK/BLOCK Masonry c O N F E R E N C E

Ing

IN-PLANE SEISMIC RESPONSE OF REINFORCED CONCRETE MASONRY

By K. C. Voon 1, J.M. Ingham 2

, and B.J. Davidson 3

ABSTRACT

There is a general consensus within New Zealand that the existing masonry design standard, NZS 4230: 7990, is overly conserva tive in its treatment of masonry shear strength, restricting cost-effective masonry designo The first section of this paper pre­sents recent New Zealand research endeavours which have assisted the development of design rules in the non-specific masonry design code, NZS 4229: 7999. These tests have been reported in New Zealand but have not been previously reparted elsewhere and are likely to be of interest ta designers in regians af high seismicity. Then, the pa­per reviews relevant experimental and analytical studies investigating the shear strength of cancrete masanry, which have been canducted since first preparatian af NZS 4230:7990. The paper concludes with some preliminary comments related to haw this recently published masanry research may influence shear provision in the next update of NZS 4230.

Key words: Seismic respanse, reinfarced masonry, code develapment, structural tes­ting, limited ductility

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1. INTRODUCTION

For many decades, masonry has been used as a common structural material in a lar­ge proportion of ali New Zealand building projects. However, since 1931 several ge­nerations of New Zealand engineers have been brought up on the 'Iesson' of the 1931 Napier earthquake (Scott, 1999) which, like many others throughout the world before and since, highlighted the extreme vulnerability of unreinforced masonry sys­tems to seismic attack. Consequently, New Zealand was amongst the first countries to develop reinforced masonry seismic design procedures based on the principie of capacity design (Priestley, 1980), which requires the dependable shear strength to ex­ceed the maximum lateralloading necessary to develop the wall flexural overstrength.

New Zealand concrete masonry manufacturers have actively promoted the use of rein­forced concrete masonry as a building medium for small-scale commercial, industrial, and residential structures, and identified the need to simplify the design procedure for this structural type. Consequently, the New Zealand concrete masonry industry has re­cently developed a non-specific design code, NZS 4229:1999 "(ode of Practice for (oncrete Masonry Building not requiring Specific Design", targeting new construction of partially and fully grouted single and double storey concrete masonry buildings. This non-specific design code is suitable for use by architects and architectural draftsmen, and is a time-saving resource for structural engineers. The design rules within this non­specific design code were developed following a comprehensive testing program con­sidering structural walls of varying geometry, including penetrations. The focus of this testing was on nominally reinforced walls designed to develop limited ductility during an earthquake, and details of the testing are reported in Section 2 of this paper.

Apart from the recent development of the non-specific design code, there is a gene­ral consensus among New Zealand researchers and designers that the existing ma­sonry design standard, NZS 4230:1990, has a number of criteria that are overly con­servative, which has restricted cost-effective masonry designo One such section is that related to masonry shear strength. The commentary to the masonry design standard (NZS 4230:Part2:1990) notes that "tests on masonry walls of both brick reinforced ca­vity masonry and concrete hollow unit masonry have indicated that properly desig­ned and detailed masonry shear walls can sustain average shear stresses well in excess of 2.0 MPa, while exhibiting a ductile flexural failure mode. It is now considered that the limits placed on the total shear in NZS 4230P:1985, were unduly conservative. Ho­wever, the extent of the experimental data remains insufficient to allow for other than a marginal relaxation of these limits in the context of Grade B and ( at this time".

2. CONCRETE MASONRY BUILDINGS NOT REQUIRING SPECIFIC ENGINEERING DESIGN

2.1. Introduction

This section of the paper publicises recent and current New Zealand research en­deavours related to the structural seismic design of reinforced concrete masonry

as this work has not yet received significant international exposure, and is poten­tially of interest to designers in seismically active countries. This research is pre­sented in the New Zealand context, which is currently a force-based limit-state design philosophy, with emphasis placed on the role this research played in the completion of the non-specific design code, NZS 4229:1999. This standard has application not only in New Zealand, but would readily serve as an appropriate document in many countries after applying only minor modification. The target readers of this standard are tradespeople rather than qualified engineers and it is therefore appropriate for the standard to be used in countries where a compara­tively poor understanding of special detailing and construction requirements may currently existo The standard enables simple masonry structures to survive signifi­cant lateral loads without loss of life.

2.2. Research conducted in support of NZS 4229:1999

Several research projects were initiated in response to the development of NZS 4229:1999. The first of these was conducted at the University of Au~kland (Bram­mer (1995), Davidson and Brammer (1996), Davidson (1996)). The researchers concluded that nominally reinforced masonry walls have reliable in-plane strength while demonstrating limited reserve ductility to ensure satisfactory seis­mic response. This information was used to develop the bracing capacity tables presented in NZS 4229:1999.

Two projects were conducted at the University of Canterbury. The first of these was conducted by Singh, Cooke and Buli (Singh (1998), Singh et aI. (1999)). The study established that ductile response could be achieved for long walls loaded out-of-plane. This study was then extended (Zhang, 1998) to investigate the per­formance of two walls that had door and window openings at structurally inap­propriate locations. The information gathered was then used in the development of the bond beam criteria in NZS 4229:1999.

2.3 Structural testing of Nominally Reinforced Concrete Masonry Walls performed at the University of Auckland

Test results from separate research projects conducted at the University of Auc­kland formed the basis of the non-specific lateral strength design procedure in NZS4229:1999. Details of these tests have not previously been reported outside New Zealand .

2.3. 1 5tructural testing by Brammer

Brammer (1995) performed quasi-static in-plane cyclic load tests on twelve no­minally reinforced cantilevered concrete masonry walls as tabulated in Table 1 be­low. Nine of these walls were partially grout-filled and the remaining three were

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Table 7. Test Wall Construction Details (ali rebar 300MPa yield, O denotes deformed bar and R denotes round bar).

Wall Width (mm) Length (mm) Effective Width, Vertical Steel Horizontal Steel Bond Beam Partially/Solid bw(mm) Stirrups f illed

1 90 1800 90 01 0 010 None Solid

2 90 2600 60 010 010 None Partial

3 90 2600 90 010 010 None Solid

4 90 4200 90 010 010 None Solid

5 140 800 60 012 016 R6 @600 Partial

6 140 1800 60 012 016 R6 @600 Partial

7 140 2600 60 012 012 R6@600 Partial

8 140 4200 60 012 016 R6 @600 Partia I

9 190 800 60 012 01 6 R6@600 Partial

10 190 1800 60 012 016 R6 @600 Partia I

11 190 2600 60 012 016 R6@600 Partia l

12 190 4200 60 012 016 R6 @600 Partial

Figure 7. Elevation of typical reinforced masonry long wall.

,-

2600

. ,."

I -

'r I

o ;

11 I ,FI J ~ I I STflJNGFlOOR

solid grout-filled . Ali walls were constructed to a height of 2400mm, but varied in length and width . None of the twelve walls had applied axial compression 10-ad. The tests were performed in order to compare the attained test behaviour with that assumed and predicted by the New Zealand masonry design code (NZS4230:1990). Particular attention was given to maximum strengths, stiffness, ductility, modes of failure, force-displacement characteristics, base course slip, and also t he shear and flexural components of displacement.

The walls were constructed in runn ing bond from standard production masonry precast units of varying width, with the relevant structural details shown in Table 1. The vertical steel was lap-spliced immediately above the foundation, and was generally spaced at 800mm centres as shown in Figure 1 (b); the exception being the two 800mm long walls which conta ined two vertical bars spaced at 600mm centres. The only horizontal steel in any wall consisted of two bars, as shown in

Figure 1 (b), placed in a bond beam within the top two block courses. The hori­zontal cyclic load was applied to the top of the wall via a steel channel as shown in Figure 1 (a), which was fastened to the top of the bond beam by cast-in bolts. The jack was fastened to the strong wall and was partially supported (vertically) by the test specimen . The wall was stabilized from moving in its out-of-plane di­rection by two parallel horizontal struts which were perpendicular to the wall and hinged to the channel and a reaction frame. The test procedure and method for evaluating the available displacement ductility used were that described by Park (1989).

Behaviour of Test Walls

The nominal wall flexural strength Fn, nominal wall shear strength Fy , the maxi­mum lateral force recorded during the test F ma< and the experimentally found bili­near yield (or ductility 1) displacements Oy are tabulated in Table 2. Two principal modes of failure were observed during the tests: a diagonal tension and a hinge­sliding mode. Diagonal tension failure is characterised by the development of early flexural cracking which is later exaggerated by diagonal cracking that ex­tends throughout the whole masonry wall, while a hinge-sliding mode is charac­terised by the propagation of a significant crack located at the wall / foundation in­terface. It was anticipated that the walls would fail in shear (diagonal tension) because of the lack of shear reinforcement, with this being preferable to the hin­ge-sliding mode, where the lateral force was resisted only by the dowel action of the vertical reinforcement once a crack opened up along the entire length of the wall/foundation interface (Priestley, 1976).

As Table 2 above shows, ali the partially grout-filled walls exhibited a diagonal tension mode of failure except for Wall 9 (800mm long x 140mm wide), which exhibited a combination of hinge sliding and diagonal tension . The two shortest

Table 2. Wal! Test Parameters and Made af Failure.

Wall F, (kN) F, (kN) F~ (kN) F~/F" F~JF, F./F, f, (MPa) õ,(mm) Observed Mode ~.

of Failure

1 28.3 38.9 32.2 1.14 0.83 1.35 0.20 1.0 Hinge Sliding 6.0

2 60.8 31.2 62.0 1.02 1.99 0.51 0.40 1.6 Diagonal Tension 3.0

3 56.8 56.2 62.8 1.11 1.12 0.99 0.27 0.6 Hinge Sliding 6.0

4 137.4 90.7 144.0 1.05 1.59 0.66 0.38 1.0 Diagonal Tension 4.5

5 10.8 15.1 12.4 1.15 0.82 1.40 0.26 5.5 Diagonal Tension 3.75

6 38.1 34.1 41.0 1.04 1.20 0.90 0.38 5.0 Diagonal Tension 4.5

7 75.0 37.4 80.4 1.07 2.15 0.50 0.52 3.1 Diagonal Tension 4.5

8 213.1 98.5 179.9 0.84 1.83 0.46 0.71 3.6 Diagonal Tension 1.0

9 11.0 13.2 13.2 1.2 1.00 1.20 0.28 2.4 Hinge Sliding/ 3.75 Diagonal Tension

10 44.3 25.9 47.0 1.06 1.81 0.58 0.44 1.1 Diagonal Tension 6.0

11 83.9 40.3 87.1 1.04 2.16 0.48 0.56 1.0 Diagonal Tension 6.0

12 203.7 60.5 197.3 0.97 3.26 0.30 0.78 1.5 Diagonal Tension 4.5

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solid filled walls (Wall 1 & 3) had a hinge-sliding mode of failure as characterised by both negligible cracking at test conclusion and significant base course sl ip.

From Table 2, it can be seen that ali walls other than Wall 8 & 12 obtained F maxlF n> 1.0, where F n was evaluated based on a rectangular masonry compression stress block:

(1 )

lhe table also shows that generally F m" was significantly larger than the nominal she­ar strength F" calculated from the equation provided by the New Zealand masonry design code:

(2)

where vm is the shear stress provided by the masonry, and is equal to the greater of 0.3 MPa or the sum of 0.03f'm and 0.3WJAg.

An important facto r can be identified from the relationship FjFn in Table 2, as the duc­tile diagonal tension mode developed even when FjFn<1.0. Also, except for Wall 8, ali walls having FjFn<1.0 had l-1,v 3.0. This indicates that the evaluated shear strength using Equation 2 is of limited relevance for structures having a reinforcement distri­bution as indicated in Figure 1 (b) and supporting little axial compression load, when required to develop only limited ductility (l-1av~2.0) . This is partially because Fv is not well predicted by the New Zealand masonry standard, NZS 4230:1990, but more im­portantly due to the frame action generated by use of the bond beam shown in Fi­gure 1 and the shear friction generated between blocks during lateral deformation.

The parameter (l-1,v is an estimate of the wall's displacement ductility capacity. Park (1989) assumed that four loading cycles to 1-1" would result in the same re­duction in wall strength as was experienced during the wall test. The ductility ca­pacity is derived from :

(3)

The values tabulated in Table 2 for 1-1", show that the largest recorded ductility ca­pacity for each wall width corresponded to a wall aspect ratio of approximately unity, and reduced for both the 800mm and 4200mm long walls. The less ducti­le walls of 800mm and 4200mm long displayed significantly greater cracking than the 1800mm and 2600mm long walls . It was deduced that the ductility ra­ting for the 800mm and 4200mm long walls were less because of the rapidly-de­veloping wide cracking that contributed to shear displacement, accelerated ini­tiation of the diagonal tension mode of failure, and created strength degradation.

Limited space prevents details of ali the tests from being reported. However, in­dicative response is demonstrated by considering more fully Wall 10 and Wall 1. These two walls are considered because Wall 10 exhibited a diagonal tension mo­de while Wall 1 failed in a hinge-sliding mode.

Wall 10 demonstrated significant strength and stiffness degradation after failure, particularly in the negative quadrant, as can be seen in Figure 2. It can be seen that the maximum strength was approximately 10% greater than the evaluated nominal flexural strength, and that this maximum force occurred during the /14 cycle in the positive quadrant. Failure occurred during the first negative excursion to J..ll O. The theoretical value for the force to cause first yield within the wall Fy is shown in the figure, as well as the point during the test when cracking was no­ted to be sufficiently wide for daylight to be seen through the wall, indicated by the notation "C".

Wall 1 in Figure 3 shows marked stiffness degradation, but significant strength de­gradation occured only in the positive quadrant. A similar value for the maximum strength occurred at the first positive excursion of both the J..l4 and 116 cycles. This maximum value was approximately 15% greater than the nominal flexural strength. Failure óccurred during the positive excursion of the second J..l8 cycle.

As Figure 2 and Figure 3 illustrate, the nominally reinforced concrete masonry walls detailed in Table 1 and 2 exhibited gradual strength degradation and in no case did any wall suffer from sudden failure. This desirable behaviour in the no­minally reinforced partially grout-filled and solid grout-filled masonry walls was created by the solid filled bond beam at the top of the walls, which caused a fra­me-type action at latter stage of testing. Also, strength 1055 and degradation due to the hinge-sliding mode of failure was not significantly greater than that due to diagonal tension failure, as can be seen in Figure 2 and 3, such that it was diffi­cult to distinguish the failure mode associated with various tests based on their measured force-displacement histories.

Figure 2. Farce-displacement behaviaur af Wafl 7 O. Figure 3. Farce-displacement behaviaur af Wafl 7.

I I I ,. : ....... 1-_ t- .. , .... '_,. _ _ c..- ...... r_

O.6· _..L..· 1

,. I

_l~"L _---'-_---''--.-'-, -'-, --:-,------'-.'---'-. --:-"----',, -,

:~ II .....: I~ i u 1:-: ::: ::::,:;::·IH-+r.t--f_ 'i'-h~~;-~+-:"j---l

I I I. I I

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2.3.2 Structural testing by Davidson

Brammer's research was extended to investigate the behaviour of walls with ope­nings and applied axial compression stress (Davidson, 1996). Two nominally rein­forced concrete masonry walls (4200mm long x 2400mm high x 190mm wide) were constructed so that they had an identical arrangement of a 2000mm x 600mm 'doorway' and a 1200mm x 600mm 'window', with the only difference being the magnitude of the applied axial compressive load. The 'doorway' and 'window' were arranged in a way to enable the vertical reinforcement to be pla­ced at 800mm centres.

Table 3. Test results for Walls with openings.

Wall o,(mm) Applied Axial Load (kN) Fmu (kN)

Al 1.2 O 90.0

A2 2.3 24 110.0

The test behaviour of Wall A 1 and A2 was similar to the test behaviour of Wall 10 and Wall 1 as described above, and it is important to note that Walls A 1 and A2 did not suffer sudden failure. A comparison of test results from Walls A 1 and A2 with Wall 12 in section 2.3.1 illustrates that the capacities of the two walls with openings were approximately half that of the closed wall, and that the compres­sion stress slightly increased the lateral strength of Wall A2 as tabulated in Table 3. It was concluded from this study that openings have a detrimental effect on the lateral strength of masonry walls while axial compression stress is beneficiaI.

2.4 Correlation of the University of Auckland test results with NZS 4229: 1999

The procedure employed in NZS 4229:1999 to establish the lateral strength of a reinforced masonry wall with vertical D12 reinforcement at 800mm centres is to determine the 'bracing capacity' based upon the height and length of individual bracing panels. The geometry of these bracing panels is dictated by the presen­ce of wall penetrations and shrinkage control joints, as shown in Figure 4.

In Table 4 the code predicted capacity of the wall tests reported in section 2.3 is correlated with the strengths obtained from experiment, noting that a 90mm structural wall width is not permissible in NZS 4229:1999.

As illustrates in Table 4, in ali cases there was a substantial safety margin betwe­en the measured wall capacity and that predicted using NZS 4229:1999.

3.0 Review of other Research

The data sources used in the preparation of the New Zealand masonry design standard, NZS 4230:1990, were published in 1980 or earlier. Consequently, the

Figure 4. Typica/ Masanry Wal! consisting af variaus Bracing Pane/s.

Shrinkage Conlrol Joinl

I ~~~~r--~::;:::=-i~~aoo;;;;;;--- less Ihan 600cmm \-O ... more Ih~ less Ihan 800mm aOOmm

Tab/e 4. Camparisan af Cade Predicted strength with Measured Strength.

Panel Equivalent Lateral4 Measured Wa ll Height Length Width

Bracing Capacity, f, ... Strength, f mu

f~, Units r:-(mm) (mm) (mm) (kN) (kN)

5 2400 800 140 155 7.8 12.4 1.59

6 2400 1800 140 498 24.9 41.0 1.65

7 2400 2600 140 920 46.0 80.4 1.75

8 2400 4200 140 1713 85.7 179.9 2.10

9 2400 800 190 155 7.8 13.2 1.69

10 2400 1800 190 510 25.5 47.0 1.84

11 2400 2600 190 943 47.2 87.1 1.85

12 2400 4200 190 1758 87.9 197.3 2.24

Al Hl=2000 Ll=1000 90.0 1.80

A2 H2=1200 L2= 1000 190 1000 50.0

H3=1200 L3= 1000 110.0 2.20

existing standard has a number of criteria that are overly conservative, including those related to masonry shear strength. As a significant volume of research on masonry has been completed since 1980, there is now considerable additional experimental and analytical information available related to masonry shear strength. However, only the following research was properly reviewed at the ti­me of writing this paper: Brunner and Shing (1996), Crisafulli (1995), Hidalgo and McNiven(1980), Larbi and Harris (1990), Schultz (1996), Shing et aI. (1989), and Shing et aI. (1990).

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3.1 Experimental Studies

Relevant details from the experimental studies previously listed are reported here. This consists of experimental testing of fully grout-filled masonry walls by Brunner and Shing (1996), Larbi and Harris (1990), and Shing et aI. (1989). The only experimen­tai testing of partially grout-filled masonry walls reported here is that of Schultz (1996). Note that ali reinforcement has been converted to equivalent metric dimensions.

Table 5. Summary of experimental studies by some researchers.

Reference Wa ll Reinforcement Width Aspect Ratio Compressive Fai lure number Vertica l Horizontal (mm) H/L Load Mode

Brunner and 1 4 x 19.1 4 x 9.5 137 1300/1400 1.9 MPa Diagonal Shing (1996) Tension

2 5 x 19.1 " 137 1300/1800 " "

3 6 x 19.1 " 137 1300/2200 " " Larbi and 1 - 4 5 x 15.9 5x9.5 47 610/6 10 1.86 MPa Diagonal

Harris (1990) Tension

5 - 7 5 x 15.9 5x l 2.7 47 610/610 1.86 MPa Flexure

8 - 10 5 x 19.1 5 x 9.5 47 610/610 1.86 MPa Diagonal Tension

Schultz 1 2 x 19.1 2 x 9.5 195 1422/2845 267 kN Diagonal (1996) Tension

2 " " " 1422/2032 191 kN "

3 " " " 1422/1422 133 kN "

4 " lxl2.7& 195 1422/2845 266 kN 1 x 15.9

5 " " " 1422/2032 177 kN "

6 " " " 1422/1422 132 kN "

Shing et aI. 1 5 x 15.9 5 x 12.7 143 1830/1830 1.38 MPa Flexure (1989) 2 " 9 x 9.5 " " 1.86 MPa Flexure

3 5 x 22.2 5 x 9.5 " " 1.86 MPa Shear

4 " " " " O Shear

5 " " " " 0.69 MPa Shear

Flexurel 6 5 x 15.9 " " " O Shearl

Sli de

7 5 x 22.2 " " " 0.69 MPa Shear

8 5 x 15.9 5 x 12.7 " " O Flexurel Shear

9 " 5 x 9.5 " " 1.86 MPa Shear

10 " " " " 0.69 MPa Flexurel Shear

11 5 x 22.2 5 x 12.7 " " O Shearl Sli de

12 5 x 15.9 " " " 0.69 MPa Flexure

13 5 x 19.1 " " " 1.86 MPa Shear

14 " 5 x 9.5 " " 1.86 MPa Shear

15 " 5 x 12.7 " " 0.69 MPa Flexurel Shear

16 5 x 22.2 " " " 1.86 MPa Shear

Prior studies done by these researchers concluded the inelastic behaviour of rein­forced masonry wall was highly sensitive to the applied axial stress and the amount of vertical and horizontal reinforcement present. Masonry walls with low vertical steel ratio exhibited a predominantly flexural behaviour, as shown in Ta­ble 5. Research reported here indicates that walls exhibiting flexural failure mode had more rapid load degradation when subjected to large axial stress due to mo­re severe toe crushing. The measured nominal strength of walls exhibiting shear failure, characterized by rapid strength degradation after failure, depended on the amount of horizontal reinforcement, the tensile strength of masonry, the do­wel action of the vertical steel, and the aggregate-interlocking mechanism, which in turn depended on the applied axial stress and the truss action of the vertical steel. The nominal shear strength tends to increase with the applied axial stress, and the amount of vertical and horizontal reinforcement present. However, the reinforcement content seems to have a more significant influence on the post­cracked ductility and energy dissipation capability than on the ultimate shear strength of a masonry wall (Shing et ai, 1989).

3.2 Analytical Studies

Experimental tests confirm that the shear strength of reinforced masonry walls is governed by several complicated mechanisms, such as aggregate-interlocking, dowel action of the flexural reinforcement, the truss action of the flexural and shear reinforcement, and the shear resistance of masonry at the compression toe of a wall pane!. With respect to these observations, recent formulae to predict the nominal masonry shear strength, Vn, have been proposed by various researchers, with the form of these equations given in Table 6:

(4)

Table 6. Equations for Masonry Shear Strength Calculation.

Masonry (omponet Horizontal Steel Vertical Load

BSSe (199B) 0.083 [4.0 - 1.75 ( ~ )]A, Jf: 0.5p.I~ A, 0.25P

Shing et aI. (1991) 10.0217p,f. + 0.166) A, K ( 2d' ) s L - S -1 l p,f', A. 10.0217)r.

Anderson and 0.2407 [f::A: O.5p.f, A, 0.25P Priestley (1992)

NZS 4230:1990 VmA, p.l, A, O

AS 3700·1998 tA, 0.8 p.f~ A, O

Notes: 1. Vm is the greater of OJ MPa or OJ (0.1 f'm + (l,) .. 0.72 MPa. 2. f. is equal to (1.50 - 0.5 H/L) MPa.

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The objective of this section was to review appropriate research, then compile this information in a form suitable for comparison with the masonry shear strength predictions given by NZS 4230:1990.

Figure 5. Camparisan af Shear Farmulae. (H=7830mm, L= 7 830mm, 0,= 7 .30 MPa, p,=0.38%, f,.,=440MPa, Ph=O, An=267 690mm2

).

Figure 6. Camparisan af Shear Farmulae.(H=7830mm, L=7830mm, 0,=7.30 MPa, p,=0.38%, f,.,=440MPa, Ph=0.24%, fyh=462MPa, An=267 690mm2

) .

ir:- ----~-~ .- ___ ~ ..... ·. fOf .......

I; 1.3 -- . -- BSscrofl,.,Q Úi - ........ -SI*'g·. 'oo .... S · ··)(.··NZS"230-' 9QO ~ 1.1 -G-- ASJ700-, ' 9!lfI _ ...... __ .... __ .... --

I) 0,9 __ _ - ---- - -- ..... -- ...... -- .... -

I -~ --~ --~ Ir 0.7 ---

.... - ... - .... - ... -

~--------------------,-.-,,-,~~)~ ! '---"---"'-- '8 f'~t""') 20

--~---

Figures 5 and 6 had been plotted in order to investigate the influence of ma­sonry compressive strength for an 1830mm high masonry wall. The structu­ral form selected was similar to Wall 1 tested by Shing et aI. in Table 5. The figures show that masonry shear formulae proposed by BSSC, Shing et aI., and Anderson and Priestley correlate well with each other, despite a notice­able margin separating the three . They also demonstrate that the shear strength provided by masonry component was conservatively predicted by the New Zealand masonry design standard at high value of, f' m, partly be­cause the beneficiai effect ofaxial compression was not considered by these standards. However, Figure 6 shows that both the New Zealand and Austra­lia masonry standards have over estimated the shear strength provided by the horizontal reinforcement, with respect to the other three recommenda­tions.

4. ADDITIONAl TESTING

Further testing is to be conducted at the University of Auckland before more ap­propriate design criteria for masonry shear strength may be incorporated into the next New Zealand masonry design standard. Initial testing will focus on pure she­ar failure so that masonry shear strength can be established. Some of the walls lis­ted in Table 1 will be tested without the bond beam in order to achieve the aim stated early. Fully reinforced walls wi ll be tested as a priority, then the study wi ll be continued to investigate the influence ofaxial load and partially grout-filled walls.

5. CONCLUSIONS

1. The maximum displacement ductility of partially grout-filled concrete masonry walls subjected to in-plane loading is attained for wall aspects ratio of approxi­mately unity.

2. Partially grout-filled masonry is a viable lateralload-resisting construction formo 3. The resistance mechanism of partially grout-filled masonry walls consists of fra­me action generated by the bond beam, dowel action of vertical steel and sliding friction between masonry panels and concrete surfaces.

4. NZS 4229:1999 is an appropriate code for the design of partially grout-filled concrete masonry walls.

5. The shear strength provided by the masonry component has been conservati­vely predicted by the New Zealand masonry design code. However, the New Ze­aland masonry standard had over estimated the contribution of horizontal rein­forcement.

6. ACKNOWLEDGEMENTS

The study presented in this paper is financially supported by EQC (Earthquake Commission Research Foundation) and the New Zealand Concrete Masonry As­sociation (NZCMA). Testing of partial grout-filled masonry was sponsored by Firth Industries Ltd, and the efforts of Hank Mooy and Bryan Mehaffy in conducting this testing are acknowledged. Completion of NZS 4229 is accredited to the as­sociated code committee, but the efforts of David Barnard as chair, and of An­drew Wilton as liason in regard to testing at the Universities of Auckland and Can­terbury is gratefully acknowledged .

7. REFERENCES

Anderson, O.L., and Priestley, M.J.N. (1992), "In Plane Shear Strength of Masonry Walls", Pro­ceedings of the 6'" Canadian Masonry Symposium, Saskatoon, Saskatchewan.

AS 3700(1998, "Masonry Structures", Standards Association of Australia, Homebush, NSW, Aus­tralia, 124p.

Brammer, O.R (1995), "The Lateral Force-Oeflection Behaviour of Nominally Reinforced Masonry Walls", ME Thesis, Oepartment of Civil and Resource Engineering, University of Auckland, New Zealand, 271 p.

Brunner, J.O, and Shing, P.B. (1996), "Shear Strength of Reinforced Masonry Walls", TMS Jour­nal, Vol. 14, No. 1, pp65-77.

BSSC (1998), "NEHRP Recommended Provisions for Seismic Regulations for New Buildings and Other Structures", Building Seismic Safety Council for the Federal Emergency Management Agency, Washington, O.C, Report No. FEMA-302, 337p.

Crisafulli, F.J., Carr, A.J ., and Park, R., (1995) "Shear Strength of Unreinforced Masonry Panels",

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proceedings of the Pacific Conference on Earthquake Engineering, Vol. 3, Melbourne, Australia, pp 77-86.

Davidson, B.J (1996), "In-Plane Cyclic Loading of Nominally Reinforced Masonry Walls with Ope­nings", New Zealand Concrete Society Conference, Wairakei, New Zealand, pp120-129.

Davidson, B.J and Brammer, D.R (1996), "Cyclic Performance of Nominally Reinforced Masonry Walls", Technical Conference of NZSEE, New Plymouth, New Zealand, pp144-151.

Hidalgo, P. A., and McNiven, H. D. (1980), "Seismic Behaviour of Masonry Buildings, Procee­dings of the 7th World Conference on Earthquake Engineering, Vo1.7, Istanbul, Turkey, Septem­ber, pp111-118.

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7. NOTATION

Ag = gross wall cross·sectional area

Ah = area of horizontal steel

Ao = net wall cross-sectional area

A, = area of vertical steel

b = width of compressive stress block

bd = deepness of masonry unit

bw = effective wall width

d = length of masonry unit in direction of shear force

d' = distance of the extreme vertical steel from the nearest edge of the wall

f 'm = masonry compressive strength

f'.b = masonry tensile strength

( = developed masonry shear stress

( , = masonry shear stress

F,od' = code predicted strength

Fm .. = experimental maximum lateral force

Fo = nominal flexural strength

F, = nominal shear strength

fy = steel yield stress

fyh

f",

H

= yield strength of horizontal steel

= yield strength of vertical steel

= wall height

jdo = lever arm to the nth bar measured from half the depth of compression block

= walllength L

M = moment on masonry section due to unfactored loads

P = axial load on masonry section due to unfactored loads

= vertical spacing of horizontal reinforcement

Vm = masonry shear stress

V = shear on masonry section due to unfactored loads

Vm = shear force provided by the masonry

Vo = nominal shear strength of reinforced masonry

V, = shear resistance provided by the horizontal shear steel

Vem = shear strength of unreinforced masonry

W, = wall self weight

õy = yield displacement

0 , = axial compression stress

1-\" = available displacement ductility factor

Ph = volumetric ratio of horizontal steel (based on net area)

P, = volumetric ratio of vertical steel (based on net area)

NOTES 1. Postgraduate Student, Department of Civil and Resource Engineering, University of Auckland,

Private Bag 92019, Auckland, New Zealand. 2. Cement and Concrete Association Lecturer, Department of Civil and Resource Engineering,

University of Auckland. Also, Executive Officer, New Zealand Concrete Masonry Association, P,O Box 448, Wellintong, New Zealand.

3. Senior Lecturer, Civil and Resource Engineering, University of Auckland, Private Bag 92019, Auckland, New Zealand.

4 . NZS 4229: 1999 defines 100 BU's as 5 kN.

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