ACI STRUCTURAL JOURNAL TECHNICAL PAPER Title no. 85-S41

Effectiveness of Welded Wire Fabric as Shear Reinforcement in Pretensioned Prestressed Concrete T·Beams

by Xiaoyi Xuan, Sami Rizkalia, and Kyuichi Maruyama

A lotal 0/ six pretensioned prestressed single T-beams with identical flexural reinforcements and shear span-la-depth ratio were tested statically up to failure to investigate the effectiveness oj welded wire fabric as shear reinforcement. The tested beams featured one beam without shear reinforcement and five beams with different types of shear reinforcement. including conventional double-legged stirrups, single-legged stirrups, and three different types of commercially available welded wire fabric (WWF).

The results indicated that the effectiveness of WWF as shear rein­forcement is the same as that of conventional stirrups under static loading conditions. It was found that using deformed WWF slightly improved the distribution of the diagonal cracks. Anchorages of WWF by means of two horizontal wires at the top and bottom are adequate according to A CI Building Code recommendations.

Keywords: beams (supports); cracking (fracturing); crack width and spacing; failure; failure mechanisms; T ·beams; prestressed concrete; pretensioning; reinforcing steels; shear strength: stirrups; strains; welded wire fabric.

The use of welded wire fabric as shear reinforcement has become popular among precast concrete manufac­turers because of its relative ease of placement and sav­ing of time and money due to reduced cutting, bend­ing, and labor. The use of welded wire fabric (WWF) as shear reinforcement also is promising for prestressed precast concrete members with deep thin webs. How­ever, few research papers are available regarding the use of WWF as shear reinforcement. ,.,

Welded wire fabric is manufactured from cold-drawn steel wires and generally tack-welded in an orthogonal mesh. The strength of WWF is usually greater than that of conventional stirrups. However, the cold-drawing process significantly decreases the ductility of WWF. The ultimate strain of the WWF is usually less than 2 percent while that of conventional stirrups is greater than 15 percent. The anchorage of WWF is provided by the bond of the vertical wires and the dowel action of the longitudinal wires welded on the vertical wires. The PCIIWRI Committee' recommended the use of two longitudinal wires at top and bottom for the anchorage

ACI Structural Journal I July-August 1988

of smooth WWF and one longitudinal wire for the de­formed WWF. The ACI Building Code' required using two horizontal anchorage wires at top and bottom for both smooth and deformed WWF. Conventional stir­rups are normally hooked at both top and bottom lon­gitudinal reinforcements. This configuration provides not only good anchorage but also good confinement of concrete and shear transfer by the dowel action of the longitudinal reinforcement. The welds of the WWF in­troduce additional weakness, due to the stress concen­tration, which could be serious under cyclic loading conditions.

RESEARCH SIGNIFICANCE This research program was undertaken to evaluate

the effectiveness of welded wire fabric as shear rein­forcement in pretensioned prestressed concrete T-beams under static loading conditions . The effectiveness will be measured in terms of the shear behavior, including shear crack initiation, shear crack pattern, crack width, and ultimate strength as compared to commonly used conventional stirrups. The experimental program also included various types and configurations of commer­cially available WWF to compare their behavior to the conventional stirrups.

EXPERIMENTAL PROGRAM Test specimens

A total of six pretensioned prestressed single T -beams with identical flexural reinforcement and shear span-to­depth ratio were tested statically up to failure. The tested beams featured different types of shear rein­forcement, including conventional double-legged stir-

Received June 10, 1987, and reviewed under Institute publication policies. Copyright © 1988, American Concrete Institute. All rights reserved, including the making of copies unless permission is obtained from the copyright proprio etors. Pertinent discussion will bc published in thc May·June 1989 ACI Struc· tural Journal if received by Jan. 1, 1989.

429

Xiaoyl Xuan is a graduate student in the Department of Ciyil Engineering at the University of Manitoba, Winnipeg, Manitoba, Canada. He received his 8Sc degree f rom Zhejiang Engineerjng q ollege. China, and his MSc f rom the Uni­yersity of Manitoba.

A CI member Sam; H. Rizkafla is an associate professor and Head of the Structures Division of the Civil Engineering Department at the University of Manitoba. He received his BSc degree from Alexandria University, Egypt, and his MSc and PhD degrees fro m North Carolina Slate University. He ;s a mem­ber of ASCE, the Prestressed Concrete Instilute, and the Canadian Society for Civil Engineering. He is also a member of the Executive Commillee of the Structures Division of the Canadian Society for Civil Engineers. His research interests are reinforced and prestressed concrete structures.

Kyulchl Maruyama is an associate professor, Civil Engineering Department, Univers ity of Nagaoka, Japan. He received his BSc and MSc degrees from the University of Tokyo, Japan, and his PhD f rom the University of Texas at Aus­tin. He is a member of the Japan COncrete Institute. His research interests are reinforced and prestressed concrete structures.

rups, single-legged stirrups, and three different types of commercially available welded wire fabric. A beam without web reinforcement was also included in this program to evaluate the contribution of concrete. All beams were designed with a nominal flexural strength higher than the shear strength.

The longitudinal tensile reinforcement consisted of three 20M-400 MPa deformed bars, two 15M-400 MPa deformed bars, and two 13 mm seven-wire stress-re­lieved prestressed strands with a specified minimum

A

I Applied Load •

Symm.

~

i[L ____ ~~----------~ "" Support

750" 1150 350 I Total s'oon • 4500 mm - r

650mm

- '-"7T- '---,.,,==-=======r.:1'1 Wire Mesh ~ 4x4 -W4 xW4

E E o '" v

2.-_13mm Prestressed Strands

3-20M--<~ ,*--- 2-15M

Section A - A

Fig. J - Dimensions of specimens

430

tensile strength of 1860 MPa. Two 95 mm (No. 3)-400 MPa deformed bars were used as longitudinal com­pressive reinforcements_ The beam flange was rein­forced with a 4 x 4-W4-W4 welded wire mesh. The di­mensions and cross section details of a typical beam are shown in Fig. 1. The various shear reinforcement con­figurations and their material properties used in this program are given in Table 1 and Fig. 2. Typical con­figuration of the welded wire fabric used in this pro­gram is shown in Fig. 3. It should be noted that the lack of ductility of the vertical wires of the deformed wire fabric is based on independent testing of nine specimens.' For a constant stirrup spacing s of 152 mm, the steel area of the shear reinforcement A. was se­lected according to the measured yield strength /y and the measured concrete strength f ; to have similar steel contribution A,f/ b. sf; , as close as possible for all the beams tested in this program where b. is the web width of the beam.

Concrete was obtained from a local ready-mixed concrete supplier. The mix contained normal portland cement and local river aggregates. The main properties (by weight) per cubic meter were: cement 325 kg, sand 925 kg, gravel 1000 kg, water 170 kg; maximum aggre­gate size 15 mm; slump 100 mm. The average concrete strength , based on 150 'x 300-mm cylinders cast and tested simultaneously with the beams, was 34.3 MPa. The beams were cured by sprinkling water and were covered with plastic sheets for seven days. The pre­stress force was released at the eighth day. Prestressing forces were measured by load cells and strain gages at­tached to the strands, which were also monitored to evaluate the prestressing losses . Age at testing varied from 28 to 42 days.

Testing apparatus and procedure Each beam was loaded directly with two equal con­

centrated loads with a span of 3000 mm and a shear span-to-depth ratio of 2.9. The test setup for a typical beam during testing is shown in Fig . 4. The average concrete strains were measured using demec points that were attached to the concrete surface to cover all pos­sible areas of diagonal cracking. Crack widths and the slides (relative movement) along the diagonal cracks were measured at different locations using three sets of Demec points, as shown in Fig. 5_

The steel strains were measured using electrical resis­tance strain gages attached to the prestressed and non­prestressed longitudinal reinforcements as well as the stirrups. Strain gages were also used to check anchor­age, slippage of the reinforcement, and to monitor the prestressing force in the prestressed strands.

Beams were loaded in small increments of load var­ied between 10 and 15 kN to determine, with reason­able accuracy, the loads corresponding to the initiation of the first flexural crack and first diagonal shear crack. The load increments were increased to 30 kN up to failure. At each increment, while the displacement was

ACI Structural Journal I July-August 1988

Table 1 - Properties of shear reinforcements Shear

Yield Ultimate reinforcement

ratio Type of Size strength strength Ultimate A./, shear (area), i .p i • . elongation,

Beam reinforcement mm' M • MP. percent b.si: PSN I-O None - - - - -PSN2-WD WWF D4.7 deformed 587· 587 6.0 0.0291

(28.4)

PSN3-D2 Double-legged NO. 2 smooth 344 510 28.0 0.0372 stirrup

PSN4-WDH WWFwith 04.7 deformed 593 · 603 6.0 0.0294 additional (28.4)

horizontal wire

PSN5-S6M Single-legged 6 mm de fo rmed 483· 693 10.5 0.0266 stirrup (31.17)

PSN6-WS WWF W4.7 smooth 578- 61 8 8.5 Om08 (30 .58)

·Yleld strength computed accordmg to Clause 3.5.3.2 of ACt 31 8M-S3.

20 ____ I}...:~~~N __

20

___ .!\C~~~_ 38mm z z H 116

'" '" ..r38mm

." 10

~~L ." 10 U 0 0

0 0 ..J ..J

23 * H70

0 0.1 0.2 0 0.1 0.2 Elongation (mm) Elongotion (mm)

PSN5 - S6M PSN3 - D2

20 20 20 ___ Pu..:18.9 KN

_l'.!!.~~~~ Pu =17. 1 ~~

z z '" z '" '" 8 ..,

! -g10 I F

0 10 .., 10 0 0 ..J 0

.3 460mm ..J

0 0.1 0.2 0.1 0 .2 0 0.1 0.2

Elongation (mm) Elongation (mm) Elongation (m m)

PSN2 - WD PSN4-WDH PSN6- WS

Fig. 2 - Load-elongation curves for shear reinforcements

held constant, readings of load, displacement of actua­tor cylinder (stroke), strain gages, and deflections by LVDTs were measured and recorded by an H.P. Data Acquisition System. Lengths of cracks were marked and readings of demec points were taken. For detailed information, consult Reference 7.

ACI Structural Journal I July-August 1988

1IIIII III IIIIIII rn Fig. 3 - Typical configuration of welded wire fabric

431

Testing Machine

Lood Cell

Spreader Beam

" II Tes! Beam II II

" I' " I

II "

Fig. 4 - Test setup

I

, b.

c;r- Demec Point \y ~

Iv Jil~ " "

4PicoiCrock

I I I I

¢ I. Gage Length

Fig. 5 - Demec gage points

Table 2 - Test results

Flexural crack Slump. J: . initiation (2P).

Beam mm MP, kN

PSNI-O 60 36. 1 130

PSN2-WD 60 38 .1 120

PSN3-D2 100 33.3 130

PSN4-WDH 100 3 1. 5 120

PSN5-S6M lOS 32.5 120

PSN6-WS 105 34.3 150

W: Crock Width S: Slide

Shear crack Ultimate initiation (2P). load (2P),

kN kN

208 m 195 435.4

210 516.2

210 494.2

210 507 .7

210 507.5

TEST RESULTS AND DISCUSSION General behavior

Load-deflection at midspan measured for all six beams tested in this program is shown in Fig. 6. Load­stroke curves for all the beams are shown in Fig . 7 since all the L VDTs underneath the specimen were removed before failure . The deflection behavior of all the beams was very similar up to the initiation of the first diago­nal shear cracks. The deflection of the beam without shear reinforcement (PSNl -O) increased rapidly after initiation of the shear cracks, compared to beams with

432

'40

roo Z " ~160

~ ~ 120

'" ~ • :a. 80 ~

<I

4

C PSNI-O + PSN2 - WO <> PSN3 -02 t::. PSN4 - WOH x PSN5 - S6M

'Q PSN6 - WS

8 12 16

Deflection of Mid- SpOIl (mm) • A

Fig. 6 - Load-deflection at midspan

'0

~o ~----~p~~p~~~---------------:~~~~~ n , '

' 00

z " ~ 160

j 120

'" I 80 ~ <I

is --_____ _ 6

CPSNI - O + PSN2 - WD <> PSN 3 - 02 t::. PSN4 - WOH It PSN5 - SSM

'Q PSN6 - WS

4 8 12 16

Stroke (mm) , 8

Fig. 7 - Load-stroke curves of specimens

zo

shear reinforcement. Failure of the beam with de­formed WWF (PSN2-WD) occurred at a considerably lower deflection value than that o f the other beams with other shear reinforcement. This behavior could be attributed to the lack of ductility of the deformed WWF. The additional longitudinal wire at the mid­height of the deformed WWF in Beam PSN4-WDH did not influence the stiffness of the beam. However, it did enhance slightly the ductility of the beam as well as the ultimate shear strength of the beam.

Shear behavior Concrete properties and measured results of all tested

beams are summarized in Table 2 . Shear strength , which corresponds to the initiation of the first diagonal crack V", and ultimate shear strength V, are compared to the predicted values using different design codes'·' ·10 and are given in Table 3. Table 3 indicates cJearly that the ACI Building Code' and the simplified method of the Canadian Code' provide the best predictions for this type of prestressed concrete T-beams . Due to the low shear reinforcement ratio used for these six beams, the general method of the Canadian Code' underesti­mated the ultimate strength of the beams by 40 to 58 percent.

ACI Structural Journal I July-August 1988

Table 3 - Shear strength ratios

ACI CSA simple

Beam C' U' C U

PSNI-O 0.98 1.76 0.99 1.79

PSN2-WD

PSN3-D2

PSN4-WDH

PSN5-S6M

PSN6-WS

c v" measured • '" V. predicted

! U = ",V';em",,:::,,;;:u:.:":;-d V~ predicted

0.90

1.02

1.04

1.03

1.01

1.43 0.93 1.47

1.63 1.01 1.62

1.70 1.01 1.68

1.79 1.01 1.77

1.69 1.01 1.69

The predicted nominal shear strength provided by concrete, according to the ACI Building Code, was close to the tested values of applied shear force correspond­ing to the first initiation of the diagonal cracking.

Test results given in Table 3 indicate that the mea­sured ultimate shear capacity of the beam without web reinforcement (PSNI-O) is 76 percent higher than the predicted value according to the ACI Building Code. This behavior is attributed to the low ratio of shear span-to-depth and the high flange width-to-web width ratios for this category of beams. Consequently, the shear crack strength V" could be reasonably used as the contribution of the concrete for shear reinforced con­crete members. The prediction according to the ACI code underestimates the ultimate shear capacity for beams with shear reinforcement by 43 to 79 percent. The conservative predictions are mainly attributed to the measured shallow angle of the diagonal cracks of 25 deg instead of the 45 deg assumed by the ACI code.

The contribution of the shear reinforcement based on the difference between the measured ultimate load V. and load corresponding to the cracking load V" is com­pared to the yield strength of the stirrups A . 1;, in Fig. 8. The high values obtained in Fig. 8 confirm the fact that the shear-carrying mechanism is not as simple as assumed by the 45 deg truss analogy. However, the consistency of the results indicates that the effective­ness of WWF as shear reinforcement is similar to that of conventional stirrups.

The effectiveness of the shear reinforcement could also be evaluated based on the ratios of the measured ultimate strength and predicted values using different design codes, as given in Table 3. It is clear that the ef­fectiveness of the smooth WWF and the deformed WWF with an additional horizontal wire at mid height as shear reinforcement is the same as that of the single­legged and double-legged stirrups. The lower ultimate strength value of the beam reinforced with deformed WWF (PSN2-WD) could be attributed to the lack of ductility of the WWF, as shown in Fig. 2.

Failure mechanism All tested beams failed in shear. The first diagonal

shear cracks were initiated in each shear span of beams

ACI Structural Journal I July-August 1988

JSCE C U

1.26 2.27

1.I6 1.71

1.31 1.90

1.33 2.02

1.32 2.13

1.29 2.00

4 .0

'-1 1f 3.5 l' _>'3.0

> <t '- 2.5 ~

~ 2.0

::l 1.5 > ~

1.0

0.5

0 PSN2 -wo

CEB C U

1.28 2.30

1.I5 1.72

1.37 2.16

1.42 2. 15

1.39 2.23

1.34 2.08

PSN3 -02

CSA general

U

-1.72

2.07

2.3 1

2.51

2.26

PSN4 -WOH

PSN5 -S6M

PSN6 -WS

Fig. 8 - Measured contribution of shear reinforce­ment

at an angle of 45 to 65 deg to the horizontal axis of the beam. For the beam without web reinforcement (PSNI-0), the failure occurred suddenly due to the propagat­ing and widening of the diagonal crack up to the load point and down to the bottom of the beam at the sup­port.

For beams with shear reinforcement, additional di­agonal cracks were formed with flatter angles as the load increased . For the beam with deformed WWF (pSN2-WD), the failure occurred suddenly by fracture of the vertical wires at the position crossed by the di­agonal shear cracks, due to the lack of ductility of that WWF. For the four beams with other types of shear reinforcement, as the diagonal crack widened, stirrups crossed by the failure crack yielded and stress redistri­bution took place before failure. Further loading caused formation of transverse cracks across the top of the flange at the central region of the failure shear span, as shown in Fig. 9. The failure crack propagated

433

Fig. 9 - Failure cracks

PSN /- 0 W~{II"I,t\~ t t

PSN3-02

~ ~ 'd,mo",,' C",k,

~u,I,ID IV t t

PSN6- WS

Fig. 10 - Crack patterns

up to the load point and down to the bottom of the beams at the support. This diagonal crack damaged the top anchorages of the stirrups close to the load point and the bottom anchorages of the stirrups close to the support. At this stage, significant drop of the load transferred by the stirrups was measured. This mecha-

434

msm introduced a sudden increase of the tension force in the remaining stirrups crossed by the failure diago­nal crack. For beams with smooth WWF (PSN6-WS) and with deformed WWF with an additional horizontal wire (PSN4-WDH), three vertical wires were fractured. For the other two beams with conventional stirrups, the diagonal cracks propagated into a horizontal crack along the transition of the web and flange, causing damage of the anchorages of the stirrups, as shown in Fig. 10. Failure occurred suddenly after formation of the diagonal cracks shown in Fig. 10.

The measured strains in the stirrups and the vertical wires of WWF were negligible before the initiation of the diagonal cracks. After cracking, the strains were dependent on the relative positions of the cracks to the strain gage locations and the crack widths. No loss of bond or anchorage in the stirrups or the vertical wires of WWF was observed, since the measured strains in­dicated continuing increase up to yielding by increasing the applied load. Thus, the anchorage of WWF by us­ing two horizontal wires was sufficient.

Inspection of the test specimens after failure indi­cated that all fractures of the vertical wires of the WWF occurred at the locations crossed by the diagonal cracks rather than the weld points. This behavior reflected that the stress concentration of the weld does not affect the behavior of the WWF.

Crack patterns Crack patterns of all the beams were similar, as

shown in Fig. 9 and 10. For Beam PSNI-O without web

ACI Structural Journal I July-August 1988

Z ISO

'" a.

~

0 ... .r; (J)

'0

~ 0-0-«

100

50

o PSNI -0

+ PSN2-WD

<> PSN3 - 02

~ PSN4- WDH

)( PSN5 - S6M

v PSN6-WS

°0~--~----~--~----~2~--~----3~--~

Moximum Crock - Width (mm)

Fig. 11 - Applied shear: maximum crack width

reinforcement, the number of diagonal cracks was con­siderably less than that of the beams with shear rein­forcement. Beams reinforced with deformed WWF (PSN2-WD and PSN4-WDH) had the most uniform and the largest number of cracks measured at the mid­height of the section. The additional longitudinal wire at midheight of the WWF had no significant influence on the crack pattern, as is evident by comparing Beams PSN2-WD and PSN4-WDH in Fig. 10. In beams rein­forced with conventional double-legged and single­legged stirrups, additional diagonal cracks were ob­served he fore failure due to the redistribution of the stresses, as a result of the ductility of the steel, as shown in Fig. 10. However, a major horizontal crack was observed along the longitudinal reinforcement due to the reduction of the concrete cover, as a result of us­ing single-legged stirrups for Beam PSNS-S6M. The beam with smooth WWF (PSN6-WS) showed a similar cracking pattern in comparison with the beams with conventional stirrups.

Crack width At each load stage, the maximum crack width Wmo>

was calculated by using the measured deformations in three directions at different locations of the beam, as shown in Fig. S. The crack width W was measured by the movement perpendicular to the crack. The move­ment parallel to the shear crack was used to measure the relative movement along the crack. Both quantities were determined based on the measured crack angle and the deformation in two pertinent directions. The maximum crack width for all the tested beams is shown in Fig. 11. The total crack width was also calculated based on the summation of the crack widths, within the two shear spans, at each load increment for all beams,

ACI Structural Journal I July-August 1988

z '" a.

~

0 ... .r; (J)

'0 ... Ii 0-« 50

oPSNI-O

+ PSN2- WD

<> PSN3- 02

~ PSN4- WOH

){ PSN5 - S6M

v PSN6 - WS

°0~--~4----~8~--~'2~--~1~6--~2~0--~2~4~~28·

Totol Crock Width (mm)

Fig. 12 - Applied shear: total crack width

as shown in Fig. 12. Slides along the cracks were found to be very small in comparison with the crack widths . Thus, the total relative movement within the beam was figured only based on the total crack width.

Both maximum crack widths and total crack widths were essentially the same for all the tested beams up to the initiation of major diagonal shear cracks. At higher load levels, beams with smooth WWF (PSN6-WS) and double-legged stirrups (PSN3-D2) exhibited slightly smaller crack widths than the other beams. However, the amounts of shear reinforcement used in those two beams were also slightly larger than the other beams, as given in Table I. The beam with the single-legged stir­rups (PSNS-S6M) had relatively the largest maximum crack width, which could be attributed to the relatively smallest amount of shear reinforcement used in this beam. The same trend could also be observed in terms of total crack width, as shown in Fig. 12. Accordingly, the variation of the crack widths of the beams with shear reinforcement was mainly proportional to the amount of shear reinforcement rather than the stirrup configuration. The additional horizontal wire at the mid height of the beam in WWF made no improvement in terms of crack width, as is evident by the identical behavior of Beams PSN2-WD and PSN4-WDH.

CONCLUSIONS Based on the test results, the following conclusions

can be drawn: I. Crack widths were essentially the same for all

types of shear reinforcement configuration up to the initiation of major diagonal cracks. At higher load level, the slight differences in the crack widths were in­fluenced by the slight variation of the percentage of shear reinforcement regardless of the type and config­uration of shear reinforcement.

435

2. Anchorages of WWF by means of two horizontal wires at the top and bottom of the vertical wires were sufficient. The stress concentration due to tack-welds and limited ductility of the cold-drawn wires does not influence the effectiveness of WWF as shear reinforce­ment. Using deformed WWF slightly improved the dis­tribution of the diagonal cracks in comparison to the conventional stirrups.

3. The effectiveness of WWF is the same as that of conventional stirrups in terms of crack width control and the shear strength of the beam.

4. The additional horizontal wire at the midheight of WWF had no significant influence on crack pattern, crack width, and the stiffness of the beam. However, it could enhance the ductility as well as the ultimate shear strength of the beam.

5. Welded wire fabric used as shear reinforcement should exhibit adequate ductility to insure the overall ductility of the member.

RECOMMENDATION Further research should address the behavior and the

effectiveness of WWF as shear reinforcement under cyclic loading conditions to examine possible problems related to anchorage and ductility.

ACKNOWLEDGMENTS This study was performed in the Structures Laboratory at the Uni­

versity of Manitoba, Winnipeg. The assistance given by Messrs. Mo­ray McVey. Ed Lemke, and Dave Fedorowich during the experimental program is sincerely appreciated. The research project was financed by the Nf:.tional Science and Engineering Research Council, Canada. Donation of the welded wire fabric provided by DUR-D-WAL Ltd., Chicago, Illinois, and the sleel forms and material provided by Con­Force Ltd., Manitoba, are greatly appreciated.

436

CONVERSION FACTORS 1 mm = 0.039 in.

1 mm l = 0.00155 in.l I kg ~ 2.21b

I kN ~ 0.225 kip 1 MPa = 0.145 ksi

REFERENCES 1. Baskin, K. S., "Behaviour of Concrete T-Beams with Welded

Wire Fabric as Web Reinforcement," thesis, University of Alberta, Edmonton, 1982.

2. Leonhardt. F., and Walther. R., "Welded Wire Mesh as Stirrup Reinforcement, Shear Tests of T-Beams and Anchorage Tests," Bautechnik (Berlin), V. 42, No. 10, Oct. 1965. (in German. English translation by W . Dilger, Wire Reinforcement Institute, McLean, Va.)

3. Robertson, I. N., and Durrani , A. J., "Shear Strength of Pre­stressed Concrete T-Beams with Welded Wire Fabric as Shear Rein­forcement," Structural Research Report No. 29, Department of Civil Engineering, Rice University, Houston, Jan. 1985.

4. Taylor, Michael A., and EI-Hammasi, Soleiman, "Web Crack­ing Behavior of Beams Using Welded Wire Fabric as Shear Rein­forcement," ACI JOURNAL, Proceedings V. 77, No. I, Jan. -Peb . 1980, pp. 12-17 .

5. " Welded Wire Fabric for Shear Reinforcement," Journal, Pre­stressed Concrete Institute. V. 25, No.4, July-Aug. 1980, pp. 32-36.

6. ACI Committee 318, "Building Code Requirements for Rein­forced Concrete (ACI 318-83)," American Concrete Institute, De­troit , 1983, 111 pp.

7. Xuan, X., "Effectiveness of Welded Wire Fabric as Shear Re­inforcement in Pretensioned Prestressed Concrete T-Beams," MSc thesis, University of Manitoba, Winnipeg, 1986.

8. "Code for the Design of Concrete Structures for Buildings." (CAN 3-A23-M84), Canadian Standards Association, Rexdale, 1984, 281 pp.

9. Japan Society of Civil Engineers, "Recommendations for Limit State Design of Concrete Structures," Concrete Library Interna­tional, No.4, Dec . 1984, pp. 32-35.

10. CE8-FIP Model Code for Concrete Structures, 3rd Edition, Comite Euro-International du Beton/Federation Internationale de la Precontrainte , Paris . 1978,348 pp.

ACI Structural Journal I July-August 1988