Jl 99 January February 7

8/10/2019 Jl 99 January February 7

1/12

Shear ehavior of Concrete

eams

Prestressed

with

FRP Tendons

Sang

Yeol Park

Ph.D.

Full-time Lecturer

Department of Ocean a

nd

Civil Engineering

Cheju National University

Cheju,

Korea

Antoine

E

Naaman

Ph.D. P.E.

Professor

Department of Civil and

Environmental Engineeri

ng

University of Michigan

Ann

Arbor, Michigan

7

FRP reinforcements

h ve

excellent properties for use in

concrete

structures including high corrosion resist nce

nd

high tensile

strength. However they have

some

technical drawbacks particularly

their lack of ductility nd low transverse strength. This study deals

with an experimental investigation

of

the shear behavior

of

concrete

beams prestressed with

FRP

tendons. In the experimental program

the shear-tendon rupture failure mode was investigated in detail nd

experimentally confirmed. Shear tests showed that premature failure

due to shear-tendon rupture is likely to occur

in

concrete be ms

prestressed with

FRP

tendons resulting in

reduced

lo d carrying

capacity. The premature fai lure is due to tendon rupture by dowel

shear at the shear-cracking pl ne nd is ttributed to

the

brittle

behavior nd low transverse resistance of FRP tendons.

T

he applicability of Fiber Rein

forced

Polymer

FRP)

rein

forcements

to

concrete struc

tures as a substitute for steel bars or

prestressing tendons is being actively

studied in numerous research laborato

ries. This is primarily because FRP re

inforcement, in comparison to conven

tional steel reinforcement, offers some

excellent advantages, including non

corrosive, non-magnetic, high strength,

and lightweight properties.

In particular, non-corrosion is the

most important property for civil en

gineering infrastructures because the

deterioration due to corrosion causes

the most serious economic and techni

cal problems

in

repairing existing

structures in many countries. Even in

prestressed concrete structures, which

have excellent durability , corrosion is

considered the

main

cause

of

long

term deterioration .

2

Therefore,

FRP

reinforcements appear to be ideal sub

stitutes for steel reinforcement

in

con

crete structures.

On the

other

hand ,

FRP

reinforce

ments also

have some

disadvantages

such as non-plastic behavior, very low

shear

or

transverse strength , suscepti

bility to stress-rupture , and high cost.

From a structural engineering view

point, the most serious of these disad

vantages are the lack of plastic behav

ior and the very low shear strength in

the transverse direction. Such charac

teristics may lead to premature tendon

rupture, particularly when

combined

PCI JOURN L


2/12

effects are present, such as at shear

cracking planes where dowel action

exists in prestressed concrete beams.

The dowel action reduces residual ten

sile and shear resistance

in

the tendon.

Thus, it

is

expected that shear-tendon

rupture failure initiated by dowel ac

tion results in less shear resistance and

shear ductility in concrete members

prestressed with FRP tendons.

RESE RCH

SIGNIFIC NCE

The majority

of

research on con

crete structures using FRP reinforce

ments has been on members that are

not critical in shear. Because there

have been very few shear tests, the

shear behavior

of

prestressed

con-

crete members using FRP reinforce

ment is not well understood. Unlike

flexural behavior, shear behavior is

quite complex by itself, even in ordi

nary reinforced or

prestressed

con

crete members.

Furthermore, the experimentally de

rived prediction equation for the shear

capacity of prestressed concrete mem

bers using steel tendons has not

yet

been proven to be

applicable

when

FRP tendons are used. This is because

the mechanical characteristics

of

FRP

reinforcement, such as no yielding be

havior ,

low

shear

or transverse

strength, and low elastic modulus, are

significantly different from those

of

steel tendons.

The few researchers

3

6

who

have

studied the shear behavior

of

concrete

beams prestressed with FRP tendons

focused mainly on their shear strength,

not on the shear

failure

mode.

Nishikawa et

aJ.

were the first to re

port that prestressed concrete beams

using CFRP tendons with low residual

elongation ability lose their shear re

sisting capacity by tendon failure at

the shear crack. Jeong and Naaman

7

speculated on the possible causes

of

the FRP shear-tendon rupture that oc

curred in some

of

their flexural tests.

In addition, the JSCE Research Sub

committee on FRP

8

comments on the

possibility

of

lowered ultimate load

due to local stress

in

tensile reinforce

ment at the crack location by dowel

action , and ACI Committee 440

9

sug

gests that special attention should be

devoted to the reduced dowel contri

bution of FRP reinforcements in the

January-February 1999

Fig

. 1. Dowel action in concrete beam.

presence

of

shear cracks. However, to

the authors ' knowledge there has

been no study on this type of shear

failure in concrete beams prestressed

with FRP tendons.

ST TEMENT

OF PROBLEM

Longitudinal reinforcement, which

is designed primarily to resist flexural

tension , is often required to carry a

shear force by dowel action across a

diagonal tension crack. If the crack

opens (rotates) slightly, a shear dis

placement will result from the rotation

of

a beam about the crack tip and the

a)

Tendon

c)

shear slip due to the shear force along

the crack face.

To

resist differential

shear displacement between the crack

faces

,

the

bars

or

tendons develop

dowel shear forces. This counteraction

of

the bars or tendons to displacement

is called dowel action (see Fig. 1).

In a diagonally cracked prestressed

concrete beam, dowel action leads to a

dowel bending moment and a shear

force in the tendon itself, in addition

to the tensile force due to the effective

prestressing

force and the applied

load. As the bending moment and the

shear force due to dowel action in

crease with loading, bending and shear

Tendon

Horizontal

cracking

and

spalling

of

concrete

cover

b)

Yiel

r r

Tendon

d)

Fig

. 2. Failure modes observed

in

test beams:

a)

shear-tendon rupture failure;

b) shear-tension failure; (

c)

shear-compression failure ; d) flexural-tension failure.

75


3/12

Table 1 Experimental variables

o

test beams for first series.

Effective

Effective

prestressing Concrete

Beam Longitudinal

depth,d

, force,

F

strength,//

ideJ]tification

tendons

in. (mm) kips (kN) psi

(MPa)

Reinforcing

index,

w

Cl

CFRP 3 p

5

6

in

.

8.60 (218) 22.35 (99.4) 6450 (44.4) 0.22

(51.5 percent)

C2 CFRP 3 p

5

6 in. 8.60 (218) 22.35 (99.4) 6750 (46.5)

0.21

(51.5 percent)

Sl

Steel 3

< >

3

/s in

.

9.00 (229)

19.48 (86.6)

6150 (42.4)

0.23

(28.3 percent)

S2 Steel 3 p

3

/s in. 9.00 (229) 19.48 (86.6) 6030 (41.6)

0.23

(28.3 percent)

CS1* CFRP 3 p

5

6 in. 8.60 (218)

22.35 (99.4) 6650 (45.9) 0.21

(51.5 percent)

C3

CFRP 2

p

5

6 in .

8.45 (215) 17 .82 (79.3) 6450 (44.5)

0.30

(61.6 percent)

Steel I

p 'h

in. 13

.

71

(61.0)

(33.2 percent)

C4 CFRP 2 p

5

6 in

. 8.45 (215)

17

.82 (79.3) 6200 (42.7)

0.31

(61.6 percent)

Steel p h in.

13

.71 (61.0)

(33.2 percent)

S3 Steel 2

3

ls

in. 8.80 (224)

17.12 (76.1)

6600 (45.5)

0.27

(37 3 percent)

Steel

I p h

in.

13 .55 (60.3)

(32.8 percent)

S4

Steel 2 p

3

ls

in.

8.80 (224) 17.12 (76.1)

5950 (41.0)

0.30

(37

3

percent)

teel

I

p h

in

.

13

.55 (60.3)

(32.8 percent)

Note:

w=

p,lbd,

) /

J, )

fo

r

FRP

tendon

s; w=

Ap,lbd,

) j

pj , ) for steel tendons, wh ere AP

=

area of prestress

in

g tendons,JP"

=

ultimate strength

of

FRP tendon

s,

JPY = yield strength of steel tendons.

* Two percent stee l fibers by volume.

< > =

strand diameter

stresses initiate simultaneously in the

FRP tendon and become larger. Under

these combined ten sile and shear

stresses, the tendon may fail prema

turely, that is, before reaching its uni

directional tensile strength.

According to current research,

6

'

0

"

the available tensile strength of FRP

reinforcements decreases as their

shear stress increases. Thus, dowel ac

tion reduces the allowable ten sile

stress in the tendon beyond that al

ready

caused by

the

effective

pre-

stressing force and applied load. Also,

it may change the failure mode of a

beam from flexural-ten sion failure to

shear-tendon rupture failure, resulting

in less load carrying capacity.

Therefore , it is expected that the

premature shear-tendon rupture failure

initiated by dowel action will result in

lesser shear res

is

tance and lesser duc

tility

in

concrete members prestressed

with

FRP tendon

s .

Fig

. 2 s

how

s

schematically the shear-tendon rupture

76

failure mode in concrete beams pre

stressed with FRP tendons and other

failure modes observed in test beams

with FRP and steel tendons.

EXPERIMENT L

PROGR M

The experimental program included

two series

of

te sts (two set s

of

beam

s) . The fi rst s

eries

comprised

nine prestressed concrete beams fabri

cated without s

tirrups

. Five

beam

s

were prestressed using CFRP tendons

and, for comparison, four beams were

prestre ssed using conventional steel

tendons.

One

beam with FRP tendons

wa

s

made of fiber reinforced concrete con

taining discontinuous steel fibers . The

main objective of this first series of

tests was to experimentally confirm

the shear-tendon rupture failure mode

in

pre

stre ss

ed concrete beams

with

FRP tendons and to compare it with

other failure modes

in

prestressed con

crete beams with steel tendons.

The second series of the experimen

tal program comprised seven FRP pre

stressed concrete beams and one non

prestressed beam shear

re i

nforced

with steel stirrups (seven beams)

or

steel fibers (one beam

.

The test pa

rameters were the pretensioning ratio,

the shear span-to-depth ratio, shear re

inforcement ratio , the use of steel

fibers ,

the

compre

ssive s

trength

of

c

oncrete

, and the type of

reinfor

ce

ment. The main goal of the second se

ries was to evaluate the parameters af-

fecting the shear strength and ductility

of concrete beam

s prestre ss

ed

with

FRP tendons.

aterials

Seven-wire CFRP strands manufac

tured by Tokyo Rope Company'

2

were

used for the test beam

s.

The

5

/

16 in

.

(7 5 mm) diameter tendon had-an ef

fective section

area

of 0.047 sq in.

PC JOURNAL


4/12

Table 2. Experimen

ta

l vari ables of tes t beams for

sec

ond seri es .

I

Beam

Longitudinal

identification

tendons

C5

CFRP I

< >

1

in .

CFRP 2

5

/i 6 in

.

'

S5 Steel 3

3

/s in .

-

CS3*

CFRP 1

< >

1

in .

CFRP 2

5

/i 6 in

.

C6

CFRP I

h in

.

CFRP 2

5

/i

6

in .

-

C7

CFRP I

< >

h in .

CFRP 2

5

/i 6 in.

r -

C8

CFRP I

'h in

.

CFRP 2

5

/i

6

in

.

f

C9 CFRP I

< >

1

in

.

CFRP 2

< >

5

1i6

in .

CIO CFRP I

< >

1

in .

CFRP 2

5

/i 6 in

.

Not

e:

=

strand

di

ameter

*Two percent st

ee

l fibers

by

volume.

(30.4 mm

2

) with a specified strength,

pu

of

307 ksi

2

120

MP

a) and the

1

/z

in. (12.5 mm ) diameter tendon had an

effective section area

of

0.11 8 sq in .

(76.0 mm

2

) with a specified strength,

pu of 315 ksi (2170

MP

a). According

to th e manufacturer, the stress-strain

rela

ti

onship of the tendons is linear

elas tic up to fa

ilur

e with a tens ile

modulu s

of

19,900 and 2 1,

000

ksi

(137 and 145 GPa) with an elongation

of 1.6 a

nd

1.5 percent at rupture, re

spectively.

The steel tendons used had a diame

ter of

3

/s and

1

/z

in. (9.5 and 12

.5

mm )

and we re of Gr ade 270 ksi ( 1860

MP a) with a te ns il e modulu s of

29,000 ksi 200 GPa). No. 2 round re

inforc

in

g bars fo r the stirrups were

Grade 40 ksi (275 MPa). Type III ce

ment, natural sand, and c

ru

shed lime

stone aggregates (pea gravel in the

second se

ri

es) with a maximum size

of

3

/s

in

.

9

.5 mm) were used for the

concrete. The fibers used for the con-


Effective

Sti

rr u

ps, Effective

prestressing

spacing,

s

depth,

d

force,

F

in.

(mm

) in. (

mm

)

kips (kN)

1-leg #2

8.67 (220)

15 .3

5 (68.3)

8 (203)

(4

1.3

perce

nt

)

12.0 1 (53.4)

'

(41.5 perce

nt

)

- H:

l -

le

g #2

9.

00 (229) 27.40

1

21.9)

8 2 03) (39.8 perce

nt

)

No stirrups 8.67 (220) 15.40 (68.5 )

(41.8 pe

rc

e

nt

)

12. 18 (54.2)

(4

2.

1 perce

nt

)

-leg #2 8.67 (220) 0 (0 percent)

8 (203) 0 (0 perce

nt

)

2-leg #2 8.67 (220)

15

.

35

(68.3)

4 (10

2)

(41.3 percent)

12 .

01

(53.4)

(41.5 perce

nt

)

1-l

eg #2 8.

67

(220)

15

.

43

(68.6)

8 203) (41.5 perce

nt

)

12

.09 (5

3.

8)

(41.8 percent)

1-leg #2 8. 67 (220)

15.46 (68.8)

8 203) (41.6 percent)

12

.12 (53.9)

(41.9 percent)

1-l

eg

2

8.67 220)

15

.72 (69.9)

8 (203) (42.3 percent)

12

.

35

(54.9)

(42.7 percent)

crete were hooked steel fibers 1.18 in.

(30 mm) in length and 0.02 in. (0.5

mm) in diameter.

Test Variabl es

Experimental variables for the test

beams in the first and second series

are summarized in Tables 1 and 2, re

spectively. All beams in the first series

we re fabr ica ted without s

tirrup

s.

Beams Cl , C2, and CS 1 were pre

stressed with CFRP tendons, while

Beams Sl , S2, S3, and S4 were pre

stressed with steel tendon

s.

Combined

CFRP and steel tendons were used for

Beams C3 and C4.

In th e second series, steel stirrups

were used for a

ll

beams except Beam

CS3. Minimum shear re

in f

orcement

was provided fo r all beams

in

the sec

ond series except Beam C7, for which

the required shear reinforcement was

provided, according

to

the ACI Code.'

3

For Beams CS 1 and CS 3, hooked steel

I

Concrete

strengt

h

c'

Reinforcing

psi (MPa)

index, w

5050 3 4.8) 0.32

5650 (39.0) 0.25

5900 (40.7)

0.28

5100 3 5.2) 0.32

-

5200 (35.9)

0. 3 1

5400 (37.2)

0.

30

I

5250 (36.2)

0.30

7050 (4

8.

6)

0.23

fibers were used in the amount of 2

percent by volume of concrete.

The effecti ve prestress ratio of FRP

tendons was about 50 percent (or 60

percent) for the first series of beams

and about 40 percent of the specific

strength of the tendons for

th

e seco

nd

series except for the nonprestressed

Bea m C6. The

pr

etensioning forces

were released 4 days after cas ting the

concrete, wh en the compressive con

crete strength had reached about 70

perce nt

of

it s 28 -day strength . The

prestress losses were calculated by the

time-step method.

4

The selected shear

span-to-depth ratio wa s 2.5 for all

be ams, ex cept for 1.5 and 3.5 fo r

Beams C8 and C9, r

es

pectively.

Test Se

tup

and D

ata

Acqu isit ion

The

lo

ading arrangement and cross

secti onal dimensions (same fo r all

be ams) are shown in Fig. 3. Th e

beams were simply supported and sub-

77


5/12

p

0 0 0

8

9.5

203)

241)

1

0.75

273)

5

w

restr

ng

SSl

tendon

1

27)

43

1092)

1 (279)

ba

65 (1651)

L

Note: Dime

n

sions in

parenthesis are

in mm

Beam C8: L 65 (1651)

,

2a =

26 (660), b

=

1

9.5

(495)

Beam

C9 : L 83 (1346) ,

2a =

61 (1549), b

=

11 (279)

Fi

g

3

Load

i

ng

arrangement and typical cross section.

jected

to

one concentrated

l

oad

at

midspan. The selected shear span-to

depth ratio was 2.5, except for 1.5 and

3.5

for Beams C8 and C9, respec

tively. Fig. 4 shows

an

overall view of

the test setup and instrumentation.

A non-contacting motion measuring

instrument

(Optotrak) was used to

measure crack

displacements and

crack widths as well as load and de

flection.

This

instrument

is

a three

dimen s ional

digitizing

and

motion

analysis system. It operates by track

ing the 3-D coordinates (x,y,z)

of

ac

tive infrared emitting diodes attached

to a test specimen. For each beam, 32

markers were glued on the surface of

the beam. At the level

of

the longitudi

nal

reinforcement, markers were

placed at 2 in. (50 mm) intervals.

The test beam was loaded using dis

placement control at a loading rate

of

0.001 in . (0.025 mm) per second. For

the first series

of

tests, each beam was

loaded for one or

two

cycles

up to

about 60 or 80 percent of expected

maximum flexural load, prior to pro

ceeding

with the final loading path .

For the

second

series of

tests, each

beam was loaded monotonically up to

failure without prior loading. Continu

ous readings

of

applied load and coor

dinates

of

infrared markers were

recorded every second.

Fi

g

4. Overa

ll

view

of

test setup and instrumentation.

The following data were

obtained

by the Optotrak system: (1) load from

the load cell

of

the Instron loading ma

chine; (2) deflection at midspan; and

78

(3) crack width and differential shear

(transverse) displacement at the crack

ing plane from the markers at the level

of

the longitudinal tendons. Although

32 markers were attached to the test

beam, only the data obtained from the

markers that were closest to and on ei

ther side

of

the critical shear-cracking

plane were utilized.

N LYSIS AND DISCUSSION

OF TEST RESULTS

Relevant test results

of

the first se

ries

of

beams

are

summarized

in

Table 3.

ompar

is

on of Test R

s

ul ts

for Series I

Beams

l

and C2 vs. Sl and S

-

To

evaluate the effect

of

FRP vs.

steel tendons, the test results

of

Beams

Cl

and C2 are compared with those

of

Beams Sl and S2. Beams Cl and C2

were

prestressed

with

FRP tendons

,

while Beams S 1 and S2

were

pre

stressed with steel tendons.

A marked difference

between

the

test beams was their mode

of

failure.

Beams Cl

and

C2

failed by

shear

tendon rupture, while Beams S 1 and

S2 failed by shear-compression. As

P IJOU RNAL


6/12

Table 3. Summary

of

shear test results for first series.

Beam

Failure

Pu

ll

I

jSll

Wsu

P scr

P,

11

ACI)

I

PJ P

Pfu AC I)

identification mode

kips kN) in. mm ) in.

mm

) in

mm )

kips

kN)

kips kN)

ACI)

kips kN)

Cl STR 41.7 185) 0.24 6. 1)

0.030 0.76) 0.016 0.41) 32.6 145) 25.7 114)

l.62

44.4 197)

C2 STR

43.7 194) 0.29

7.4)

0.027 0.69) 0.024 0.61) 28.5 li S)

25.9

li S)

1.69 44.7 199)

Sl sc

49.4 219) 0.35 8.9)

0.092 2.34) 0.032 0.8 1)

27.3 121) 24.7 1

10)

2.00 48.4

2

1

5)

r

S2

sc

47.5

21

1) 0.38 9.7) 0.096 2.44) 0.046 1. 17) 27.7 123)

24.6 109) 1.93

48.2 2 14)

CSI Ff -

22 1) 5.3)

0.

009 0.23) 0.0 I0 0.25) 34.0 IS I) 25.8

li S)

44.6 198)

1-

STR

+-

49.8

Z22

l 0.23 5.8)

3

0.034 0.86) 0.026 0.66) 35.5 I57) 3 1.1 138)

1.

60

55

.7 248)

I

-

C4 STR

51

.0 227) 0.24 6

.1

)

0.037 0.94) 0.023 0.58)

38.6 172)

30.9 137)

1.65

I

55.2 246)

S3

ST

57.0 254) 0.27 6.9)

N/A

N/A 38.0 169)

32.1 143)

1.78 57.3 255)

S4 ST

54.1 241)

0.28 7.1)

N/A

N/A

I

40.0 178) 3

1.

6 141)

1.71 56.1 250)

ole:

STR : Shear-tendon rupture failure;

Ff: Fl

exura l-tension failure: SC: Shear-co mpress ion fa ilure:

ST

: Shear-tension failure.

P

= ultimate load

6

= ultimate deflec tion

Du= ult imate shear djspl acement

Wsu =ultimate shear crack width

P

a

=s hear cra

ck

ing load

P. AC I) =design she

ar

strength us

in

g ACI Code

P

1

ACI) = des ign flex ural strength using ACI

Co

de

Fig. 5 Shear-tendon rupture failure and crack pattern

of

Beam C1 .

Fig. 6 Shear-compression failure and crack pattern

of

Beam 51

shown

in

Figs.

5

and

6,

Beam CJ was

split into two segments

by

the tendon

rupture

at the critical shear-crack

plane, while Beam S I remained to

gether.

Also, Beam C1 had smooth

failure faces , while Beam S1 had the

concrete crushed

in

the compression

zone

and

spalled

off

in the

tension

zone. The angle of the critical shear

cracking was about

50

to 55 degrees

in

Beams C

I

and

C2

, and about

45

de

grees in Beams S

I

and S2.

Different types of failure led to dif

ferent shear resisting

capacities

. On

average , Beams C 1 and

C2

, which

failed by shear-tendon

rupture,

had

about 12 percent less shear carrying

capacity than Beams S

I

and S2, which

failed by shear-compression in the

concrete. Also, the

average

ultimate

deflection of Beams C 1 and C2 at


midspan was about

30

percent less

than that of Beams S

1

and S2.

For the beams that failed by shear,

the measured ultimate loads were con

siderably higher than the design shear

strength computed using

the

ACI

Code. Beams C

1

and C2 had about

65

percent higher ultimate shear

strengths, while Beams S I and S2 had

about

95

percent higher ultimate shear

strengths.

It can be seen from Figs.

7, 8,

and

9

that the general

shapes of the

load

deflection response

of

Beams

Cl

and

S

I

their load-shear displacement, and

their load-shear crack width are very

similar. However, the values

of

loads,

deflections , shear displacements and

shear crack widths at ultimate were

significantly different. The response

curves

of Beams

C2 and S2 are not

shown

in

the

figures

because they

were very similar to those of Beams

Cl

and

Sl

, and to maintain the clarity

of

the figures.

A

notable

difference between the

two types

of

reinforcement

FRP

vs.

steel)

is in

the maximum vertical shear

displacement at the critical shear

cracking

plane. As can be

seen

in

Table

3

and Figs.

8

and

10,

the aver

age differential shear displacement of

Beams C

I

and C2 at ultimate load was

about 30 percent

of

that

of

Beams S1

and S2.

Also, the average

crack

width of

Beams

Cl

and

C2

at failure load was

about 50 percent of that

of

Beams S1

and S2. Moreover , as shown

in

Fig .

10, the relationship between the shear

displacement and crack width was al

most linear for all beams.

t

is

also ob-

79


7/12

Shear crack

width

mm)

0

2

Deflection mm)

4 6 8

10

12

0 0.2 0.4 0.6 0.8

1

2

60

r r ~ ~ ~ r r ~ ~ r r ~ ~ ~ ~ ~ r ~

250

250

50

Beam

CSl .. Beam Sl

50

Beam

CS1

. Be

am

81

......

200

_40

'

Beam

C1 solid)

0 .

g3

' '

'

j

20

10

50

....._

40

'

.

30

' '

'

j

20

10

Beam

q

200

.

50

0

L ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ o

o ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ o

0

0.1

0.2 0.3

0.4

0.5

0 0.01 0.02 0.03 0.04 0.05

Deflection in)

Shear

crack width in)

Fig. 7. Load-deflection curves for Beams

C1

51 and

CS1. Fig. 9.

Load-shear crack width curv

es

for

Beams

C1

51

and CSl.

Shear

displacement

mm)

0 0.5 1 1.5 2 2.5 3 3.5

6 0 r r r n r M . ~ O T . T r r r r n r M ~ . ~ ~ ~ ~

50

....._ 40

'

.

d 3o

' '

'

20

Bea m CS1

Beam

Sl

........

Beam

C

_

..-

---

Shear displacement in)

0 .12

0.1

200

c 0.08

Q)

E

Q)

0.

06

c

'

;

0.04

....

'

0

0 .02

Cll

0

Shear

crack width mm)

0 0.2 0.4 0.6 0.8 1 1.2 1.4

3

Beam 81

2.5 6

s

2

c

Q)

E

1.5

'

c

1

'

;

:

0.5

..c:

Cll

0

0 0.01 0.02 0.03 0.

04

0.05 0.06

Shear

crack width in)

Fig. 8. Load-shear displacement curves for Beams C1

51

and CS1 .

Fig. 10. Shear displacement-crack width curves for Beams C1

51

and CSl

served that Beam CS 1 had a small

er

slope i.e., stiffer response) than that

of Beams C and S1

Beams l and C2

vs. Sl - In

order to assess the effects of adding

fibers to the concrete matrix, the test

results

of

Beams

Cl

and C2 are com

pared to the results of B

eam

CS 1,

which was made of fiber reinforced

concrete.

As mentioned earlier, Beams C1

and C2 failed by shear-tendon rupture,

whi le

Beam

CSl failed by flexural

tension Figs. 5 and 11). The addition

of

fibers changed the failure mode and

led

to

smaller crack widths and a

larger number

of

cracks . The reduc

tion in crack width led to a reduction

in

differential shear displacement,

which changed the failure mode from

80

shear -tendon rupture to flexural

tension failure. Also, the load carrying

capacity of Beam CS1 was 15 percent

larger than the average load carrying

capacity of Beams C1 and C2, and its

deflection at ultimate was 20 percent

smaller.

Moreover, due to the effects of the

fi

bers,

Beam

CS 1 was

co

nsiderably

stiffer than Beams C1 and C2 see Fig.

7) and its ultimate differential shear

displacement and crack width were al

most one-third and one-half the aver

age values of Beams C1 and C2, re

spectively see Figs . 8, 9 , and 10).

Unlike beams with

FRP

tendons that

failed by flexural mode, Beam CS 1

did not experience the very loud bang

and had no large longitudinal cracks.

This fact is attributed to the effects of

fibers, including higher bond strength,

high

er

co n

fi

nin g forces , and higher

fracture toughness

resulting

in a

higher capacity

of

energy absorption

at FRP tendon rupture.

Beams C3 and C4

vs.

S3 and S4

- Beams C3 and C4 were prestressed

with two RP tendons and one steel

tendon, while Beams S3 and S4 were

prestressed with three steel tendons.

They had about the same prestressing

index.

The failure modes

of

these beams

were significantly different. Beams C3

and C4 failed by shear-tendon rupture

see Fig. 12), while Beams S3 and S4

failed by shear-tension see Fig. 13).

The latter

is

characterized by splitting

debondi

ng)

alo

ng

the tension

rein

forcement at the end

of

a diagonal ten-

PCI JOURNAL


8/12

Fig 11.

Flexural tension

failure and crack

pattern

o

Beam CS1

Fig

12.

Shear tendon

rupture failure and

crack pattern

o

Beam C4.

Fig 13.

Shear tension failure

and crack pattern o

B

ea

m 54.

sion crack. The shear-tension failure

in

Beams

S3

and S4 also had a differ

ent failure pattern from

the

shear-

compression failure in Beams S 1 and

S2, which could be explained by their

higher prestressing force or prestress

ing index).

As can be seen in Fig . 14, the load

deflection curves

of

Beams C4 and S4

are very similar up to the

ultimate

loads. Moreover, there is no distinct

difference

in

the crack pattern except

that

the

diagonal

tension

cracks

of

Beams S3 and S4 occurred at the final


stage

of

loading with relatively low

angles of about 30 to 35 degrees and

led to failure .

On the other hand, the critical shear

cracks in Beams C3 and C4, which de

veloped from the early stages of load

ing with angles of 50 and 40 degrees,

respectively, caused the failure. The

average ultimate deflection of Beams

C3 and C4 was slightly smaller than

that

of

Beams S3 and S4. On average,

Beams C3 and C4 had about 10 per

cent less shear resistance than that

of

Beams S3 and S4.

Test esults of Test

Series

II

Table 4

presents

the

summary

of

relevant test results for the second set

of beams, and Figs. 15 16,

17

and 18

show their load-deflection, load-shear

displacement, load-shear crack width,

and shear displacement-crack width

curves, respectively.

omparison of Test esults

for Series II

eams CS and

SS - Beam S was

prestressed with FRP tendons, while

Beam S5 was prestressed with steel

tendons. The effective

prestressing

ratio

in

the FRP tendons was about 40

percent. Both beams failed by shear

tension failure;

however

, Beam CS

had about

13

percent less shear resist

ing capacity than Beam SS see Table

4). It appeared here that a reduction in

the

effective prestressing

ratio

changed the shear failure mode in

Beam

S

from shear-tendon rupture to

shear-tension. The measured ultimate

load of Beams

S

and SS was, respec

tively , 23 and 40 percent higher than

the design shear strength calculated

from the ACI Code.

The shear-tension failure that oc

curred

in

Beams C5 and

SS

was sud

den and explosive. The concrete cover

of the test beams suddenly cracked

and spalled off along the longitudinal

tendon s. Their crack patterns were

very similar except that the angle of

the critical shear crack of Beam S

about 45 degrees) was slightly steeper

than that

of

Beam about

40

degrees).

The general shape of the load-

deflection curves see Fig . 15) and

load-shear displacement curves see

Fig. 16) of Beams

S

and

SS

are simi

lar. At the failure load, Beams S and

SS

had about the same ultimate deflec

tion and shear displacement. The dif

ference in their ultimate shear strength

can be attributed to the difference

in

the elastic moduli of the tendons. Steel

tendons

would have higher tensile

forces than FRP tendons at the same

deflection. Beam SS showed about the

same response as Beam S until shear

cracking load , and higher stiffness

after shear

cracking

. The ultimate

shear crack width of Beam S at the

failure

shear plane

was

about

two

times that of Beam SS The relation-

81


9/12

Fig . 14.

Comparison o load

deflection curves for

Beams C4 and 54.

50

2

0

0.1

ship between the shear displacement

and crack width was almost linear for

the two beams (see Fig. 18).

Beams

CS, C7,

and

CS3 -

Beams

C5, C7, and CS3 were identically pre

stressed with three FRP tendons. For

Beams C5 and C7 , respectively, mini

mum and required shear reinforce

ments

according to

the

ACI Code

were provided with steel stirrups. For

Beam CS3, 2 percent by volume of

hooked steel fibers was added to the

concrete matrix as a substitute

for

steel stirrups.

There was a marked difference be

tween the modes

of

failure

of

these

beams. Beams C7 and CS3 failed by

shear-tendon rupture, while Beam C5

failed by shear-tension as mentioned

earlier. Increasing the amount

of

stir

rups or adding steel fibers, increased

Deflection (mm)

4 6 8

10

. ..

Beam

S4

Beam

C4 (solid)

0.2 0.3

0.4

Deflection (in)

12

250

200

1soz-

c

' 0

100

g

:1

50

0

0.5

the ultimate load but changed the fail

ure mode from shear-tension failure to

shear-tendon rupture failure. The FRP

tendons

in

Beams

C7

and CS3

snapped at a flexural-shear-cracking

plane. All three tendons in Beam CS3

ruptured simultaneously, while two

tendons

in

the lower row of Beam C7

ruptured at failure.

The angle

of

the failure flexural

shear plane was about 65 degrees in

Beam C7 and about

60

degrees in

Beam CS3. The fail ure plane

of

Beam

C7 initiated at the bottom location of

the first

steel

stirrup nearest to the

loading point. The addition

of

steel

fibers in Beam CS3 led to smaller

crack widths and a larger number of

cracks . The flexural crack width in

Beam CS3 was too small to be mea

sured even at the ultimate load.

Table 4. Summary o shear test results for second series.

Beam

Fa

ilure

P. 6.

6,.

w .

identification mode

kips (kN) in. (mm) in. (mm)

in

. mm )

cs

ST

41.9 (186)

0.35 (8.9) 0.108 (2.74) 0.066 (1.68)

ss

ST 48 .3 (215) 0.32 (8.1) 0.097 (2.46) 0.036 (0.

91

)

CS3 STR

50.6 (225) 0.

27

(6.9) 0.036 (0.

91

) 0.043 ( 1.09)

C6

ST

15 .8 (70) 0.55 (14.0)

- -

C7

STR

47 .1 (210) 0.43 (10 .9) 0.024 (0.

61

) 0.023 (0.58)

C8

cs

57

.7 (

25

7) 0.

12

(3.0)

-

-

C9

ST

29.9 (133) 0.54 (13.7)

-

-

C lO ST

44.4 (197) 0.48(

12.

2) 0.079 (2.00) 0.04 1 (1.04)

Note: STR: Shear-tendon rupture failure; ST: Shear-tension frulure ; CS: Compress

iOn-

strut fru.lure.

Pu=ultimate load

, ; ultimate deflection

su

=ultimat

e s hear displacement

w su =ultimate shear crack width

P

; shear cracking load

P

'

ACD; design shear strength using ACI Code

P i

(ACI) ; design

fl

exural strength using ACI Code

82

Ps

cr

kips (kN)

28 (125 )

30 (133)

32

1

42)

10 (44)

26 (116)

52 (

231

)

20 (89)

28 (125)

As can be seen in Table 4, the ulti

mate shear resisting capacity of Beam

C7 was 12 percent larger than that of

Beam C5. The load-deflection re

sponse of Beam C7 was very similar

to that

of

Beam C5 except that Beam

C7

has slightly higher stiffness after

shear cracking and larger deflection

(about 25 percent) at the ultimate load

(see Fig. 15) . At the flexural-shear

plane, the ultimate shear displacement

of Beam C7 was about one-quarter

and the corresponding

crack

width

was about one-third that of Beam C5 .

Beam CS3, which failed by shear

tendon rupture, had about 20 percent

larger shear resisting capacity than

Beam C5, which failed by shear ten

sion. Due to the effects of steel fibers,

Beam CS3 was considerably stiffer

than Beam

C5

from the beginning to

the failure load, with a steady continu

ous change

in

curvature in the load

deflection curve after shear cracking

(see Fig. 15). The ultimate deflection

of

Beam CS3 was about

75

percent

of

that of Beam C5 (see Table 4). The ul

timate shear displacement and corre

sponding shear

crack

width

at

the

shear-cracking plane were one-third

and two-thirds of that of Beam C5,

respectively .

The reason why the tendon-rupture

of

Beams C7 and CS3 occurred at a

shear displacement smaller than that

of Beam C5 is thought to be due to the

strengthened

concrete

cover due to

hook action of the steel stirrups and

steel fibers. The strengthened concrete

P. ACn

P.fP

. Pfu (ACI)

kips (kN) (ACI)

ki ps (kN)

34.0 (1

51

) 1.23

48.9 (

21

8)

34 .6 (154) 1.40 47.7 (212 )

33.8 (150)

1.50

50.5 (225)

25.0 (

111

)

1.02 49.0 (218)

46.7 (208) 1.01

49.2 (219)

38

.6 (172)

1.49

86.2 (383)

23

.2 (103 ) 1.29

34.6 (154)

34.8 (155 )

1.28

52.1 (232)

PCI JOURN L


10/12

Deflect ion (mm)

0 6 8

10 12

14 16

70

STR: Sh

ear

tendo n rupture

60

C8

(CS)

ST: Shea r -tension failure

CS:

Co

mpression

s

trut failure

50

CS3 (STR)

55

(S I)

Cii 40

;g

al

30

j

20

C6 (S I)

10

300

60

250

50

200

40

z

~

;;

0.

30

100

.

j 20

50

10

0 0.5

Shea r crack widt h (

)

1 1.5

CS3 (

STR)

2 5

250

200

150

z-

c

STR: Shea r-te

nd

on r

uptur

e

ST: Shear -tension failu re 100 j

50

0

0

0

~ ~ ~ ~ ~

0

0. 1 0. 2

0.3 0.4

0.5 0.6

0.7

0 0.02

0.04 0.06

0.08 0. 1

De flection (in)

Shea r crack width (in)

F

ig.

1

5.

Load-deflec

ti

on cu

rves

of second set of beams. Fig . 17. Load

-s

hear crack width curves of seco

nd

set of beams.

60

50

40

;;

Q.

30

j

20

10

0

0 0.5

CS3 (STR)

Shear di splaceme

nt

(mm)

1.5 2 2 .5 3

S5 (S I)

ST

R:

Sh

ea r

-tendon ru pt ure

ST: Sh

ear

-tension failure

3.5

250

200

150z

c

'

00j

50

~ ~

~

0

0.02 0.04 0.06

0.08 0.1 0.12 0.14

She

ar displace ment (in)

Shea r crack width (mm )

0 0.5 1 1.5 2.5

0.

14

3. 5

0.12

C5 (S I)

3

5

0. 1

2.5

c

0 .08

2

E

1J

1J

.,

STR: Shea r-tendon r upture

0.06

ST:

Sh

ear-te ns ion failur e

5 g.

:0

:0

:;

0.04

1

c::

eli

n

0.02

0.5

0

0

0

0.02

0.

04 0.0 6

0 .08

0.1

Shear

c

ra

ck

width

(in)

Fig. 16. Load-shear displace ment curves of second se t of

beams.

Fig. 18. Shear di splacement-c rack width c

urv

es of seco

nd

set

of beam

s.

cover prevents

th

e opening

of

a hori

zontal crack along the lon

gi

tudinal re

info rcement , res ultin g in increased

dowel fo rce on CFRP tendons.

As a result of

th

e premature shear

tendon rupture, the ultimate shear re

si

st

in

g capacity

of

Beam C7, which re

quired stirrups according to the ACI

Code, was considerably reduced and

was about 7 percent less th an that of

Beam CS3, which conta

in

ed 2 percent

hooked stee l f ibers without stirrups

(see Tabl e 4). In both cases, the beams

fai led by shear-tendon rupture at the

fl

ex ural-shear-cracking plane. At the

ul

timate load, Beam CS3 had about a

35 percent smaller ultimate de

fl

ec

ti

on.

As shown in Figs. 16 and 17 , Beams

C7 and CS3 sh

ow

a ve ry st

ee

p r

e-

sponse in the

ir

load-shear displacement

January-

February 1999

and load- shear crack width c

ur

ves .

Like other beams, Beams C7 and CS3

al so showed

an

almost

Lin

ear relation

ship betwee n th e shear displ ace ment

and crack width (see

Fig.

18).

Beams CS and C6 - To evaluate

the effects

of

prestress

in

g on the shear

performance, Beam C6 was fabricated

with FRP tend ons with no pres tress,

i.e., similar to rein forced concrete.

Beam C6 failed by shea r-tension,

which was the same fa ilure mode as

Beam C5 . However, the ultimate shear

strength of Bea m C6 was about

40

percent of th at

of

Beam C5 (see Table

4). Thi s suggests that increas in g the

prestress ing force is a possible way to

in

crease the shear resisting capac

it

y

of

concrete beams prestressed with FRP

and steel tendons. The angle of the

failure shea r pl ane of Beam C6 was

about 35 degrees, while that of Beam

C5 was 45 degrees. As shown

in

Fi

g.

15 , Beam C6 has much lower stiffness

and about 60 perce nt large r ultimate

deflection than Beam C5.

Beams

CS,

C8 and

C9

- To eval

uate the effects of shear span-to-dep th

ratio, Beams C8 and C9 were tes ted

with a shear span-to-depth ratio of 1.5

and 3

.5

, respect

iv

ely, and compared to

Beam C5, which was tested at a shear

span-to-depth ratio of

2.

5.

Beam C8 failed by crushing of

th

e

compression strut and Beam C9 failed

by shear-tension, which was th

e same

failure mode as Beam CS For Bea m

C8 with a shear span-to-depth ra

ti

o of

1.5 , the applied l

oa

d seemed to be car

ried mainly by the compress ion strut

83


11/12

connecting the loading point to the

support, at

an

angle of 35 degrees. For

Beam C9 with a shear span-to-depth

ratio of 3.5, the failure crack angle

was about

25

degrees, which is much

smaller than that

of

Beam

C5

(about

45 degrees).

The ultimate shear resisting capaci

ties

of

Beams C8 and C9 were about

40 percent higher and about 30 per

cent lower than that

of

Beam C5, re

spectively. At the ultimate load, the

deflections of Beams C8 and C9 were

about 65 percent smaller and about 55

percent larger than that of Beam C5 ,

respectively. Thus, it is concluded that

the beams with a larger shear span-to

depth ratio have less shear resisting

capacity and larger shear ductility.

Beams CS and ClO -

To evaluate

the effects of concrete compressive

strength Beam

C 10 was made of

higher strength concrete than Beam

C5

, with.fc = 7050 psi (48.6 MPa.

The

test results showed

that

the

compressive strength

of

concrete had

no significant effect on the ultimate

shear strength

of

concrete beams with

FRP tendons but had a significant ef

fect on their ultimate deflection at low

prestressing ratios. As shown in Fig.

15, the load-deflection curve

of

Beam

ClO was similar to that

of

Beam C5 ,

except for a 6 percent higher shear

strength and a 35 percent larger de

flection at ultimate.

A relatively soft snapping sound

was heard prior to complete failure in

Beam ClO. Examination of Beam ClO

after failure revealed that three wires

in the

z

in. (12.5 mm) FRP tendon

and three wires in the h in. (7.5 mm)

FRP tendons were broken at the fail-

84

ure shear plane leading to the shear

tension failure observed. The failure

shear crack angle of Beam C I 0 was

about

35

degrees, while that of Beam

C5

was about

45

degrees.

CONCLUSIONS

On the basis of this experimental in

vestigation, the following conclusions

can be drawn:

1 The shear-tendon rupture failure

is a unique mode of failure, which, un

less properly designed for, is likely to

occur

in

concrete beams prestressed

with FRP tendons . This premature

failure is due to tendon rupture by

dowel

shear at the shear-cracking

plane.

t

is attributed to the poor resis

tance of FRP tendons

in

the transverse

direction and their brittle behavior.

2 The ultimate shear resisting ca

pacity of beams prestressed with FRP

tendons was about

15

percent less than

that

of

beams prestressed with steel

tendons, regardless

of

their shear fail

ure mode.

3. The shear-tendon rupture failure

occurred at the flexural-shear-cracking

plane in beams with FRP tendons,

even when the effective prestress ratio

was low (about 40 percent) and there

quired amount

of

steel stirrups was

provided according to the ACI Code.

4 Adding steel fibers is a possible

way to improve the shear resistance

of

concrete beams prestressed with FRP

tendons by avoiding or delaying shear

tendon rupture failure.

5. Differences in the properties of

FRP and steel tendons appear to have

no significant effect on the initial por

tion of load-deflection response

of

prestressed concrete beams subjected

to a center point loading with a shear

span-to-depth ratio of 2.5.

6 The ultimate shear displacement

and crack width

of

prestressed beams

that failed by shear-tendon rupture

were about one-third and one-half, re

spectively,

of

those

of

similar beams

with steel tendons. For all beams

tested,

an

almost linear relationship

was observed between the shear crack

width and the differential shear dis

placement at the critical shear-crack

ing plane.

7. Although only one specimen was

tested for each parameter, the follow

ing observations were made for beams

prestressed with

FRP tendons:

Increasing the shear span-to-depth

ratio from 1.5 to 3.5 led to a de

crease

in

shear resistance but an in

crease in shear ductility (displace

ment).

Adding stirrups in sufficient quan

tity changes the failure mode from

shear-tension to shear-tendon rup

ture in beams with a low effective

prestress ratio of about 40 percent.

Increasing the compressive strength

of concrete slightly increases the

shear strength and considerably in

creases the corresponding deflection.

CKNOWLEDGMENTS

This research was supported in part

by the Department

of

Civil and Envi

ronmental Engineering at the Univer

sity of Michigan. The authors are also

grateful to Tokyo Rope Manufacturing

for supplying the carbon fiber rein

forced plastic

strands

used in this

study.

PCI JOURN L


12/12

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Documents

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