Bond Clause Proposal for FRP

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Bond clause proposal for FRP-bars/rods inconcrete based on CEB/FIP Model Code 90with discussion of needed tests

Ralejs Tepfers

Department of Structural Engineering and Mechanics Report 04:2

Concrete Structures

CHALMERS UNIVERSITY OF TECHNOLOGYGöteborg, Sweden

Göteborg 2004





REPORT 04:2

Bond clause proposal for FRP-bars/rods in concrete based on CEB/FIP Model Code 90 with discussion of

needed tests

RALEJS TEPFERS

Department of Structural Engineering and MechanicsConcrete Structures

CHALMERS UNIVERSITY OF TECHNOLOGY

Göteborg, Sweden 2004



Bond clause proposal for FRP-bars/rods in concrete based on CEB/FIP Model Code90 with discussion of needed tests

© RALEJS TEPFERS, 2004

ISSN 1651-9035

Report no. 04:2Archive no. 120Department of Structural Engineering and MechanicsConcrete Structures

Chalmers University of TechnologySE-412 96 Göteborg

SwedenTelephone: + 46 (0)31-772 1000

Department of Structural Engineering and Mechanics

Göteborg, Sweden 2004



I

Bond clause proposal for FRP-bars/rods in concrete based on CEB/FIP Model Code90 with discussion of needed tests

RALEJS TEPFERS

Department of Structural EngineeringConcrete Structures Chalmers University of Technology

ABSTRACT

The bond clauses in the CEB/FIP Model Code 1990, (MC90), were written for steelreinforcement in concrete. After finalization of the MC90 the further work withinCEB and after the merger with FIP in the new organization fib should be directedtowards covering up lacunae and also adapting the Model Code to new materials. The

basic MC90 concepts should be kept when applied to new materials. Models should

be used to justify the clauses. However, the bond clauses in MC90 were not fully based on models, because there are very many options for bond failures and it isdifficult to cover up these with few models. Reinforcing steel is more or less just onematerial. FRPs, on the other hand, are made of a lot of different materials withmanifold shapes and all these FRPs perform differently in many contexts. Especiallyin severe environments, where steel is not durable, FRPs may perform well. To makeuse of the fiber composites, it is necessary to bring these materials to codes of practiceand the closest is to try to adapt the code for steel reinforced concrete to FRPreinforced. The clauses should be given with boxed values open for individualcoefficients for different FRPs. The necessary coefficients should be determined forthe code models using systems of appropriate and coupled test methods.

Key words: Concrete, Reinforcement, FRP, Bond, Code Clauses.



II

Normförslag för vidhäftning av FRP armering i betong baserad på CEB/FIP ModelCode 90 med diskussion av nödvändiga tester

RALEJS TEPFERS

Institutionen för konstruktionsteknikBetongbyggnad

Chalmers tekniska högskola

SAMMANFATTNING

Vidhäftningsnormen i CEB/FIP Model Code 1990 (MC90) har skrivits förstålarmering. Sedan MC90 hade fullbordats avsågs det fortsatta arbetet inom CEB ochefter sammanslagningen med FIP till den nya organisationen fib att täcka uppobserverade brister i MC90 och att även anpassa Model Code till nya material. Detgrundläggande upplägget i normen skulle bibehållas när den anpassades till nyamaterial. Modeller skulle användas för att belägga normkraven. Emellertid varvidhäftningsnormen i MC90 inte helt modellbaserad, då det finns alltför mångavarianter på vidhäftningsbrott och det är mycket svårt att täcka upp alla med fåmodeller. Stålarmering är mer eller mindre ett material. FRP å andra sidan tillverkasav en mängd olika material, med många utformningar och alla dessa uppför sig olika ien mängd sammanhang. Särskilt i besvärliga miljöer, där stål inte är beständigt, kanFRP armering klara sig mycket bra. För att kunna dra nytta av fiberkompositer måstede introduceras i normer och det som ligger närmast är att anpassastålarmeringsnormen till FRP armering. Normkraven för vidhäftning bör ges med

boxade värden öppna för olika typer av FRP armering. De nödvändiga koefficienternaför medellerna i normen bör bestämmas med system av adekvata och koppladetestmetoder.

Nyckelord: Betong, Armering, FRP, Vidhäftning, Normer



CHALMERS , Structural Engineering and Mechanics, report 04:2 III

Contents

ABSTRACT I

SAMMANFATTNING II

CONTENTS III

PREFACE V

1 INTRODUCTION 1

2 MODEL CODE 90 BOND CLAUSES ADAPTED TO FRP-REINFOR-

CEMENT AND RELATED TEST METHODS WITH COMMENTS 3

2.1 Non prestressed reinforcement 3

2.1.1 Introduction 3

2.1.2 Design bond stress for reinforcing bars 3

2.1.2.1 Code clauses 3

2.1.2.2 Necessary tests for determination of ?-values 4

2.2.3 Code clauses 14

2.2.3.1 Basic anchorage length 14

2.2.3.2 Design anchorage length 15

2.2.3.3 Test method for determination of influence of transverse pressure 19

2.2.3.4 Design lap length of bars in tension 19

2.2.3.5 Design lap length of bars permanently in compression 20

2.2 Prestressed concrete 21

2.2.1 Anchorage of prestressing tendons 21

2.2.2 Anchorage of pretensioned prestressing reinforcement 21

2.2.3 Design bond strength 22

2.2.4 Basic anchorage length 22



CHALMERS , Structural Engineering and Mechanics, report 04:2IV

2.2.5 Transmission length 23

2.2.6 Determination in test of the transmission length 24

2.2.7 Design anchorage length 25

2.2.8 Development length 252.2.9 Transverse stresses in the anchorage zone of prestressed tendons 26

2.3 Tension stiffening effect 26

2.4 Rotation capacity 26

3 REFERENCES 27



CHALMERS , Structural Engineering and Mechanics, report 04:2 V

Preface

This report is elaborated with the aim to be used for the bond clause work within the “fédérationinternationale du béton” fib, Task Group 4.5 “Bond Models”.

The paper is meant to give the necessary background for elaboration of bond code clauses forFRP reinforcement in concrete based on CEB/FIP-s Model Code 1990. The needed coefficientsin the formulas have to be determined by tests. Adequate test methods have to be agreed upon.Here test methods are proposed, but these could still be improved.

Chalmers University of Technology and the company “Ralejs Tepfers Consulting” havesponsored the work. The given support is gratefully recognized.

Göteborg, December 2003

Ralejs Tepfers





CHALMERS , Structural Engineering and Mechanics, report 04:2 1

1 INTRODUCTION

The bond clauses in the CEB/FIP Model Code 1990, (MC90), were written for steelreinforcement in concrete. After finalization of the MC90 the further work within

CEB and after the merger with FIP in the new organization fib should be directedtowards covering up lacunae and also adapting the Model Code to new materials. The basic MC90 concepts should be kept when applied to new materials. Models should be used to justify the clauses. However, the bond clauses in MC90 were not fully based on models, because there are very many options for bond failures and it isdifficult to cover up these with few models. It is a comprehensive task to developclauses based on models and it requires some years of work of a task group doing alsoadditional research work. Such work has been started by the new fib, Task Group 4.5“Bond Models” for the next generation of Model Code. Here is presented what can bedone for time being.

The bond of steel reinforcement in concrete has been investigated for a century andstill there is a lot to be done. Reinforcing steel is more or less just one material and itis practically the same world over. FRPs, on the other hand, are new materials, a lot ofdifferent materials with manifold shapes and all these FRPs perform differently inmany contexts. Especially in severe environments, where steel is not durable, FRPsmay perform well. To make use of the fiber composites, it is necessary to bring thesematerials to codes of practice and the closest is to try to adapt the code for steelreinforced concrete to FRP reinforced. The clauses should be given with boxed valuesopen for individual coefficients for different FRPs. The necessary coefficients should

be determined for the code models using systems of appropriate and coupled testmethods.

Bond of ordinary steel reinforcement in concrete is dependent on very many parameters as pull-out resistance, geometry of the concrete member, placing of the barin the member section, cover splitting, confinement by concrete and surroundingreinforcement, the order of bond crack appearance and bond stress distribution alongthe bond length. The bond of the very many existing types of FRP reinforcement isdependent of even more parameters. The surface of the FRP bars is weaker and softerthan that of steel bars and may fracture instead of concrete and does not create in bondcontact points to concrete as high local stress concentrations as the harder steel barsdo. FRP-bars fixed with epoxy paste or mortar in grooves for strengthening have twointerfaces, bar to mortar and mortar to concrete, which increase the number

parameters and introduce new possible failure modes. The fundament for bondresistance estimation should be an accepted bond philosophy linked to appropriatemodels. A system of bond tests should provide necessary coefficients for the modelsin clauses.

The material compositions, forms and shapes of FRP reinforcement are too early tostandardize, because the development has not matured and arrived to a few materials,appropriate shapes, dimensions, surface designs, stress-strain relations, E-modulus,fatigue resistances and bond performances. It is not yet clear in which way theappropriate bond should be best achieved. If the FRP surfaces should be sand covered,have lugs, wavy deformations or combinations of what ever. Certain bond function

should also be ensured when glass transition temperature Tg of the resin is exceededor during a fire. May be a wavy surface configuration is to prefer, which force the



CHALMERS , Structural Engineering and Mechanics report 04:22

more fire resistant fibers to take part in bond action, if the viscosity of the resindecreases.

FRP reinforcement can also be used for flexural and/or shear strengthening of existingmembers. For this application, FRP-bars are placed in grooves cut on the surface of

the member to be strengthened and there fixed with cement mortar or epoxy paste. Insuch application, the performance of bond between the FRP rod and the mortar orresin and then between mortar or resin and concrete is critical for the effectiveness ofthe technique. The presence of two interfaces increases the number parameters neededto characterize the global “joint” behaviour and introduce new possible failure modes.

In research it is necessary to tell what kind of material has been tested, who hasmanufactured it and when. The FRP materials change, improve and alter in

performance. Therefore dating is important. Old results cannot directly be applied toimproved FRP materials. If we in our investigations do not tell what kind of materialwe test and just call the materials A, B and C, then in due time this knowledge will be

lost. It should not be like this. Also the manufacturers will profit of openness.

The MC90 clauses in the following will be adapted as far as possible to function ofinternal FRP-reinforcement. Necessary test methods will be proposed for estimationof the coefficients in formulas.




2 MODEL CODE 90 BOND CLAUSES

ADAPTED TO FRP-REINFORCEMENT AND

RELATED TEST METHODS WITH

COMMENTS

2.1 Non prestressed reinforcement

2.1.1 Introduction

The FRPs are elastic up to tensile failure while hot rolled deformed steelreinforcement yields. The anchorages should provide necessary bond for a load ofabout 10% above the capacity of bars in the concrete member. For steel reinforcementthe yielding leads to successive break down of anchorage and much above the yield

strength cannot be accounted for. The bond in MC90 clauses is not enough to enablethe steel bars to fail in tension. FRPs have no defined yield stress. The existing FRPreinforcements for time being have 30-80% lower modulus of elasticity than steel. Forordinary use in concrete member of non-prestressed reinforcement the deformationsof the structure will in most cases limit the stresses in FRP reinforcement to 30-80%of those in steel reinforced concrete structures. There should be an accepted level ofsafety of tensile failure for FRP bars and a design stress. An agreement has to bereached about how much of the tensile strength of the FRP bar should be anchored at

bond failure. Or, if stress 10% above the tensile design strength of FRP bar/rod should be anchored.

In the following the formulas are numbered according to the MC90.

2.1.2 Design bond stress for reinforcing bars

2.1.2.1 Code clauses

The coefficients in the code should be given as open boxed values, which aredetermined in tests for each type of FRP bar/rod. [ ] means boxed value.

f bd = 321 ηηη k D k T f ctd < f bd FRP; ………………(1)

where

f ctd is the design value of concrete tensile strength (= f ctk,min/1.50)

f bd FRP is the design bond strength in the surface of the FRP-bar/rod (determined in pull-out test with short bond length and central placement of bar/rod).

k D is modification factor for durability k T




η1 considers the type of reinforcement: is modification factor for temperature

[η1] = 1.0 for plain bars (certain surface resin just to lower the bond for ductilityreasons (rotation capacity) for the structure might give values < 1.0);

[η1]= (1.4 for indented steel bars and η1 = 2.25 for ribbed steel bars) The value forFRP bars must be determined for each bar type individually and depend on thesurface roughness, tendency to split the surrounding concrete and if the bondfailure is a concrete failure or is a shearing off the FRP bar ribs.

η2 considers the position of the bar during casting:

η2 = 1.0, when good bond conditions are obtained, as for:

all bars with an inclination of 45o-90o to the horizontal during concreting;

all bars with an inclination less than 45

o

to the horizontal, which are upto 250 mm from the bottom or at least 300 mm from the top of theconcrete layer during concreting. (Here the conditions for round barshave been taken into account. There might be FRP shapes differentfrom round, which must be paid attention to)

η2 = 0.7 for all other cases and for bars in structural parts built with slip forms.(A further reduction of coefficient might be necessary for unfavorable FRPshapes).

These requirements can be expected to be applied for FRP bars/rods also.

[η3] considers the bar diameter (for FRP with other shape than round specialreduction should be done)

For steel bars η3 = 1.0 for φ≤32 mm

For steel bars η3 = (132 - φ)/100 for φ>32 mm

with φ in mm.

2.1.2.2 Necessary tests for determination of ?-values

Pul l -out test for determination of maximum pull -out resistance

Appropriate tests are pull-out tests with central placement of bar and short bondlength, usually 3 bar diameters so the bond stress distribution along the bar becomesalmost even. However, depending on bar surface configuration longer bond lengthsshould be chosen to become representative. These tests should be used to determinethe maximum possible bond capacity of the FRP-bars/rods to obtain information ofmaximum possible bond resistance at estimation of η1-values. Bond stress-free barend slip relationships for different reinforcing bars and rods are shown in figure 1.Several of FRP bars/rods fail in surface layer of bar/rod and give f bd FRP. The curve for




steel reinforcement (in figure 1 Swedish Ks 600) may form a master curve to which theFRP curves can be related.

Figure 1. Pull-out test specimen with short bond length and centric placement of barand bond stress–free bar end slip relations for different reinforcement bars. Concrete

compressive strength 43-48 MPa. Master curve for steel deformed bar Swedish Ks600 envelops the slip curves.

FRP bars/rods have lower internal shear strength in resin between the fibers than thatof steel, which results in bond shear failure switch from concrete into the surface layerof FRP bar/rod when concrete becomes stronger than about 30 to 40 MPa, figure 2.For stronger concretes the bond shear resistance becomes independent of concretestrength. It is necessary to determine this highest bond shear strength f bdFRP for theFRP bar using pull-out test with short bond length and central bar placement.

0

5

10

15

0 20 40 60

fcu (Mpa)

÷ * ( M P a )

GFRP

CFRP

Figure 2. Influence of concrete strength on the bond strength of EUROCRETE FRPbars, Achillides, (1998).For stronger concretes than 35 MPa the bond strength

stabilizes at f bd FRP.

The bar diameter appears to play an important role in the bond behaviour of FRP barsto concrete. Larger diameter bars develop lower average bond strength than smallerdiameter bars. Bond development in FRP bars is affected by the diameter, due to theirlow shear stiffness in the axial direction. The value of the shear stiffness of FRP barsdepends mainly on the shear stiffness of the bar resin and the shear strength capacity

at the resin-fibre interface. When an FRP bar is pulled in tension through the surface,there can be a differential movement between the core and the surface fibres, which




results in a non-uniform distribution of normal stresses through the cross section ofthe bar. This is illustrated by the idealised stress distribution of figure 3.

σmax.

Axial load imposed by grip mechanism

FRP bar

Distribution of normal stresses

σmin

σav

Figure 3. Indicative distribution of normal stresses on a FRP bar cross-section subjected to axial load, Achillides (1998).

The actual developed bond strength between the bar and the concrete is directlyrelated to the value of the normal stress that occurs close to the surface of the bar(σmax). On the other hand, the calculated bond strength that the user anticipates as the“real” bond strength, is proportional to the value of the average normal stress over thewhole cross section of the bar (σav). As the diameter of the bar increases, thedifference between σmax and σav is also expected to increase, especially when the axialshear stiffness of the bar is relatively low as in the case of FRP bars, and the “real”

bond strength of the bar decreases. This effect, known in the literature as the “shearlag effect”, appears to be more significant for higher normal bar stresses. For steel

bars, this effect is less important, since the shear strength of steel is significantlyhigher. The shear lag effect and the bar diameter influence have to be included in the

?3-value for the FRP-bar/rod.

Tests for estimation of spli tting resistance of surrounding concrete

In most structures bars have limited confinement by concrete side covers of about 20 to50mm. For these bars the bond failure happens as a splitting of the concrete cover. Theless confined bars are surrounded by two concrete covers cx and cy in the corners ofstructure, figure 4. These corner bars should form the basis for estimation of the η1-values for splitting of concrete.

Figure 4. Concrete member section showing corner bars with covers c x and c y.

The concrete cover spli tting resistance along the bar can be studied in a pul l -out test

with eccentr ic placement of bar , figure 5, showing also bond stress-free bar end sliprelations for the specimens. Strain gauges should be used to monitor the appearance of




the cover crack. In very strong concretes the plasticization of stress in the cover is notso pronounced and the elastic stress configuration results in earlier cover cracking asmotivated by concrete strength increase. If the bar is very hard compared to concrete,stresses from concrete shrinkage may add to the bond ring stresses. Also differencesin thermal elongation between concrete and FRP bar in its radial direction may give

rise to extra stresses and may for some matrix resins be of interest to study. This testis not so well fitted to study the final splitting off of the surrounding concrete, becausethe support friction is disturbing. Benchmark results obtained from pull-out tests withordinary deformed steel reinforcing bar of corresponding diameter should be used forcomparison.

Figure 5. Bond stress-free bar end slip for pullout specimens with eccentric bar

placement. The bond stress levels for cracking of the concrete cover along the bar aremarked for the specimens in the diagram, Tepfers (1993).

In figure 6 schematic bond stress-slip relationships for pull-out tests are shown. Thediagram (a) represents a pull-out shear bond failure in concrete or FRP bar surface.The diagram (b) shows situation, when cover crack opens and in this very momentalso bond failure happens. This type of failure may happen for some certain membergeometries. Diagram (c) illustrates the relationship, when cover crack open and there

after the surrounding concrete resists higher splitting forces before it is pressed off




forming an ultimate splitting crack pattern. The tensile strength of concrete along thecrack pattern determines the fracture resistance. This type of splitting bond failure isthe usual one.

Free-end slip

A v e r a g e b o n d s t r e s s , τ

τu

Free-end slip

,

τsp

τu

Free-end slip

u

p

u

Figure 6. Bond stress-slip relationships for three types of failure. (a) shear bond failure; (b) cover crack induced failure; (c) failure by splitting off surroundingconcrete. Load could be raised after forming of cover crack. DeLorenzis & Tepfers

(2002).

The cover crack along the bar should not appear in serviceability stage because itgives rise to intrusion of corrosive agents to the bars/rods.

F igure 7. Tensile stress distributions in 1 elastic, 2 partly cracked elastic and 3 plastic stage and effect of concrete cover thickness c y upon the bond capacity of pullout specimens on occurrence of concrete cover crack along the bar. Open circles

Swedish Ks600 deformed steel reinforcing bars, closed circles GFRP bars (Hughes Brothers) and cross marks C-BARs

.




In figure 7 are shown the tensile stresses caused by bond stress in a concrete ring withthickness of the concrete cover cy surrounding the bar according to Tepfers (1973) and

bond stress f cbc when cover cracks related to cover thickness (bar diameter is ø andtensile strength of concrete is f ct):

In elastic stage: f cbc / f ct = ( 1 / tan a) [(cy + ø/2)

2

– (ø/2)

2

] / [(cy + ø/2)

2

+ (ø/2)

2

];In partly cracked elastic stage: f cbc / f ct= (cy + ø/2) / (1.664 ø tan a);

In plastic stage: f cbc / f ct = 2cy /ø tan a;

For the curves in figure 7 the angle a, is chosen to be 45o, which is the most likely.For the plastic stage a line is shown also for α = 30o. The results from anchoreddeformed bars, steel and FRP (C-BAR), are grouped between the plastic and elasticstages with the results being closer to plastic stage for thin concrete covers and to the

partly cracked elastic stage for thicker covers. This is natural, because, when

dimensions increase the plasticization becomes less pronounced. The C-BAR is softerthan steel bar and splits cover at higher bond stress. The hard steel bar gives stressconcentrations in concrete at lug tips, which favors start of cracking. The results fromGFRP bars (Hughes Brothers) with sanded surface come close to a = 30o line for

plastic stage, show smaller angle a and have less splitting effect on concrete, probablyforming a softening concrete zone close to bar, which can transfer some tension. Inthis way the results obtained from different FRP bars in eccentric pull-out tests can beevaluated for η1-value by comparison with benchmarks of steel reinforcing barsmodel curves for different stages.

The “Ring pull -out test ”, figure 8, is a more sophisticated test, which enables to

determine the splitting tendency of a bar in a direct way. With the "Ring pull-out test",Tepfers & Olsson (1992), the angle a of the bond forces in different stages of load can beestimated. The splitting tendency of the bar/rod increases, when the angle a increases.The ring pull-out test is a small cylindrical concrete body with axially placed bar/rod.The bond length is 3 bar diameters and the height of the concrete cylinder is equal withthe bond length. A thin steel cylindrical shell surrounds the concrete cylinder. At loadingthe radial and longitudinal bond force components are separated by a ring support withseveral teflon sheet layers, which prevents radial forces to be taken by support. Thecircumferential strain of the steel cylinder caused by radial bond force components ismeasured with strain gauges. The bond force component relation determines the angle a,which may change and increase when load increase. The measured free bar end slip and

ring strains are shown in figure 9.




Figure 8. Ring test for estimation of splitting tendency of reinforcing bars.

Figure 9. Bond stress – free bar end slip and steel ring strain – slip relations for Ring

test.

Overlap splice test

For basic information on η1 of the splitting resistance of the surrounding concrete,confining reinforcement and influence of uneven bond stress distribution along the

bars the investigation of the overlap strength of spliced reinforcement is an adequate

test procedure, figure 10. Several splice lengths, different concrete strengths, differentcove confinement and confinement by stirrups should be tested. It is discussed if the

bond resistance in an overlap splice is the same or lower that that for single anchored bar with the same concrete covers. However, studying the resistance in an overlapsplice the result will be on safer side for single anchored bar.

Figure 10. An example of beam lay out for testing the strength of tensilereinforcement overlap splices.




Figure 11. Distribution of bond stresses of the failure modes A, B and C. f bu = smallest ultimate failure pattern bond stress of appropriate type. f bc = bond stresswhich initiates the cover crack. Examples of ultimate failure patterns and the side

pressure options of overlapped bars. Tepfers (1973).

When the bond goes to failure for two side by side overlapped reinforcing bars the bond stress distribution along the bars can be as in the three modes according to figure11.

Mode A shows bond evenly distributed when the cover along the overlap length iscracked. The bond stress resistance f bu of the cover cracked surrounding concrete ishigher then the bond stress at cracking of concrete cover f bc.

Mode B shows uneven bond stress distribution with cover cracks at ends of overlapsplice. The bond stress resistance f bu of the cover cracked surrounding concrete is lowerthen the bond stress at cracking of concrete cover f bc. The cover cracked and not cover

parts of splice determine the maximum bond resistance. In cover cracked parts the abilityto slip of the bar increases and the bond stress becomes evenly distributed.

Mode C shows uneven bond stress distribution. When the bond stress f bu is reached,which cracks up the cover, it immediately results in pressing off the surrounding




concrete. The failure has zipper character and gives no warning of visible cracking before failure.

Evaluation of bond is to analyse possible resistance options by checking geometry,cracking sequences and stress distributions. There is also the question of single or

double pressure from side by side lapped bars, figure 11. It is not possible to pull out asingle bar and compare the result with that of overlap spliced bars with the samelength without analysing what happens. The cracking sequences and stressdistributions may be very different. For certain FRP-bars the bond force angle a may

be less than 45o, which means that these bars have less splitting tendency and givegood anchorage, when concrete cover determines the resistance. However these barsmay give less resistance, when confinement is excellent because of weak surfacelayer. FRP-bars with glossy surface and ribs give pronounced splitting forces andearly failure by pressing off concrete cover, but may give high pull-out resistancewhen confinement is good, because of strong FRP-bar surface layer.

Retrofi tting and strengthening of concrete structur es – rein forcing bars in groves

For retrofitting and/or strengthening of concrete members, near-sur face mounted

(NSM) FRP bars are put into grooves cut onto the concrete member and solidified toit with epoxy or mortar. The bond behavior of NSM FRP bars can be studied witheccentric pull-out tests, where the bar is fixed in groove accordingly, figure12.Different bond failure modes are encountered, depending on the combined effectof the test variables, De Lorenzis and Nanni (2002):

b p

B p = 300

dimensioni in mm F

F

F /4

2 3 0

z

d b F /4

70

3 5

u y

l a

l m

3

5

H p = 300h p

70

s p

z p = 160

y

Dimensions in mm

Figure 12. Pull-out specimen for determination of maximum bond strength for

reinforcement in grooves and bond stress–slip relations for CFRP rods, De Lorenziset al (2002).

- Pull-out, i.e., failure at the interface between rod and groove- filling material. Thisis the critical failure mode for sand covered rods, provided that the groove size issufficient to avoid splitting failure. It could be roughly estimated that a groove sizeequal to 1.5 times the actual rod diameter is enough to make pull-out thecontrolling failure mechanism

- Splitting of the groove-filling material, combined or not with cracking of thesurrounding concrete along inclined fracture planes. This is the critical brittlefailure mode for ribbed rods and for spirally wound rods, at least when the groove




surface is sufficiently rough to avoid failure at the interface between concrete andgroove-filling material. The ultimate load increases for increasing groove size.

- Failure at the interface between concrete and groove- filling material. Thismechanism is critical when the groove surface is smooth and has pseudo ductilecharacter due to residual friction. The average bond strength decreases as the

groove size increases, due to the non-uniform distribution of the bond stressesalong the perimeter of the groove.- Shearing of the ribs on the rod surface. The rod surface configuration is of crucial

importance. The superficial pattern of FRP rods to be used as NSM reinforcement,while being sufficiently rough to avoid failure by pull-out at the rod-epoxyinterface (apparently light sand covering is not enough for this purpose), shouldgenerate radial stresses as low as possible to delay splitting and cracking

phenomena and to allow a pseudo ductile bond–slip behavior.

I nf luence of temperature

The above tests have to be performed at normal temperature 20oC and also attemperature just above the glass transition temperature Tg of the FRP resin and athigher temperature for evaluation of influence of fire. If the matrix resin is susceptibleto temperature rise, a modification factor k T should be introduced describing theinfluence on bond strength change. It is necessary that the glass transition temperatureTg is declared for all FRP products to avoid unexpected degrading of strength. FRPsare sensitive for temperature rise as it becomes evident from table 1. Thick concretecovers are needed for fire protection and for time being these can be chosen as forsteel reinforcement until better knowledge is obtained.

Table 1.Bond between FRP bars/rods and concrete at elevated temperature. Thevalues in table give the highest and lowest value for 2 different FRP types.

The table is based on values given in Blontrock (1999).

Temperature (°C) Bond strength in % ofvalues at 20°C

100 35-80

150 20-40

220 10-20

Modif ication factor for durabil ity

Information from durability tests and life cycle analysis has to be established for thetested FRP reinforcement to determine the bond strength modification factor k D,which should be related environment and time, Dejke & Tepfers (2001, 2002), figure13. If the modification is in code in form of a reduction of tensile strength of FRP,

then a reduction of bond strength should not be necessary.




Desiredlife

Bond

strength

Dimensioningstrength

Change in bond strength

Time

Figure 13. Shematic diagram showing environmental influence on FRP bond strength, Dejke & Tepfers (2001, 2002).

Conclusions

The bond resistance level for steel reinforcing bars has been established from researchand practical application during very long time. The practice shows that the codesystem works. The bond performance of FRP-bars/rods should be compared with the

bond function and the established requirements for steel reinforcing bars. However,the some FRP bars/rods may give relatively low bond strength when confinement isgood, but have little splitting tendency and thereby develop good bond whenconfinement is only by concrete covers. Some other FRP show the opposite. Thesplitting of concrete cover in general comes at a higher load for FRPs then fordeformed steel bars of the same dimension and surrounding concrete covers. The

FRPs have softer surface and at bond force transfer create less stress concentrations inthe surrounding concrete. The bond splitting model curves, figure 7 enables tounderstand and to evaluate the bond splitting function of the bars. The η-values forFRP-bars/rods should be determined taking into consideration the presented testmethods and the relation to the η-values for steel reinforcement.in

2.2.3 Code clauses

2.2.3.1 Basic anchorage length

The basic length necessary for the transfer of the yield force of a bar or rod ofdiameter φ , or if square, maximum side dimension, is

l b = φ f yd/4 f bd; …………………………… (6.9-5)

f yd is the design yield stress (FRP reinforcement is fully elastic and has no yield stress.

An acceptable level of design stress for FRP must be given here).




2.2.3.2 Design anchorage length

The design anchorage length l b,net can be calculated from:

l b,net = 54321 ααααα l b As,cal/As,ef ≥ l b,min; .…………………(6.9-6)

where:

As,cal is the calculated area of reinforcement required by the design;

As,ef is the area of reinforcement provided.

α1, α2,α3,α4, and α5 are coefficients given below and defined as:

α1 coefficient taking into account the form of the bar (straight, bent, loop)

α2 coefficient taking into account the influence of one or more transverse bars (φt >0.6 ), figure 14 (6.9.6) along the design anchorage length l b,net

α3 coefficient taking into account the effect of confinement by the concrete cover

α4 coefficient taking into account the effect of confinement by transversereinforcement

α5 coefficient taking into consideration the effect of the pressure transverse to the plane of splitting along the design anchorage length.

l b is taken from equation (6.9-5).

Figure 14. Model Code 1990 figure 6.9.6.

l b,min denotes the minimum anchorage length:




* for bars in tension: l b,min > max {0.3l b; 10 φ , or if square maximum sidedimension; 100 mm}

* for bars in compression: l b,min > max {0.6l b; 10 φ , or if square maximumside dimension; 100 mm}.

The limitations of l b,min are give

• to ensure minimum active anchorage length• to take into account tolerances

The product (α3α4α5) is limited:

- for high bond bars: α3α4α5 > 0.7,

- for plain or indented bars (intermittent bond) or rods: α3α4α5 = 1.

Figure 15. Model Code 1990 figure 6.9.7.




Table2. Model Code 1990 table 6.9.1 giving the a-coefficients.

Minimum concrete cover for the model to be valid is one bar diameter.

The a-values should be for FRP reinf orcement accordingly: .

Form of bars:

Straight bars α1 = 1.0.

For FRP reinforcement bends do not contribute to anchorage because the bends are not sufficiently stiff. Thereforeα1 = 1.0.

Transverse bars with in termeshed fibers:

α2 = 0.7 if transverse bar is able to anchor at least 33% of tensile load.




Other types of bar crossings then intermeshed fibers have not enough strength.

Confinement by concrete:

α3 = 1 - 0.15(cd - 3φ)/φ ≥ 0.7 but ≤ 1.0, cd see figure 15.

Bends and loops do not contribute favorably because of limited bendingresistance of the FRP bars.

Confinement by not f iber in termeshed transverse reinforcement:

α4 = 1 - K λ ≥ 0.8 but ≤ 1.0

with λ = (∑ st A - ∑ min, st A )/ s A

∑ st A is the cross-sectional area of the transverse reinforcement along the

design anchorage length l b,net.

∑ min, st A is the cross-sectional area of the minimum transverse

reinforcement = 0.25 s A for beams and 0 for slabs

s A is the area of a single bar with maximum bar diameter.

K values are given in figure 16.

Figure 16. Coefficient K for not intermeshed transverse reinforcement confinement.

K = 0.10 for anchored bar As in an edge of surrounding reinforcement

K = 0.05 for transverse reinforcement Ast in the cover for anchored reinforcement.

K = 0 for transverse reinforcement Ast on the inner side of anchored reinforcement.

Confinement by transverse pressure:

α5 = - 0.04p ≥ 0.7 but ≤ 1.0




Comments concerning bond of FRP reinforcement on the Table 2

The contribution in bond strength of FRP bars/rods by confinement of transversereinforcement is probably less than that for steel reinforcement, because the modulusof FRP is less. Therefore the confinement effect is limited with α4 = 0.8.

For FRP-bars the transverse pressure will give less contribution to the anchorageeffect than for steel bars. The FRP-bars are softer than steel bars in transversedirection and will not attract as much pressure force as steel bars do. However, theharder steel bars introduce more splitting forces in concrete transverse to the directionof the pressure than the softer FRP bars do, which may cause earlier splitting fracture.These two effects are counteracting and therefore no change has been proposed for theα5-value.

For bars in compression α1, α2, α3, α4, and α5 are given the value 1.0. Thecompressive force, which has to be anchored, will be lower than the tensile one forFRP reinforcement, because the compressive strength of FRP-bars is considerablylower than the tensile one, which in its turn will require less bond strength.

2.2.3.3 Test method for determination of influence of transverse pressure

The effect of the transverse pressure can be studied with a hinged beam according tofigure 17. When the specimen is loaded the tensile force in reinforcement increasetogether with transverse pressure and simulate the situation at supports of memberswith very intense anchoring force transfer. The confining effect of stirrups can also bestudied adding these in the studied zone.

Figure 17. Hinged beam test for determination of influence of transverse pressure onanchorage, Magnusson (1997).

2.2.3.4 Design lap length of bars in tension

The design lap length is:




lo = 65431 ααααα l b As,cal/As,ef ≥ lo,min

where:

l b is the basic anchorage length

lo,min > max{0.3 α6 l b; 15φ; 200 mm}

α1, α3,α4 and α5 can be taken from 2.2.3.2 above.

However, for the calculation of α4, ∑ min, st A should be taken as 1.0 s A , with s A =

area of one spliced bar.

6α is a coefficient given in table 3 (MC90 Table 7.8.2) as a function of the percentage

of the reinforcement lapped within 1.3 lo from the center of the lap length consideredfigure 18.

Table 3. MC90 table 7.8.2

Percentage of lapped bars relative to the total cross-section of FRP bars/rods

≤20% 25% 33% 50% >50%

α6 1.2 1.4 1.6 1.8 2.0

For transverse distribution reinforcement α6 can be taken equal to 1.0.

Figure 18. Percentage of lapped bars in one section.MC90 figure 6.9.8.

2.2.3.5 Design lap length of bars permanently in compression

The design length of lap lo should comply with the condition lo > l b.




One bar of transverse reinforcement should be placed outside of each end of the laplength and within 4φ of the ends of the lap length.

2.2 Prestressed concrete

2.2.1 Anchorage of prestressing tendons

It is necessary to ensure that the anchorage device in the case of post-tensionedtendons or the anchorage length in the case of pretensioned tendons are able totransfer the design strength of the tendon to the concrete. If sleeves or mechanicalsplices (couplers) are used, these should be so located that the required strengths can

be obtained in all sections and the anchorages specified above can be attained.

Figure 19. Course of steel stresses along the anchorage zone of a pretensioned

member. MC90 figure 6.9.9.

2.2.2 Anchorage of pretensioned prestressing reinforcement

Two different bond situations should be considered due to the transverse deformationsof the tendon. A “push-in” along the transmission length, where the tendons becomethicker at release, and a “pull-out”, which refers to the anchorage length where theopposite occurs when the steel stress is increased due to loading.

The bond strength of pretensioned prestressing tendons depends on the loading case.The highest value applies to the transmission length - length to introduce the

prestressing force ("push- in" along transmission length, where the tendons become

thicker at release). Beyond that length lower bond strength has to be taken intoaccount ("pull-out", which refers to the anchorage length where the steel stress is




increased due to loading and tendons become thinner). This results in bilinear diagramfor the embedment length that is required to develop the design steel stress.

2.2.3 Design bond strengthThe design value of the bond strength for prestressing FRP tendons is:

f bpd = η p1 η p2 f ctd< f bd FRP;

where:

f ctd = f ctk (t)/1.50 is the lower design concrete strength; for the transmission length thestrength at the time of release, for the anchorage length the strength at 28 days;

f bd FRP is the design bond strength in the surface of the FRP-tendon (Determined in pull-out test with short bond length and central placement of bar/rod. Push-inor push-out situation in this case does not matter).

η p1 takes into account the type of prestressing tendon:

[η p1] = (1.4 for indented and crimped wires made of steel), (1.2 for 7-wire steelstrands.) The value for FRP tendon must be determined individually and

depend on the surface roughness, tendency to split the surrounding concreteand if the bond failure is a concrete failure or is a shearing off the FRP tendonsurface. FRP tendons usually have better bond properties than steel tendonsdue to rougher surface.

η p2 takes into account the position of the tendon:

η p2 = 1.0 for all tendons with an inclination of 45o-90o with respect to thehorizontal during concreting;

η p2 = 1.0 for all horizontal tendons which are up to 250 mm from the bottom

or at least 300 mm below the top of the concrete section during concreting;

η p2 = 0.7 for all other cases.

2.2.4 Basic anchorage length

The basic anchorage length defines the length that is required to develop the fullstrength in an untensioned tendon.

The basic anchorage length of an individual pretensioned tendon is:




l bp = Asp/(φπ )(f ptd/f bpd)

where:

f ptd = f ptk /[1.15]. The coefficient 1.15 have to be reconsidered for FRP tendons.

[f ptk ] is the characteristic stress level of tendon. For fully elastic FRP tendon this valueshould be introduced here. (For steel tendons f ptk is defined in MC 90 Clause 2.3.4.3).

The factor Asp/(φπ ) depends on the type of tendon. For circular cross-section Asp/(φπ )= ø/4 and for 7-wire strand (7/36)ø.

ø is the tendon diameter, or if square, maximum side dimension.

2.2.5 Transmission length

The transmission length of a pretensioned tendon is

l bpt =α8 α9 α10 l bp σ pi/ f pd

where:

α8 considers the way of release:

α8 = 1.0 for gradual release;

α8 = 1.25 for sudden release;

α9 considers the action effect to be verified:

α9 = 1.0 for calculation of anchorage length when moment and shear capacityis considered;

α9 = 0.5 for verification of transverse stresses in anchorage zone;

α10 considers the influence of bond situation:

[α10]= (0.5 for steel strands);

[α10]= (0.7 for indented or crimped steel wires);

σ pi is the FRP tendon stress just after release.




The use of narrow spaced stirrups or helices around the tendons and transverse prestressing may result in a shorter transmission length. This is not considered due tolack of experimental data.

The following has to be verified, because of lacking data for FRP tendons.

Tendon release that is obtained by sawing through the concrete and the FRP should beconsidered as gradual release. The text is valid for steel strands.

The transmission length can be estimated from draw in value (de) of the tendons at theend face of the concrete member. Assuming a linear FRP stress along the transmissionlength, this draw in shall be

de< 0.5 s pi/ t p l bpt with α9 = 1.0 in the expression for l bpt .

When the concrete member is sawn from a longer production unit, the draw-in cannot be estimated properly.

The basic anchorage length is related to “pull-out”. The transmission length isconnected to “push-in”. The ratio between the two is given by α10 .

2.2.7 Determination in test of the transmission length

Figure 20. Measured strain distribution in a square 70x70mm concrete rod 2000mmlong, prestressed with CFRP CFCC strand. A length of about 400mm is necessary fortransfer of the prestressing force. Tepfers et al (1992) tests vs. Cox et al (1999) model.

The test specimen can be a square concrete rod with centrally placed prestressed FRPstrand/rod. The ends of rod have spiral reinforcement to prevent splitting of rod end.The side of the section may be 5 to 7 times the rod diameter. The rod is prestressed,the concrete cast and when it has hardened the strain gauges are place on two oppositesides of the concrete rod with 50mm intervals so the expected transmission length is




covered with margin. Then the prestressed FRP rod is released and the compressivestrain on concrete surface is measured and compared with calculated strain in concretecaused by full transfer of prestressing force. The transmission length is the distancefrom end of rod to the point where the measured strain stabilize, see figure 20.

For FRP rods the transmission length is often shorter than for steel strands because ofrougher surface. This may result in too short transmission lengths causingconsiderable splitting forces on surrounding concrete.

2.2.7 Design anchorage length

The design anchorage length of a pretensioned prestressing tendon is:

l bpd = l bpt + l bp(σ pd - σ pcs)/f pd

where:

σ pd is the tendon stress under design load (σ pd ≤ f pd)

σ pcs is the tendon stress due to prestress including all losses.

If it is necessary, the required anchorage capacity may be obtained by additional endanchorages or non-prestressed reinforcement.

2.2.8 Development length

The development length is the distance from the end face to the concrete cross-section beyond which the distribution of the longitudinal stresses is considered linear.

For a rectangular cross-section and straight tendons situated near the bottom edge ofthe concrete section the development length is:

( ) bpt bpt p l l hl >+=22 6.0

where h is the total depth of the concrete section.






bond and tension stiffening effect is lost. Thereby the full length of the steel baryields, elongates and makes rotation of the member possible.

FRP bars/rods may have elastic fracture elongation of 3 to 4%. However, the diametercontraction is not of the same magnitude as for steel bars when these yield. For hybrid

fiber bars with successive failure of fibers and stress-strain relation like that of steel bar the “yielding” elongation do not result in diameter decrease of the samemagnitude as for steel bar. The fractured fibers are just broken and do not contributeto reduction in bar/rod diameter. Therefore the bond between the concrete cracks willnot be lost and the FRP bar main elongation will be only concentrated in vicinity ofthe cracks and this elongation will not be enough to cause necessary rotation formoment redistribution. A possible way of obtaining rotation capacity is to reduce the

bond capacity by making bars/rods with intermittent bond along the bar, Lees &Burgoyne (1997). However, the bar needs also to have good bond in the anchoragezones. This is not easy to produce bars with different bond capacity in different placesalong the bar length and to believe that the right part of the bar comes into the right

position in structure.

3 REFERENCES

1973 Tepfers, R. (1973). “A Theory of Bond Applied to Overlapped TensileReinforcement Splices for Deformed Bars”. Doctor thesis. Work No 723, Publ73:2. Division of Concrete Structures, Chalmers University of Technology,

Göteborg, May 1973, p. 328.

1992 Tepfers R., Olsson P-Å. (1992). “Ring Test for Evaluation of Bond Properties ofReinforcing Bars”. International Conference "Bond in Concrete - From

Research to Practice", Riga, Latvia, October 1992. pp. 1-89 - 1-99.

1992 Tepfers R., Molander I., Thalenius K., (1992), “Experience from testing ofconcrete reinforced with carbon fiber and aramid fiber strands”. XIV. NordicConcrete Congress & Nordic Concrete Industry Meeting, 6.-8. August 1992.Icelandic Concrete Association, Reykjavik. p.p. 337-347.

1993 CEB-FIP Model Code 1990. Design Code. Comité Euro-International duBéton. Thomas Telford Services Ltd, London 1993. p. 437.

1996 Machida A.: Designing Concrete structures with continuous fiber reinforcingmaterials. Dept. of Civil and environmental Engineering, Saitama University,Urawa, Saitama 338, Japan. First International Conference on Composites inInfrastructure, ICCI'96, 15-17 January 1996, Tucson Arizona, USA.Department of Civil Engineering and Engineering Mechanics, University ofArizona, Tucson, Arizona 85721, USA. Tucson 1996. p. 13.

1997 Japan Society of Civil Engineers, JSCE: Recommendations for design and

construction of concrete structures using continuous fiber reinforcingmaterials. Concrete Engineering Series 23. Tokyo 1997. p. 325.




1997 Lees J. M., Burgoyne Ch. J.: Rigid body analysis of concrete beams pre-tensioned with partly-bonded AFRP tendons. Non-Metallic (FRP)Reinforcement for Concrete Structures. Proceedings of the ThirdInternational Symposium. Vol. 2, Sapporo October 1997. pp. 759-766.

1997 Magnusson J. (1997). “Bond and Anchorage of Deformed Bars in High-Strength Concrete”. Chalmers University of Technology, Division of ConcreteStructures, Licentiate thesis, Work No. 1113, Publication 97:1, Göteborg,

November 1997. p. 234.

1997 Tepfers R., Karlsson M.: Pull-out and tensile reinforcement splice tests usingFRP C-BARs. Chalmers University of Technology, Division of BuildingTechnology. Work No: 13. Publication No: 97:2, Göteborg June 1997.Contribution to FRPRCS-3 Third International Symposium on Non-Metallic(FRP) Reinforcement for Concrete Structures in Sapporo 14-16 October 1997.

pp. 357-364.

1998 Achillides Z.: Bond behaviour of FRP bars in concrete. Ph D thesis, Centre ofCivil and Structural Engineering, Department of Civil and StructuralEngineering, The University of Sheffield, Sheffield July 1998. p.355.

1998 Tepfers R., Hedlund G., Rosinski B.: Pull-out and tensile reinforcement splicetests with GFRP bars. Proceedings of the ICCI´98, Second InternationalConference on Composites in Infrastructure, January 3-4, 1998, Tucson,Arizona, Vol. II pp. 37-51.

1999 Blontrock H., ”Properties of Fiber Reinforced Plastics at Elevated

Temperatures with Regard to Fire resistance of Reinforced ConcreteMembers”, Proceedings from Fourth international Symposium, FiberReinforced Polymer Reinforcement for Reinforced concrete Structures, ACIEd. Dolan C. W., Rizkalla S. H. and Nanni A., Baltimore 1999, pp. 43-54.

1999 Cox J. V., Guo J.: Modeling Stress State Dependency of Bond Behaviour ofFiber Reinforcedc Polymer Tendons. Fourth FRPRCS InternationalSyompsium, Baltimore. ACI International SP-188-68, Editors Dolan C. W.,Rizkalla S.H., Nanni A. 1999. pp. 791-805.

2000 fib, (CEB-FIP) Bulletin 10, "Bond of reinforcement in concrete", State of the art

report prepared by Task Group Bond Models, former CEB, Task Group 5.2.CH-1015 Lausanne , August 2000. p. 427.

2001 Dejke V., Tepfers R.: Durability and service life prediction of GFRP forconcrete reinforcement. FRPRCS-5. Edited by Ch. Burgoyne, University ofCambridge, Volume 1. Thomas Telford, London 2001. ISBN 0 7277 3029 0. pp505-514.

2001 Tamuzs V., Apinis R., Modniks J., Tepfers R.: The performance of bond of FRPreinforcement in concrete. SAMPE (Society for the Advancement of Materialand Process Engineering) Symposium, Long Beach, California, May 6th-10th,2001. pp. 1738-1748.



2002 De Lorenzis, L., and Nanni, A. (2002). "Bond between Near-Surface MountedFiber-Reinforced Polymer Rods and Concrete in Structural Strengthening",ACI Structural Journal, Vol.99 No.2, March-April 2002. pp. 123-132.

2002 De Lorenzis L., Tepfers R.: (2002). “Bond of FRP Reinforcement in Concrete -

a Challenge”. Contribution to Conference on Mechanics of CompositeMaterials, MCM-2002 in June 9-13, 2002, Riga, Latvia.

2002 Dejke V., Tepfers R.: Beständighet och livslängd hosfiberkompositarmerad betong. (Durability and service life of concrete reinforced with fibercomposites). Bygg & teknik, 94th edition, ISSN 0281-658X, Stockholm October#7-2002. pp. 16-22.

2002 Tepfers R.: (2002). ”Test system for evaluation of bond properties of FRPreinforcement in concrete”. Proceedings of the third International Symposiumon Bond in Concrete – from research to standards. Budapest November 20-22,

2002, pp. 657-666.

Documents

Bond Clause Proposal for FRP