FRCM composites for strengthening corrosion-damaged ...

FRCM Composites for Strengthening Corrosion-damaged

Structures: Experimental and Numerical Investigations

Thèse

Mohammed Elghazy

Doctorat en génie civil

Philosophiae doctor (Ph.D.)

Québec, Canada

© Mohammed Elghazy, 2018

Les matrices cimentaires renforcées de fibres (MCRF) pour

renforcer les structures en béton endommagées par la

corrosion: investigations expérimentales et numériques

Thèse

Mohammed Elghazy

Sous la direction de :

Ahmed El Refai, directeur de recherche

iii

Résumé

La corrosion des armatures en acier est l'un des mécanismes les plus destructifs pour les structures

en béton armé. La corrosion nuit non seulement à l'intégrité structurale et à l’aptitude au service

de la structure endommagée, mais peut aussi entraîner des défaillances inattendues ou des ruptures

fragiles. Malgré les dispositions rigoureuses de la plupart des codes de pratique pour éviter la

corrosion, des signes de dommages dus à la corrosion sont toujours signalés.

Récemment, des systèmes à matrice cimentaire renforcée de fibre (MCRF) ont été proposés

comme une technique innovante de renforcement/réparation pour les structures en béton afin de

surmonter les inconvénients associés à l'utilisation des systèmes de polymères renforcés de fibres

(PRF). Bien que l'utilisation de composites MCRF pour renforcer les éléments en béton non

endommagés ait prouvé son efficacité, très peu est connu sur la viabilité de leur utilisation pour

renforcer les éléments en béton endommagés à divers niveaux dus à la corrosion. De plus, les

comportements de post-réparation et la durabilité à long-terme des éléments corrodés et renforcés

par les systèmes MCRF et qui seront probablement exposés aux mêmes conditions

environnementales qui prévalaient avant leur réparation, n'ont pas retenu l'attention des chercheurs

dans la littérature. De plus, la plupart de nos infrastructures, telles que les ponts et garages de

stationnement, sont susceptibles d'être endommagées par la corrosion tout en étant soumises à des

charges oscillatoires qui provoquent de la fatigue. À ce jour, aucune information n'est disponible

sur l'effet de la combinaison de la charge de fatigue et de la corrosion dans les structures renforcées

par les systèmes MCRF.

Dans ce travail, les comportements monotones et de fatigue en flexion des poutres en béton

endommagées par la corrosion et renforcées par des systèmes MCRF ont été étudiés en plus de

leur performance à long-terme, c'est-à-dire après une exposition à un environnement corrosif après

leur renforcement. Le travail comprend des investigations expérimentales et numériques. Les

prédictions analytiques et les formulations théoriques actuellement disponibles dans les codes de

conception ont été aussi vérifiées par rapport aux résultats expérimentaux. Le programme

expérimental consistait à tester trente (30) poutres en béton à grande échelle de 150 × 250 × 2800

mm. Les poutres ont été construites et testées en configuration de charge à quatre points. Un

processus accéléré de corrosion a été utilisé pour corroder les armatures d'acier en traction dans le

tiers central des poutres. Les paramètres d'essai comprenaient le niveau de corrosion (représenté

iv

par 10, 20 et 30% de perte de masse dans l'acier de traction), le type de système de renforcement

utilisé (Polyparaphénylène benzobisoxazole (PBO-MCRF), MCRF de carbone et PRF), la quantité

de composites MCRF (1, 2, 3 et 4 couches), le schéma de renforcement MCRF (couches ancrées

aux extrémités par rapport aux couches continues sous forme U) et le régime de chargement

(monotone et fatigue).

Les résultats des tests ont montré que l'utilisation de composites MCRF améliorait

significativement le comportement en flexion des poutres corrodées. Les composites MCRF ont

contrôlé le mode de défaillance des poutres renforcées plutôt que le niveau de corrosion des barres

d'acier. Les poutres renforcées par la MCRF ont montré une augmentation de leurs résistances

ultimes variant entre 7 et 65% de celles des poutres vierges (poutres ni corrodées ni renforcées) en

fonction du type, de la quantité et du schéma de la MCRF utilisée. L'exposition des poutres

réparées par la MCRF à d’autres cycles de corrosion a entraîné une réduction de 23% de la perte

de masse de l'acier. Le schéma en U était plus efficace que le schéma d'ancrage aux extrémités à

retarder le délaminage des couches de MCRF dans les poutres renforcées et testées à court terme.

Il a également atténué l'effet des fissures de corrosion longitudinales et, par conséquent, a

augmenté l'efficacité du renforcement MCRF. Les essais de fatigue ont montré que la corrosion

des barres d'acier diminuait considérablement la résistance à la fatigue des poutres non renforcées.

Le renforcement avec des composites MCRF a augmenté la durée de vie en fatigue des poutres

endommagées par la corrosion de 38 à 377% de celle des poutres corrodées non-renforcées.

Cependant, le renforcement par MCRF n'a pas restauré la durée de vie en fatigue des poutres

vierges.

Dans l'étude numérique réalisée dans ce travail, des modèles d'éléments finis (ÉF)

tridimensionnels (3D) ont été développés pour simuler le comportement non linéaire des poutres

corrodées et renforcées par des composites MCRF et PRF à l'aide du progiciel ATENA-3D. Les

résultats de l'analyse numérique étaient en bon accord avec ceux obtenus expérimentalement en

termes de modes de défaillance, de déformations, de capacités de charge et de flèches. Les modèles

ÉF développés ont été capables de capturer le comportement non-linéaire des poutres testées avec

une bonne précision. Une étude paramétrique a ensuite été menée pour étudier l'effet de la

résistance à la compression du béton et de l'épaisseur de recouvrement des armatures sur l'efficacité

de renforcement des systèmes composites. Il a été observé que la rupture des poutres renforcées

v

par des FRCM était indépendante de la résistance à la compression du béton ou de l'épaisseur de

de recouvrement et était régie uniquement par le glissement du tissu dans la matrice.

Sur le plan analytique, les équations de conception de l’ACI-549.4R-13 (ACI 2013) ont été

évaluées à l'aide des données expérimentales obtenues à partir des tests. Il a été conclu que les

formulations théoriques de l’ACI-549.4R-13 peuvent raisonnablement prédire les résistances

ultimes des poutres renforcées ancrées à l'extrémité mais sous-estimer celles des poutres ancrées

en continu sous forme U. Un facteur de schéma de 1,1 a ensuite été proposé pour calculer la

résistance nominale des poutres renforcées par MCRF sous forme U.

Le résultat de ce travail a été publié (ou soumis pour publication) dans cinq articles de revues et

cinq conférences, comme détaillé tout au long de la thèse.

vi

Abstract

Corrosion of steel reinforcement is one of the most destructive mechanisms for reinforced

concrete (RC) structures. Corrosion not only impairs the structural integrity and the serviceability

of the damaged structure, but it may also lead to unexpected and brittle failures. Despite the

rigorous provisions of most codes of practice to avoid corrosion, evidences of corrosion damage

are still being reported.

Recently, fabric-reinforced cementitious matrix (FRCM) systems were proposed as an innovative

strengthening/repair technique for RC structures to overcome the drawbacks associated with the

use of the well-documented fiber-reinforced polymer (FRP) systems. While the use of FRCM

composites to strengthen un-damaged RC members has proven its efficiency, very little is known

about the viability of their use to retrofit RC members with various levels of corrosion damage. In

addition, the post-repair performance and the long-term durability of the FRCM-strengthened

corroded members, which most likely will be exposed to the same environmental conditions that

have prevailed prior their repair, have not received attention in the literature. Moreover, most of

our infrastructures such as bridges and parking garages are susceptible to corrosion damage while

continuously being subjected to oscillatory loads that cause fatigue. To date, no information is

available about the effect of combining fatigue loading with corrosion in FRCM-strengthened

structures.

In this work, the monotonic and fatigue flexural behaviors of corrosion-damaged RC beams

strengthened with FRCM systems were investigated in addition to their long-term performance,

i.e. after further exposure to corrosive environment following their strengthening. The work

includes experimental and numerical investigations. The analytical predictions and theoretical

formulations that are currently available in the design codes have been verified against the

experimental results. The experimental program consisted of testing thirty (30) large-scale RC

beams of 150×250×2800 mm. The beams were constructed and tested under four-point load

configuration. An accelerated corrosion process was utilized to corrode the bottom steel

reinforcement in the middle third of the test specimens. The test parameters included the level of

corrosion damage (represented by 10, 20, and 30% mass loss in the tensile steel), the type of the

strengthening system used (Polyparaphenylene benzobisoxazole (PBO-FRCM), C-FRCM, and

FRP), the amount of FRCM composites (1, 2, 3, and 4 layers), the FRCM strengthening Scheme

vii

(end-anchored versus continuously wrapped layers), and the loading regime (monotonic and

fatigue).

The test results showed that the use of FRCM composites significantly enhanced the flexural

behavior of the corroded beams. FRCM governed the failure mode of the strengthened beams

rather than the level of corrosion damage of the steel bars. FRCM-strengthened beams showed an

increase in their ultimate strengths that ranged between 7 and 65% of that of the virgin (neither

corroded nor strengthened) beam based on the type, amount, and Scheme of the FRCM used.

Exposing the repaired beams to post-repair corrosion resulted in 23% reduction in the steel mass

loss. The U-wrapped scheme was more efficient than the end-anchoring scheme in delaying the

delamination of the FRCM plies in the short-term repaired beams. It also mitigated the effect of

the longitudinal corrosion cracks and consequently increased the post-repair strengthening

effectiveness of FRCM systems. Fatigue tests showed that corrosion of steel bars dramatically

decreased the fatigue life of the unstrengthened-beams. Strengthening with FRCM composites

increased the fatigue life of the corrosion-damaged beams by 38 to 377% of that of the corroded-

unstrengthened beams. However, FRCM strengthening did not restore the fatigue life of the virgin

beams.

In the numerical study carried out in this work, three-dimensional finite element (FE) models

were developed to simulate the nonlinear behavior of the corroded beams strengthened with FRCM

and FRP composites using the software package ATENA-3D. The results of the numerical analysis

were in good agreement with those obtained experimentally in terms of failure modes, strains,

load-carrying capacities, and deflections. The developed FE models were able to capture the non-

linear behavior of the tested beams with good accuracy. A parametric study was then conducted

to investigate the effect of concrete compressive strength and thickness of concrete cover on the

strengthening effectiveness of the composite systems. It was observed that failure of FRCM-

strengthened beams was independent of the compressive strength of concrete or the thickness of

the concrete cover and was governed only by fabric slippage within the matrix.

Analytically, the design equations of ACI-549.4R-13 (ACI 2013) were assessed using the

experimental data obtained from the tests. It was concluded that the theoretical formulations of

ACI-549.4R-13 can reasonably predict the ultimate strengths of the end-anchored strengthened

beams but underestimated those of continuously-anchored beams. A scheme factor of 1.1 was then

viii

proposed to calculate the nominal strength of beams strengthened with continuously-wrapped

shape of FRCM.

The outcome of this work has been published (or submitted for publication) in five journal

articles and five conferences, as detailed throughout the thesis.

ix

Table of Contents

Résumé ……………………………………………………………......………………….…….. iii

Abstract ………………………………..………..……..……………………………..………… vi

Table of Contents …………………….………..……..…...……….………………...…..……… ix

List of Figures ……….…………….…..………..……..……………….………………....…… xiv

Acknowledgements……….………….…...……..……..…………………..……………....…… xx

Chapter 1 ....................................................................................................................................... 1

1.1 Background and Problem Definition................................................................................ 1

1.2 Thesis Structure ................................................................................................................ 3

Chapter 2 ....................................................................................................................................... 6

2.1 General ............................................................................................................................. 6

2.2 Corrosion of Steel Reinforcement .................................................................................... 6

2.2.1 Corrosion Mechanism of Steel Bars in Concrete ...................................................... 6

2.2.2 Accelerated Corrosion Process ................................................................................. 9

2.2.3 Effect of Steel Corrosion on Concrete Structures ................................................... 11

2.2.4 Behavior of Corroded-RC Beam under Monotonic Loads ..................................... 12

2.2.5 Behavior of Corroded-RC Beam under Fatigue Loads........................................... 15

2.3 FRCM Composites ......................................................................................................... 16

2.3.1 FRCM Acceptance Criteria and Design ................................................................. 18

2.3.2 FRCM Tensile Characterization ............................................................................. 18

2.3.3 Bond Behavior of FRCM Composites .................................................................... 20

2.4 FRCM-Strengthened Beams Under Monotonic Load .................................................... 21

2.5 Fatigue and Durability of FRCM-strengthened Beams.................................................. 25

2.6 Findings of Literature Review........................................................................................ 26

Chapter 3 ..................................................................................................................................... 27

3.1 Research Significance .................................................................................................... 27

3.2 Research Objectives ....................................................................................................... 27

3.3 Methodology .................................................................................................................. 28

3.3.1 Experimental Work Program .................................................................................. 28

3.3.2 Numerical Analysis ................................................................................................. 31

3.3.3 Analytical Investigation .......................................................................................... 32

Chapter 4 ..................................................................................................................................... 33

Résumé ...................................................................................................................................... 34

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4.1 Abstract .......................................................................................................................... 34

4.2 Introduction and Background ......................................................................................... 34

4.3 Experimental Program.................................................................................................... 36

4.3.1 Test Specimen ......................................................................................................... 37

4.3.2 Accelerated Corrosion Aging ................................................................................. 39

4.3.3 Materials ................................................................................................................. 40

4.3.4 FRCM Equivalent Axial Stiffness .......................................................................... 42

4.3.5 FRCM Repair Schemes........................................................................................... 43

4.3.6 Repair Technique .................................................................................................... 44

4.3.7 Test Setup and Instrumentation .............................................................................. 45

4.4 Test Observations ........................................................................................................... 46

4.4.1 Corrosion Cracks and Mass Loss ............................................................................ 46

4.4.2 Modes of Failure ..................................................................................................... 46

4.4.3 Load-deflection Response ....................................................................................... 48

4.4.4 Strength Analysis .................................................................................................... 49

4.4.5 Ductility Performance ............................................................................................. 53

4.4.6 Strain Response ....................................................................................................... 54

4.5 Theoretical Predictions ................................................................................................... 58

4.5.1 Design Provision ..................................................................................................... 61

4.6 Conclusions .................................................................................................................... 61

4.7 Notation .......................................................................................................................... 63

Chapter 5 ..................................................................................................................................... 65

Résumé ...................................................................................................................................... 66

5.1 Abstract .......................................................................................................................... 66


5.3 Experimental Program.................................................................................................... 68

5.3.1 Test Specimen and Materials .................................................................................. 69

5.3.2 Accelerated Corrosion Process ............................................................................... 71

5.3.3 FRCM Systems ....................................................................................................... 72

5.3.4 Strengthening Schemes ........................................................................................... 75

5.3.5 FRCM Installation .................................................................................................. 77

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5.3.6 Instrumentation and Test Setup .............................................................................. 77

5.4 Test Results .................................................................................................................... 78

5.4.1 Corrosion Observations .......................................................................................... 78

5.4.2 Failure Mechanisms ................................................................................................ 80

5.4.3 Strength Response ................................................................................................... 81

5.4.4 Fabric Strains .......................................................................................................... 86

5.4.5 Ductility and Energy Absorption ............................................................................ 87

5.5 Discussion ...................................................................................................................... 89

5.6 Predicted Strength Results ............................................................................................. 92

5.7 Conclusions .................................................................................................................... 95

Chapter 6 ..................................................................................................................................... 97

Résumé ...................................................................................................................................... 98

6.1 Abstract .......................................................................................................................... 98


6.3 Experimental Program.................................................................................................. 100

6.3.1 Test Specimen and Materials ................................................................................ 101

6.3.2 Accelerated Corrosion Process ............................................................................. 103

6.3.3 FRCM Composites................................................................................................ 104

6.3.4 FRCM Schemes .................................................................................................... 106

6.3.5 Repair Methodology ............................................................................................. 108

6.3.6 Test Setup and Instrumentation ............................................................................ 108

6.4 Test Results and Discussion ......................................................................................... 108

6.4.1 Corrosion Crack Pattern ........................................................................................ 108

6.4.2 Steel Mass Loss..................................................................................................... 109

6.4.3 Flexural Cracks Pattern and Failure modes .......................................................... 111

6.4.4 Flexural Response ................................................................................................. 114

6.4.5 Load-carrying Capacities ...................................................................................... 116

6.4.6 Ductility and Energy Absorption .......................................................................... 118

6.4.7 Strain Response ..................................................................................................... 120

6.5 Predicted Strengths ....................................................................................................... 123

6.6 Conclusions .................................................................................................................. 124

Chapter 7 ................................................................................................................................... 126

xii

Résumé .................................................................................................................................... 127

7.1 Abstract ........................................................................................................................ 127

7.2 Introduction and Background ....................................................................................... 127

7.3 Experimental Program.................................................................................................. 129

7.3.1 Test Specimen ....................................................................................................... 130

7.3.2 Accelerated Corrosion Technique......................................................................... 132

7.3.3 FRCM Composites................................................................................................ 133

7.3.4 FRCM Strengthening Configuration..................................................................... 135

7.3.5 Strengthening Procedure ....................................................................................... 137


7.4 Test Results and Discussion ......................................................................................... 139

7.4.1 Corrosion Crack Patterns and Actual Steel Mass Loss ......................................... 139

7.4.2 Monotonic Test Results ........................................................................................ 140

7.4.3 Fatigue Test Results .............................................................................................. 143

7.5 Conclusions .................................................................................................................. 151

Chapter 8 ................................................................................................................................... 154

Résumé .................................................................................................................................... 155

8.1 Abstract ........................................................................................................................ 155

8.2 Introduction and Background ....................................................................................... 155

8.3 Experimental Investigation .......................................................................................... 158

8.3.1 Test Matrix ............................................................................................................ 158

8.3.2 Test Specimen ....................................................................................................... 158

8.3.3 Externally-bonded Composite Systems ................................................................ 160

8.3.4 Strengthening Procedure and Configuration ......................................................... 161


8.4 Numerical Simulation .................................................................................................. 162

8.4.1 Constitutive Laws ................................................................................................. 163

8.5 Results and Discussion ................................................................................................. 170

8.5.1 Crack Pattern at Failure ........................................................................................ 170

8.5.2 Load-deflection Response ..................................................................................... 172

8.5.3 Load-carrying Capacity ........................................................................................ 177

8.5.4 Strain Response ..................................................................................................... 179

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8.5.5 Ductility ................................................................................................................ 182

8.6 Parametric Studies ........................................................................................................ 183

8.6.1 Effect of Concrete Compressive Strength (fc′)..................................................... 183

8.6.2 Effect of Concrete Cover ...................................................................................... 186

8.7 Conclusions .................................................................................................................. 188

Chapter 9 ................................................................................................................................... 191

8.8 Summary ...................................................................................................................... 191

8.9 Conclusions .................................................................................................................. 191

8.9.1 Effect of Corrosion on RC Beams ........................................................................ 191

8.9.2 Short-term Performance of Corroded Beams Strengthened with FRCM ............. 192

8.9.3 Long-term Performance of Corroded Beams Strengthened with FRCM ............. 194

8.9.4 Validation of ACI-549.4R-13 Design Equations .................................................. 195

8.9.5 Fatigue Performance of Corroded Beams Strengthened with FRCM .................. 196

8.9.6 Numerical simulation ............................................................................................ 197

8.10 Recommendation for Future Work .............................................................................. 198

8.11 Impact of Current Research.......................................................................................... 198

References .................................................................................................................................. 199

Bibliography .............................................................................................................................. 212

xiv

List of Figures

Figure 2.1: Corrosion process of steel bars inside concrete [5] ...................................................... 8

Figure 2.2: A schematic of microcell and macrocell types of corrosion [5] ................................... 8

Figure 2.3: FRCM composite system ........................................................................................... 17

Figure 2.4: Different fabric configurations [64] .......................................................................... 17

Figure 2.5: Actual and idealized stress-strain curve of FRCM coupon in tension [67]................ 19

Figure 2.6: Failure mechanism of a fiber bundle embedded in cementitious matrix [74] ............ 21

Figure 4.1: Typical dimensions and reinforcement details of the test specimen .......................... 38

Figure 4.2: Specimens connected in series inside the corrosion chamber .................................... 40

Figure 4.3: Strengthening materials: a) unbalanced PBO fabric, b) unidirectional carbon fabric,

and c) unidirectional carbon fabric ............................................................................................... 41

Figure 4.4: Idealized tensile stress-strain curve of FRCM coupon specimen [54] ....................... 43

Figure 4.5: Repair schemes: (a) Scheme I and (b) Scheme II ...................................................... 44

Figure 4.6: Repair procedure: a) removing the deteriorated concrete, b) patch repair, c)

roughening the concrete surface with sandblasting, and d) FRCM application ........................... 45

Figure 4.7: Positions of the electrical strain gauges along the outer fabric .................................. 45

Figure 4.8: Corrosion cracks pattern for specimen CU ................................................................ 46

Figure 4.9: Typical modes of failure: (a) SY-CC in beam CR-1P-I, (b) FD in beam CR-2P-I, (c)

FS in beam CR-2P-II, (d) MC-SFM in beam CR-3C-II, and (e) LR in beam CR-1FRP-I ........... 48

Figure 4.10: Effect of number of PBO-FRCM plies on the load-deflection curves ..................... 50

Figure 4.11: Effect of the repair scheme on the load-deflection curves ....................................... 50

Figure 4.12: Effect of FRCM materials on the load-deflection curves......................................... 51

Figure 4.13: Normalized ultimate load versus the equivalent stiffness ........................................ 53

Figure 4.14: Load-strain curves for specimens with repair Scheme I .......................................... 55

Figure 4.15: Load-strain curves for specimens with repair Scheme II ......................................... 56

Figure 4.16: Strain profile in the PBO fabric for specimen CR-4P-I ........................................... 57

Figure 4.17: Strain profile in the PBO fabric for specimen CR-4P-II .......................................... 58

Figure 4.18: Strain profile in the carbon fabric for specimen CR-3C-II ...................................... 58

Figure 4.19: Stress and strain distribution at ultimate stage ......................................................... 59

Figure 5.1: Test specimen geometry and reinforcement details. (All dimensions in mm) ........... 71

Figure 5.2: Specimens inside the environmental chamber during a dry cycle ............................. 72

xv

Figure 5.3: FRCM systems: a) PBO-FRCM (Unbalanced PBO fabric) and b) C-FRCM

(Unidirectional carbon fabric)....................................................................................................... 73

Figure 5.4: Stress-strain relationships for FRCM-tensile coupons [97] ....................................... 74

Figure 5.5: Strengthening schemes: a) Scheme I and b) Scheme II ............................................. 76

Figure 5.6: FRCM installation procedure: a) removing the deteriorated concrete, b) patch

repairing and sandblasting, and c) installation of PBO-FRCM composite ................................... 77

Figure 5.7: Profile of steel bars: a) uncorroded bar, b) corroded bar extracted from CSA-4P-I

(12.6% mass loss), c) corroded bar extracted from CSB-3C-II (18.6% mass loss), and d)

corroded bar extracted from CUC (22.5% mass loss) .................................................................. 79

Figure 5.8: Actual and theoretical mass loss versus the duration of corrosion process ................ 79

Figure 5.9: Modes of failure: a) FRCM delamination, b) fabric slippage with partial fabric

debonding within the matrix, and c) matrix cracking with extensive fabric slippage .................. 81

Figure 5.10: Load-deflection relationships for corroded-unstrengthened beams ......................... 82

Figure 5.11: Load-deflection relationships for corrosion-damaged FRCM-strengthened beams: a)

beams of Group A, b) beams of Group B, and c) beams of Group C ........................................... 83

Figure 5.12: Normalized strength versus the FRCM equivalent axial stiffness, Kf ...................... 85

Figure 5.13: Load versus outer fabric strain for beams strengthened in a) Scheme I and b)

Scheme II ...................................................................................................................................... 86

Figure 5.14: Normalized ductility index versus stiffness factor 𝛽𝑓 % ......................................... 89

Figure 5.15: Normalized energy absorption index versus stiffness factor 𝛽𝑓 % .......................... 89

Figure 5.16: Effect of corrosion damage on the ultimate strength of strengthened beams .......... 90

Figure 5.17: Normalized strength versus stiffness factor 𝛽𝑓 % .................................................... 91

Figure 5.18: Predicted versus experimental flexural response for beams strengthened with a) two

layers of PBO-FRCM in Scheme I, b) four layers of PBO-FRCM in Scheme II, and c) three

layers of C-FRCM in Scheme II ................................................................................................... 94

Figure 6.1: Schematic of the testing procedure of the short- and long-term beams ................... 101

Figure 6.2: Typical dimensions and reinforcement details of the test beam (all dimensions in

mm) ............................................................................................................................................. 102

Figure 6.3: A schematic of the electrical connection .................................................................. 104

Figure 6.4:FRCM systems; a) PBO-FRCM (unbalanced PBO fabric) and (b) C-FRCM

(unidirectional carbon fabric) - all dimensions in mm ................................................................ 105

xvi

Figure 6.5: Idealized tensile stress-strain curve of FRCM coupon specimen ACI 549.4R-13 [54]

..................................................................................................................................................... 106

Figure 6.6: FRCM repair schemes; a) Scheme I and b) Scheme II ............................................ 107

Figure 6.7: Corrosion cracks patterns; a) typical corrosion cracks pattern for short-term

specimens (beam CU); b) beam CRL-4P-I; c) beam CRL-4P-II; and d) beam CRL-3C-II ....... 110

Figure 6.8: Failure mode of a) beam CU due to steel yielding and concrete crushing; b) beam

CRS-4P-I due to FRCM delamination; c) beam CRL-4P-I due to premature FRCM delamination;

d) beam CRL-4P-II due to PBO-fabric debonding from matrix; and e) beam CRL-3C-II due to

fabric slippage ............................................................................................................................. 112

Figure 6.9: Load-deflection relationships of a) short-term beams; b) long-term beams; c) beams

repaired with PBO-FRCM (short-term and long-term); and d) beams repaired with C-FRCM

(short-term and long-term) .......................................................................................................... 116

Figure 6.10: Load-strains relationships for a) beams repaired with PBO-FRCM and b) beams

repaired with C-FRCM ............................................................................................................... 122

Figure 7.1: Geometry and reinforcement details of the test specimen (all dimensions in mm) . 132

Figure 7.2: Specimens connected in series inside the corrosion chamber .................................. 133

Figure 7.3: a) Unbalanced PBO fabric and b) Unidirectional carbon fabric. ............................. 134

Figure 7.4: FRCM strengthening schemes: a) Scheme I and b) Scheme II ................................ 136

Figure 7.5: FRCM strengthening procedure: a) removing the deteriorated concrete after

corrosion, b) patch repair, c) PBO-FRCM application, and d) C-FRCM application ................ 137

Figure 7.6: Test setup .................................................................................................................. 138

Figure 7.7: Corrosion crack pattern for specimen FCU .............................................................. 139

Figure 7.8: Profile of the steel bars: a) Uncorroded bar, b) corroded steel bar extracted from

beam MCS-4P-II, and c) fatigue rupture of a corroded steel bar extracted from FCS-3C-II ..... 140

Figure 7.9: Failure modes of the monotonically tested beams: a) Steel yielding followed by

concrete crushing; b) FRCM delamination; c) Fabric slippage and partial debonding; and d)

Matrix cracking followed by fabric slippage .............................................................................. 141

Figure 7.10: Load-deflection relationships of the monotonically tested beams ......................... 142

Figure 7.11: Variation of fatigue life of the strengthened beams with the stiffness factor 𝛽𝑓 .... 144

Figure 7.12: Load-deflection Hysteresis for a) beam FCU; b) beam FCS-4P-I; c) beam FCS-4P-

II; and d) beam FCS-3C-II .......................................................................................................... 147

xvii

Figure 7.13: Fatigue cracks at midspan of the strengthened beams (Side view) ........................ 148

Figure 7.14: Fatigue rupture of steel bars in a) beam FCU and b) beam FCS-4P-I ................... 148

Figure 7.15: Effect of fatigue cycles on the stiffness of the tested beams .................................. 149

Figure 7.16: Effect of fatigue cycles on the concrete and fabric strains ..................................... 150

Figure 8.1: Geometry and details of steel, P-FRCM, and C-FRP reinforcement of the tested

beams (all dimensions in mm)..................................................................................................... 159

Figure 8.2: a) PBO fabric used in the FRCM composite system and b) Carbon sheets used in the

FRP composite system ................................................................................................................ 160

Figure 8.3: Strengthening procedure of corroded beams: a) removing the deteriorated concrete,

b) patch repairing and sandblasting, c) installation of P-FRCM composite, and d) installation of

C-FRP sheets ............................................................................................................................... 162

Figure 8.4: Constitutive laws of concrete and cementitious matrix: a) compressive hardening law,

b) compressive softening law, and c) tensile softening law ........................................................ 164

Figure 8.5: C-FRP/concrete interfacial bond stress-slip model according to Lu et al. [117]

adopted for various concrete mixes ............................................................................................ 167

Figure 8.6: PBO-fabric /matrix interfacial bond stress-slip model according to D’Ambrisi et al.

[120] ............................................................................................................................................ 168

Figure 8.7: a) Meshing of P-FRCM strengthened beams, b) reinforcement layout for beams

strengthened with 4 layers of P-FRCM, and c) reinforcement layout for beams strengthened with

C-FRP sheet ................................................................................................................................ 169

Figure 8.8: Numerical and experimental crack patterns at failure for a) beam UUa, b) beam CS-

A-1C, and c) beam CS-A-4P ...................................................................................................... 172

Figure 8.9: Numerical and experimental load-deflection responses for beams of group A ....... 173

Figure 8.10: Numerical and experimental load-deflection responses for beams of group B ..... 176

Figure 8.11: Gain and decline in % in the ultimate loads, Pu, with respect to that of the control

beam (UU) .................................................................................................................................. 177

Figure 8.12: Fiber and concrete strain response ......................................................................... 180

Figure 8.13: Tensile steel strain response for beams CS-A-1C and CS-A-4P............................ 181

Figure 8.14: Strain profile of the internal and external reinforcement at ultimate ..................... 181

Figure 8.15: Effect of the concrete compressive strength on the load-deflection response for

beams a) CS-A-1C, b) CS-A-2P, and c) CS-A-4P...................................................................... 185

xviii

Figure 8.16: Effect of the concrete cover on the load-deflection response for beams a) CS-A-1C,

b) CS-A-2P, and c) CS-A-4P ...................................................................................................... 188

xix

To the researchers and engineers who appreciate the value of science and

knowledge

To the ones who were

and will always be

in my mind….

in my heart….

.…and in my Soul

Mohammed Elghazy

xx

Acknowledgements

Every journey has to have an ambitious and a defined end. My journey as a PhD student started

in December 2014. My main objective was to acquire my PhD at Laval. During the last three years,

I have developed another goal for myself: to explore my own feasibility as a researcher and how

to strengthen myself as an individual. The outcomes were greatly dependent on people, financial

status, and the surrounding environment.

Therefore, I would like to express my sincere appreciation and gratitude to my thesis supervisor

Prof. A. El Refai for his guidance and continuous encouragement during the course of this work.

Thanks for being a friend and a firm supervisor. I would like also to express my sincere gratitude

to our collaborators Prof. A. Nanni from University of Miami and Prof. U. Ebead from Qatar

University, for their encouragement and support during my research.

I would also like to acknowledge and thank all the technicians of the structural laboratory at

Laval University, especially, Mr. R. Malo for his precious assistance through the experimental

tests. The help provided by Mr. K. Attia, Mr. A. Abbadi, and Mr. N. Allam during the tests is

greatly appreciated.

I would like to acknowledge Laval University and Qatar Foundation for the financial support.

The donation of composite materials provided by Ruredil, Italy and Simpson Strong-Tie, USA

represented by Mr. Brad Erickson is greatly appreciated.

I would like to express my sincere thanks and appreciation to my mother. Without her

unconditional love, support, and encouragement, the achievement of this work would not be

possible. I cannot end my acknowledgement without expressing my deep appreciation to my sister

from the bottom of my heart for her encouragement and endless love

https://courriel.ulaval.ca/owa/?ae=Item&t=IPM.Note&id=RgAAAACSza%2biCuLGT4f618YW8XCPBwAclwyo5YCoQ5TOXPQv7ABGApejm2e7AACia1O9LRU4RYdcG8XuV3bpAAHAd2YLAAAJ

1

1. Chapter 1

Introduction

1.1 Background and Problem Definition

Repair/strengthening of reinforced concrete (RC) structures is motivated by several factors

including aging, change in use, increased loads, code compliance, and environmental damages

such as corrosion. Corrosion of steel reinforcement is one of the major durability concerns for

concrete structures especially in coastal areas and in cold regions where de-icing salts are heavily

used. Pitting corrosion reduces the cross-sectional area of the steel bars and may lead to significant

loss of their ductility [1,2]. The expansion of corrosion products causes concrete cracking and

impair the composite action between steel and concrete. As a result, the load-carrying capacity and

the service life of the corroded member are considerably jeopardized [3–5].

Corrosion of steel reinforcement combined with fatigue stresses significantly reduces the fatigue

life of RC structures and may lead to unexpected failures especially in bridges [6,7]. According to

the U.S. Federal Highway Administration report published in 2002 [8], approximately 15% of the

highway bridges in USA are considered structurally deficient. Their maintenance and repair costs

exceed 8.3$ billion dollars annually. In Canada, the total direct costs of corrosion were 23.6$

billion dollars in 2003 [9]. The statistics for Europe, Asian Pacific countries, and Australia are not

by any means better than those in North America. Government agencies and industrial firms are

looking for more durable and less costly materials and techniques to maintain and repair our

infrastructures. Hence, structural engineers are in continuous search for new construction materials

and innovative rehabilitation techniques for such deficient structures.

In the last decades, externally-bonded strengthening technologies based on organic matrices,

referred to as fiber-reinforced polymers (FRP), have proven success in restoring the serviceability

and strength of RC structures [10–12]. However, several problems associated with the use of FRP

have been documented. FRP materials are flammable and prone to deterioration and the loss of

their mechanical and bond properties at high temperatures [13]. Their epoxy-based agents lose

their stiffness and strength and change from a stiff material to a viscous material with poor

properties when exposed to elevated temperatures, which affect their bond strength [14,15].

2

Structures located in hot climates or those at risk of fire can easily be vulnerable when exposed to

elevated temperatures. The toxicity nature of epoxy and its poor thermal compatibility to the

concrete substrate add another dimension to the drawbacks of FRP systems [16].

In order to overcome the drawbacks associated with the use of FRP composites, the need to

replace the organic binder (epoxy) by an inorganic binder has been raised. The use of fabric-

reinforced cementitious matrix (FRCM) systems has been introduced as an alternative promising

strengthening technique to FRPs. The FRCM system is a composite material consisting of fabric

meshes made of long dry-woven intermittent in two orthogonal directions and embedded in

cement-based matrix that serves as a binder. The embedded grid is shielded between the mortar

layers thus minimizing its fire vulnerability. In addition, the compatibility between the mortar used

and the concrete substrate is inherited since both materials have cement as a common “base”. The

cement-based mortars used in FRCM also act as barriers against chloride ions penetration thus

protecting the main reinforcing bars from corrosion attack. Its lightweight, high tensile strength,

and ease of application makes the system very appealing to engineers. With their innovative

features, FRCM systems can ensure the endurance of the rehabilitation process and consequently

the sustainability of the strengthened structure.

Recently, significant efforts have been made to introduce FRCM composites in the construction

industry and the use of FRCM composites to strengthen RC structures was initiated in several field

applications. A significant amount of research has been devoted to study the flexural and shear

behaviors of undamaged RC members strengthened with FRCM composites. Nonetheless, the

feasibility of using FRCM composites to strengthen corrosion-damaged RC structures has received

little attention. The challenge in using FRCM to repair/strengthen corrosion-damaged concrete

members rises from the fact that the technique uses cementitious mortars rather than epoxies,

which necessitates proper surface preparation to ensure adequate mechanical and chemical

adhesion to the concrete substrate. Factors such as the absorption properties of the substrate, its

degree of carbonation, its moisture condition, and its cleanness, in addition to the existence of

micro-cracks and/or contaminants on the repaired surface may affect the desired performance.

Unfortunately, surfaces of corroded concrete structures are characterized by their random texture,

unpredictable crack distribution, and concrete fragmentation. Volume expansion resulting from

corrosion of steel bars in addition to the existence of corrosion products cause the weakness of the

3

surrounding concrete and the loss of its integrity. In fact, these conditions might consist a serious

obstacle to the use of FRCM in strengthening corroded structures. To the authors’ knowledge, very

few of the previous studies in which FRCM technique was adopted has addressed this problem,

not to mention the long-term performance of the strengthened members being exposed to the same

conditions that have caused their deterioration. Therefore, the work presented herein aimed at

tackling these problems and filling the gap in our knowledge on the use of FRCM composites in

strengthening corrosion-damaged RC elements.

1.2 Thesis Structure

The purpose of this work is to investigate the short- and long-term structural performance of

corrosion-damaged RC beams strengthened with FRCM composites. This thesis is organized in

nine chapters as follows:

• Chapter (1) provides background on the subject as well as the definition of the research

problem.

• Chapter (2) provides a comprehensive literature review on topics that are related to the work

of this study. This includes a review on the corrosion mechanism in RC structures, a review on

the flexural behavior of corrosion-damaged RC members under both monotonic and fatigue

loads, a full historical review on previous research that was carried out to determine the

mechanical properties of FRCM composites, and an inclusive review on the flexural behavior

of RC members strengthened with FRCM composites.

• Chapter (3) highlights the research significance along with the research objectives of this work.

A detailed description of the utilized methodology to achieve these objectives is also included.

• Chapter (4) presents the first journal paper submitted to the American Society of Civil

Engineers (ASCE) Journal of Composites for Construction. The paper is titled “Corrosion-

Damaged Reinforced Concrete Beams Repaired with Fabric-Reinforced Cementitious Matrix

(FRCM)”. The paper reports on the feasibility of using FRCM to strengthen/repair RC beams

with moderate level of corrosion damage. The test results of eleven large-scale RC beams were

presented including those of nine beams that were subjected to accelerated corrosion process

for 70 days to obtain an average mass loss of 12.6% in the tensile steel reinforcing bars and

4

those of two beams that were tested as controls. The effect of various levels of corrosion damage

on the flexural response of the FRCM-strengthened beams was investigated throughout Chapter

5. In this paper, the test parameters included the number of fabric plies (1, 2, 3, and 4), the

FRCM repair scheme (end-anchored and continuous U-wrapped strips), and FRCM materials

(carbon and PBO). The analysis and discussion of the test results were presented in terms of

mode of failure, load-deflection response, load carrying capacity, ductility performance, and

strain responses. In addition, the experimental results were compared to the current design

guidelines of ACI-549.4R-13.

• Chapter (5) presents the second paper that was published in the Journal of Composite

Structures. The paper is titled “Effect of Corrosion Damage on the Flexural Performance of RC

Beams Strengthened with FRCM Composites.” The paper reports on the effect of the level of

corrosion damage on the flexural response of RC beams strengthened with FRCM composites.

Three theoretical tensile steel mass losses were considered in this study, namely 10, 20, and 30

%, which represented moderate, severe, and very severe degrees of corrosion damage. The test

parameters also included the fabric type (PBO and carbon), the number of FRCM layers (two,

three, and four), and the strengthening scheme (end-anchored and continuously wrapped). The

test results were analyzed and discussed to highlight the effect of various test variables on the

failure mechanism, the flexural performance, the fiber, steel, and concrete strain responses, the

ductility, and the flexural strengths of the tested beams. In addition, the experimental results

were compared to the current design guidelines of ACI-549.4R-13.

• Chapter (6) presents the third journal paper that was submitted to the Journal of Construction

and Building Materials and was titled “Post-repair Flexural Performance of Corroded Beams

Rehabilitated with Fabric-Reinforced Cementitious Matrix (FRCM) under Corrosive

Environment”. The test results of nine RC beams were reported in this paper including those of

one beam that were neither corroded nor repaired, another beam that was corroded and not

repaired, and six beams that were corroded and then repaired in two phases. Beams of Phase I

(short-term) were subjected to an accelerated corrosion process for 210 days before being

strengthened with FRCM whereas beams of Phase II (long-term) were initially subjected to

accelerated corrosion for 70 days, then repaired with FRCM and exposed to further corrosion

5

for 140 days. The test parameters also included the type of FRCM materials (PBO and Carbon)

and the FRCM repair scheme (end-anchored and continuous wrapping).

• Chapter (7) presents the fourth journal paper that was submitted to the ASCE’s Journal of

Composites for Construction. The paper is titled “Fatigue and Monotonic Behavior of

Corrosion-damaged Reinforced Concrete Beams Strengthened with FRCM Composites.” This

paper aimed at investigating the potential of using FRCM composites to restore the fatigue life

of RC beams that were severely damaged due to corrosion (20% tensile steel mass loss). The

paper provides better understanding on the flexural behavior of FRCM-strengthened beams

damaged due to corrosion. The test results of twelve beams were reported. The test parameters

included the fabric material (PBO and Carbon), the number of FRCM plies, the strengthening

configuration, and the type of loading (monotonic and fatigue). The results were discussed and

presented in terms of fatigue life, fatigue behavior, modes of failure, strain response, and

stiffness degradation.

• Chapter (8) presents the fifth paper that was submitted to the Journal of Engineering Structures

and titled “Finite Element Modeling and Experimental Results of Corroded Concrete Beams

Strengthened with Externally-bonded Composites”. This paper presents the numerical

simulation of corroded RC beams strengthened with carbon-FRP (CFRP) and PBO-FRCM

composites under flexural loading using ATENA software package. The CFRP composite was

modeled as discrete reinforcement bonded directly to the concrete substrate without binder

while the PBO-FRCM was modeled using a more detailed approach that involved modeling the

fabric and the matrix layers. Interfacial bond stress-slip models were adopted at the

CFRP/concrete and the PBO-fabric/matrix interfaces to simulate the failure mechanism

observed during the test. The load-carrying capacities, load-deflection responses, and load-

strains responses were evaluated and compared with the experimental results to validate the

accuracy of the model. The validated models were used in a parametric study to investigate the

effect of varying the concrete compressive strengths on the flexural behavior of the FRCM-

strengthened beams after corrosion.

• Chapter (9) includes a summary of this study and the overall conclusions based on the

experimental, analytical, and numerical results obtained. Recommendations for future work are

also suggested in this chapter.

6

2. Chapter 2

Literature Review

2.1 General

The corrosion of the steel reinforcement in concrete structures is a major challenge facing

structural engineers while maintaining ageing infrastructures. Corrosion usually causes

engineering and economic problems, which has led to a considerable amount of research work

devoted to study the performance of corroded-RC flexural elements under both monotonic and

fatigue loads. In recent years, advanced composite materials in the form of externally-bonded

fabric-reinforced cementitious matrix (FRCM) have been introduced as innovative strengthening

techniques for RC structures. Several publications have been published in recent years to document

the effectiveness of FRCM composites to strengthen undamaged concrete structures. Nevertheless,

their application to strengthen/repair corrosion-damaged concrete elements is still lacking.

This chapter provides a comprehensive review on the corrosion of steel reinforcement in concrete

structures as well as the use of externally-bonded FRCM composites to strengthen RC flexural

elements. A review of the available literature concerning the research work on corroded FRCM-

strengthened flexural elements is presented along with the factors that affect their structural

behavior.

2.2 Corrosion of Steel Reinforcement

2.2.1 Corrosion Mechanism of Steel Bars in Concrete

The high alkaline environment of concrete normally protects the steel reinforcement against

corrosion by creating a passive film of iron oxides at the steel/concrete interface [5]. In addition,

the concrete cover works as a barrier to protect the steel bars against harsh environmental

conditions [5]. Exposure to sea salt spray in marine environment or de-icing salts in cold regions

are the most common causes of corrosion in concrete worldwide. Steel corrosion in concrete can

be triggered by chloride attacks and concrete carbonation. Chloride ions diffuse through the

concrete pores and cracks until they reach the surface of the steel bars. Chlorides work as catalysts

to initiate the corrosion reactions in the weak spots on the steel surface. In case of concrete

7

carbonation, the carbon dioxide gas (CO2) in the atmosphere dissolves in water (H2O) to form the

carbonic acid (H2CO3). The acid then penetrates into the concrete and drops its pH level to less

than 8.5. Consequently, the passive layer previously formed on the bar surface decades and

corrosion reactions are then initiated [5,17].

Corrosion of steel in concrete is an electrochemical process. It involves the transfer of charge

(electrons) from one location to another [17]. The corrosion reaction is similar to a galvanic cell

that consists of an anode and a cathode as shown in Figure 2.1. At the anode, the iron (steel bar)

dissolves or oxides and the ferrous electrons are released as illustrated in Equation (2.1). The

released electrons react with water and oxygen at any other location of the steel bar that acts as a

cathode to form dissolved hydroxide ions (Equation (2.2)).

Fe Fe+2 + 2e− (Anode reaction) Eq. (2.1)

2e−+H2O +1

2O2 2OH− (Cathode reaction) Eq. (2.2)

The ferrous ions (Fe+2) at the anode react with the hydroxide ions (OH−) that are released from

the cathode to form ferrous hydroxide [Fe (OH)2] as per Equation (2.3). In the presence of

sufficient amount of water and oxygen, the ferrous hydroxide reacts to produce ferric hydroxide

[Fe (OH)3] as shown in Equation (2.4). Finally, the ferric hydroxide [Fe(OH)3] decays to hydrated

ferric oxide [Fe2O3. H2O] or rust [5]. This reaction is shown in Equation (2.5). The volume of the

corrosion products is much greater than the volume of steel (volume of rust is about 6 times of that

of iron). Therefore, longitudinal cracks parallel to the reinforcing bars generally appear in the

concrete cover as indication of corrosion. The cracks allow to more quantities of moisture and

oxygen to reach the steel bars, which boosts the rate of corrosion reactions.

Fe+2+2OH− Fe (OH)2 Eq. (2.3)

4Fe (OH)2+ O2 + 2H2O 4Fe (OH)3 Eq. (2.4)

2Fe (OH)3 Fe2O3. H2O + 2H2O Eq. (2.5)

8

Figure 2.1: Corrosion process of steel bars inside concrete [5]

Corrosion of steel in concrete can occur as micro-cells (uniform iron removal) or as macro-cells

(local iron removal). Figures 2.2a and 2.2b show schematics of micro-cell and macro-cell types of

corrosion, respectively. At the beginning, corrosion pits start to take place before they expand and

propagate and uniform corrosion can be seen along the steel bars. The pits are usually initiated by

the chloride attacks at the weak spots on the bar’s surface, which creates an electrochemical

potential difference between the anodes and cathodes. In micro-cells, both the cathodes and anodes

are very close as shown in Figure 2.2a. Macro-cell reactions can occur over large distances on the

surface of the steel bar or even between two different bars. In macro-cells, the typical pits are

separated by large passive area of steel that acts as small concentered anodes (Eq. (2.1)) that are

surrounded by large cathodes as shown in Figure 2.2b (Eq. (2.2)).

Figure 2.2: A schematic of microcell and macrocell types of corrosion [5]

a) Microcell b) Macrocell

9

2.2.2 Accelerated Corrosion Process

A significant amount of experimental work has been carried out to investigate the influence of

corrosion of steel reinforcement on the structural performance and durability of concrete structures

[3,18,19]. Most researchers used accelerated corrosion techniques to achieve the desired level of

corrosion damage within reasonable amount of time. Table 2.1 summarizes some of previous

accelerated corrosion tests that have been conducted.

Accelerated corrosion can be used to simulate natural corrosion without remarkable change in

the structural response that would be encountered due to natural corrosion [20–22]. The corrosion

process is accelerated by electrical polarization of the steel reinforcement. Chloride salt is usually

used to activate the corrosion process. Some researchers added chloride salts to the concrete mix

[10,20,23] while others partially submerged the test specimens in a salted solution [18,22,24,25].

During accelerated corrosion, the steel bar is connected to the positive terminal of a power supply

to enforce the bar to act as anode and consequently Fe2+ ions start to dissipate from the surface of

the bar. Embedded stainless steel bars or tubes or external plates made of copper or stainless steel

are then connected to the negative terminal of the power supply to act as cathode. Both methods

are presented in Table 2.1.

The level of corrosion-damage of steel bars can be quantified as the percentage of the mass of

the steel bars lost to rust. The steel mass loss due to corrosion can be measured experimentally or

predicted theoretically. Measuring the mass of lost metal, after the completion of the tests, is the

most widely adopted measure of corrosion levels in laboratory tests. The actual mass loss can be

determined according to ASTM G1-90 [26] standard, which provides mechanical, chemical, and

electrolytic techniques to remove corrosion from the steel bars.

On the other side, the mass loss of steel bars can be predicted theoretically using Faraday’s law

[18,22,24,27]. This law relates the mass loss to the corrosion current and to the time of exposure

to corrosion as follows:

m =I t a

n F Eq. (2.6)

where m = mass loss (in grams); I = intensity of the impressed current (in Ampere); t = the time

of corrosion (in seconds); a = the atomic mass of iron (55.847gm); n = the number of electrons

10

transferred during the corrosion reaction (n = 2 in case of iron); and F = Faraday’s constant, which

equals to 96 500 C/equivalent.

For practical application, the current density, i, instead of the current intensity, I, as follows:

m =(S×𝑖 ) t a

n F Eq. (2.7)

where S is the surface area of the corroded steel and i is the impressed current density.

Table 2.1: Summary of some previous accelerated corrosion tests

Study Specimen type Current density

(μA/cm2) Cathode type Corrosion environment

Almusallam et al.

(1996) [25] Bond pull-out 10400 External stainless-

steel plate

Constant immersion in

3% NaCl solution

Alonso et al. (1998)

[28] Bond pull-out 3, 10, or 100 External stainless-

steel plate

Concrete cast with 3%

NaCl by weight of

cement

Mangat and Elgarf

(1999) [24] Beams 1000, 2000, 3000,

or 4000

External copper

plate

Constant immersion

3.5% NaCl solution

Masoud et al.

(2001) [10] Beams 150

Internal stainless-

steel bar


NaCl by weight of

cement

El-Maaddawy and

Soudki (2003) [20] Concrete prisms

100, 200, 350, or

500

Internal stainless-

steel bar


NaCl by weight of

cement

Malumbela et al.

(2010) [21] Beams 189 External stainless-

steel bars

4 days immersion in 5%

NaCl solution followed

by 2 days drying

Mancini et al.

(2014) [23] Concrete prisms 200 External stainless-

steel plate


NaCl by weight of

cement

Ou and Nguyen

(2016) [22] Beams 600 External copper

plate

Constant immersion

5% NaCl solution

In natural corrosion, the measured corrosion current density is in the range of 1 to 100 μA/cm2

[28–31] while the impressed current density used in accelerated corrosion ranges between 3 and

10400 μA/cm2 [25,30]. The density of the impressed current should be kept constant during the

accelerated corrosion to ensure uniform corrosion mechanism [20]. Much research has been

carried out to investigate the optimal current density that should be used in accelerated corrosion

techniques to simulate natural corrosion within reasonable amount of time [24,27,28]. It has been

reported that the use of a current density higher than 250 μA/cm2 in accelerated corrosion increases

11

the corrosion cracks width [32], decreases the bond strength [18,27], and may affect the

mechanical properties of the steel bars [19,22] in comparison to those measured during natural

corrosion at similar mass losses. These observations were attributed to the change of the

composition of the corrosion products due to the use of high current densities [27]. El-Maaddawy

and Soudki (2003) [20] concluded that varying the current density level between 100 and 200

μA/cm2 had no impact on the concrete strain response. However, increasing the current density

above 200 μA/cm2 caused a significant increase in the concrete strain response and the widths of

corrosion cracks due to higher concentrations of corrosion products around the steel bars. In

addition, it was concluded that the use of current densities less than 200 μA/cm2 would be small

enough to obtain corrosion-damage similar to that reported in the field.

2.2.3 Effect of Steel Corrosion on Concrete Structures

Corrosion of steel reinforcement in concrete structures damages the reinforcement itself, the

surrounding concrete, and ruins the composite action between the steel bars and concrete.

Corrosion significantly affects the tensile behavior of the steel bars. The external surfaces of the

corroded bars significantly change due to the formation of irregular pits. Corrosion alters the shape

of the bar cross-section, which varies randomly along the corroded length. It also reduces the cross-

sectional area of the bar and consequently its load-carrying capacity. The non-uniform distribution

of corrosion pits cause stress concentration, which reduces the residual strength and ductility of

the corroded bars compared to that of the uncorroded ones having the same cross section [33,34].

The level of corrosion-damage (i.e. the mass loss of the bars) and the distribution of pits along the

length of the corroded bars significantly influence their stress-strain response in both tension and

compression [33]. For the same corrosion mass loss, corrosion has a more significant effect on the

tensile behavior of plain bars than that of deformed bars and has a more pronounced impact on the

smaller bars than on the larger bars [2].

In an interesting study by Ou et al [22], the tensile behavior of naturally-corroded and artificially-

corroded steel bars were compared. The current density used was 600 μA/cm2. The test results

indicated that the tensile strength and ductility of both naturally- and artificially-corroded bars

were decreased by increasing the level of corrosion damage. The yield and ultimate tensile

strengths of artificially-corroded bars were similar to those of naturally-corroded bars. However,

the ultimate strains of the naturally-corroded bars were smaller than those of the artificially-

12

corroded ones. These observations indicated that corrosion pits were more uniformly distributed

in artificially-corroded bars than in naturally-corroded ones.

Corrosion damage has a more pronounced impact on the fatigue behavior of the corroded steel

bars rather than their tensile behavior [35,36]. Corrosion of steel bars significantly decrease their

fatigue life under tensile and compressive cyclic loads [33]. In addition, corroded bars fail in a

very brittle manner by sudden fatigue rupture. The fatigue life of corroded bars decreases

dramatically with the increase of corrosion level and the fatigue stress range [19].

On the other side, the expansion of the corrosion products results in expansive stresses in the

concrete surrounding the corroded bars. Once these stresses exceed the concrete tensile strength,

cracking and spalling of concrete cover occur. Moreover, corrosion products work as an isolation

layer at the steel/concrete interface, which deteriorates bond and loosens the composite action

between steel and concrete [37]. When deformed bars corrode, the height of their lugs decreases,

and rust fills the inner cracks of the deteriorated concrete. Therefore, their bond performance

become similar to that of plain bars, where just friction is the dominant parameter [38,39]. In

advanced stages of corrosion, the cracked concrete can peel off, and bond between the two

materials can be totally lost. Many researchers have investigated the effect of corrosion on the

bond characteristics at concrete/steel interface. Most studies have found that bond increases

slightly as the steel bar corrodes and decreases rapidly as the corrosion cracks take place [40]. It is

important to note that the reduction in bond are dependent of the thickness of the concrete cover,

the bar diameter, and the presence of stirrups (confinement) [41,42].

2.2.4 Behavior of Corroded-RC Beam under Monotonic Loads

Much research work has been devoted to study the structural behavior of corroded RC beams

under monotonic loads. In the following, a summary of some of the previous work is presented.

In an early study by Almusallam et al. (1996) [43], the structural behavior of RC elements with

different levels of corrosion damage was investigated. The test specimen was 63.5×305×711 mm

and reinforced by 5 steel bars of 6 mm diameter. The whole length of the steel reinforcement was

subjected to an accelerated corrosion process. A sharp reduction in the ultimate strength was

observed as the degree of corrosion increased. For instance, 10 and 28% mass loss resulted in a

reduction in the ultimate strengths of the specimens by 20 and 65%, respectively. The corroded

13

specimens failed by bond-shear failure due to the loss of bond between steel and concrete. Thus,

the ductility of the corroded specimens was significantly reduced.

Vidal et al. (2007) [44] conducted a long-term experimental program to study the structural

behavior of 36 beams of 150×280×3000 mm. The beams were reinforced with two steel bars of 12

or 16 mm diameter. The beams were stored in a chloride environment under service loads for 14

and 17 years to corrode naturally. The test results showed that the steel mass loss was 15 and 30%

after 14 and 17 years respectively. In addition, the flexural cracks during the corrosion exposure

had no significant impact on the rate of corrosion. The inspection of the corroded bars indicated

that neither the distribution of the corrosion pits nor their intensity was uniform along the steel

bars despite the fact that all specimens were subjected to uniform environmental conditions. The

flexural strength and stiffness of the corroded beams were significantly influenced by the level of

corrosion damage.

Torres-Acosta et al. (2007) [45] investigated the behavior of twelve simply-supported RC beams

(100×150×500 mm) with several levels of corrosion damage. The whole length of the steel bars

was subjected to an accelerating corrosion process. Longitudinal cracks parallel to the steel bars

were observed at the end of the corrosion process. Their width increased as the corrosion level

increased. All of the corroded beams showed a more brittle response as the degree of corrosion

damage became higher. The residual flexural strengths also decreased at higher corrosion levels

and were significantly influenced by the depths of the pits on the surface of steel bars.

Gu et al. (2010) [46] investigated the flexural behavior of naturally- and artificially-corroded RC

beams. The naturally-corroded beams with steel mass loss up to 15% failed due to flexural while

increasing the mass loss to 20% changed their mode of failure to rupture of steel bars. On the other

hand, the beams subjected to an accelerated corrosion process up to 25% mass loss failed due to

flexural whereas those subjected to 40% mass loss failed due to rupture of steel bars. This

observation was attributed to the non-uniform characteristics of the corrosion damage and the

significant change in the mechanical properties of naturally-corroded steel bars. The load-carrying

capacities of the corroded beams were reduced due to the reduction in cross-sectional area of the

steel bars and the degradation in their mechanical properties due to corrosion while the decline in

the beams’ stiffness was mainly attributed to the loss of bond at the steel/concrete interface.

14

Malumbela et al. (2010) [47] conducted an experimental study to investigate the influence of

varying the steel mass loss on the ultimate flexural capacity of corroded RC beams. The test

specimen had dimensions of 153×254×3000 mm and were reinforced with three deformed steel

bars of 12 mm diameter. The middle 700 mm of the steel bars were subjected to an accelerated

corrosion process. The results showed that the maximum mass loss was obtained in the middle of

the corroded zone with less mass loss reported at the two edges of the corrosion zone. The use of

a sustained load during the corrosion stage had no significant effect on the rate of corrosion. The

ultimate flexural capacity of the tested beams decreased linearly with the measured steel mass loss.

The ultimate capacity of the beams decreased at a rate of 0.70% per each 1% steel mass loss.

Xia et al. (2012) [48] conducted an experimental study on the effect of corrosion of steel

reinforcement on the stiffness and flexural strength of RC beams. The beams were tested under

four-point load configuration. The tensile reinforcement was corroded within the maximum

moment zone only to provide sufficient anchored length to avoid debonding failure. Nine levels

of corrosion damage (4 to 12% steel mass loss) were achieved using an accelerated corrosion

process. At the end of the corrosion process, longitudinal cracks parallel to the steel bars were

observed. The average and maximum crack widths were in the range of 0.32 to 2.38 mm, and 0.68

to 4.5 mm, respectively. The average and the maximum crack widths increased as the degree of

corrosion increased. The test results showed that the corroded and uncorroded beams had the same

stiffness until 60% of the ultimate load after which the stiffness of the corroded beams significantly

decreased. The residual flexural strength of the corroded specimens decreased as the level of

corrosion increased. The corrosion level changed the failure mode from under-reinforced ductile

failure to brittle collapse due to the rupture of the steel bars at the highest degree of corrosion

damage (12% tensile steel mass loss).

Dang and François (2014) [49] studied the structural behavior of RC beams (280×150x3000) mm

naturally-corroded under harsh environmental conditions over several years. The test observations

revealed that the non-uniform distribution of the corrosion intensity due to the random formation

of pits on the surface of the corroded bars not only affected the flexural strength of the beams but

also changed their ductile failure (steel yielding followed by concrete crushing) to brittle rupture

of the steel bars. The reduction in ultimate capacities corresponding to a given steel mass loss was

lower than the reduction in ultimate deflections. Therefore, it was concluded that the service life

15

of the corroded structure may be limited by the reduction of ductility rather than the reduction of

the load-bearing capacity. This observation was related to the significant change in the mechanical

properties of the corroded steel bars.

2.2.5 Behavior of Corroded-RC Beam under Fatigue Loads

The fatigue life of concrete structures is usually governed by the endurance of the steel

reinforcing bars under repetitive loads and is rarely controlled by concrete [50–52]. Corrosion

damage significantly decreases the fatigue endurance of the steel bars, which significantly reduces

the fatigue life of RC beams. Many researchers investigated the effect of corrosion on the fatigue

performance of RC beams. A summary of some of these investigations is presented as follow:

Yi et al. (2010) [53] carried out a laboratory study to investigate the performance of RC beams

with corroded reinforcement under fatigue loading. The test specimens (150×300×3600 mm) were

reinforced with two deformed steel bars of 20 mm diameter each. The beams were subjected to

accelerated corrosion process to achieve eight degrees of corrosion damage in the range of 3.25 to

11.6 % steel mass loss. All of the beams were subjected to the same level of fatigue load oscillating

between 11 and 52% of the load-carrying capacity of the control (virgin) beam. The test results

revealed that corrosion of the steel reinforcement significantly reduced the fatigue life of the

beams. The fatigue life decreased as the level of corrosion damage increased. For example, the

non-corroded beams safely sustained 2 million fatigue cycles while beams with 3.25 and 11.6%

steel mass loss failed after 626,000 and 89,000 cycles, respectively. In addition, all of the corroded

beams failed due to the sudden rupture of one of the steel bars. The rupture occurred at locations

of corrosion pits where maximum mass losses were encountered.

The performance of corroded RC beams under repeated loads were investigated by Sun et al.

(2015) [6]. The test results showed that corrosion damage of the steel reinforcement had a

significant adverse effect on the fatigue life of the beams. All of the corroded beams failed by

brittle fatigue fracture of one of the steel bars. The fatigue performance of the corroded RC beams

was described by two stages. At the first stage, flexural cracks occurred accompanied by rapid

increase in the mid-span deflection of the beams and reduction in their flexural stiffness. The

second stage represented the largest part of the fatigue life of the beams. In this stage, deflections

and stiffness of the beams became more stable until sudden rupture of the corroded bars occurred.

16

Recently, Zhang et al. (2017) [7] investigated the effect of corrosion on the fatigue behavior of

seven RC beams (120×200×1500 mm). Six out of the seven beams were subjected to an

accelerated corrosion process to achieve different levels of corrosion damage in their tensile steel

reinforcement (2.8 to 15.4 % steel mass loss). The corroded beams were subjected to fatigue with

lower and upper limits of 10 and 60% of the ultimate flexural capacity of the virgin beam,

respectively. The test results showed that the fatigue life of the corroded beams decreased rapidly

with the increase of the level of corrosion damage. Strains in both concrete and steel increased

with the increase of the fatigue life due to the accumulated fatigue damage. The midspan

deflections of the beams increased rapidly at approximately 5% of the fatigue life followed by a

stable stage until sudden failure of the beams occurred. The beams failed by sudden fatigue rupture

of the steel bars. The microscopic scan of the fracture surface of the bars showed that fracture

consisted of three zones namely, the crack initiation region, the crack growth region, and the abrupt

rupture zone.

2.3 FRCM Composites

Fabric-reinforced cementitious matrix (FRCM) system is a composite material consisting of

fabric meshes made of dry fibers embedded in cement-based matrix (serving as binder) as shown

in Figure 2.3. When adhered to concrete or masonry, the FRCM system acts as an externally-

bonded reinforcement. FRCM is considered as the natural evolution of ferrocement and FRP

composites. FRCM have recently emerged as a viable technology for repairing and strengthening

RC and masonry structures [54]. FRCM has been mentioned in the literature review with many

acronyms such as textile-reinforced mortar (TRM) [55–57], textile-reinforced concrete (TRC)

[58], and mineral-based composites (MBC) [59].

17

Figure 2.3: FRCM composite system

The fabric is a manufactured planar textile structure made of fibers, yarns, or both, which is

assembled by various means such as weaving, knitting, tufting, felting, braiding, or bonding of

webs to give the structure sufficient strength and other properties required for its intended use [54].

Figure 2.4 illustrates different fabric configurations. The fabric can be made of high performance

fiber such as carbon, Polyparaphenylene benzobisoxazole (PBO), alkali-resistant glass (AR-glass),

and basalt. Many researchers reported the significant influence of fabric material, geometry, and

configuration on the mechanical properties of the FRCM composite [54,60–63]. The penetration

of the cementitious matrix to the fabric is a fundamental factor for bond between the fabric and

the matrix to develop and consequently for the mechanical properties of FRCM composites.

Therefore, the fabrics consist of meshes with total coverage area less than 2/3 of the total area of

the fabric to provide sufficient matrix penetration [54,64].

Figure 2.4: Different fabric configurations [64]

18

2.3.1 FRCM Acceptance Criteria and Design

FRCM composites should not be produced randomly by selecting and mixing available

commercial materials. Therefore, the developed FRCM systems must go through various tests and

meet certain criteria to evaluate the mechanical properties, bond strength, durability, and fire

resistance of the composite to be accepted for structural applications. The ICC Evaluation Services

(ICC-ES) proposed guidelines for the test methods to evaluate the properties of FRCM composites

as well as their acceptance criteria [65] . The International Building Code (IBC) (Section 104.11.1)

[66] requires a product report that evaluates the characteristics of the FRCM system according to

AC434 guidelines [65] to be approved for strengthening and repair applications.

The American Concrete Institute (ACI) has developed design guidelines for the structural

applications of FRCM composites (ACI-549.4R, 2013) [54]. This document provides the

necessary tools for the design and use of FRCM systems. It covers background information, FRCM

material properties, and design guidelines for axial, flexural, and shear strengthening applications.

It is important to note that the design methodology provided ACI-549.4R was examined and

validated through the analytical investigation of this thesis.

2.3.2 FRCM Tensile Characterization

FRCM composites are typically applied to the tension face of RC members to work as

supplemental tensile reinforcement. Thus, characterizing their tensile behavior is a fundamental

parameter for design. Much research was carried out to investigate the behavior of FRCM coupons

under tensile loading [67–70]. The results of some of these investigations can be summarised as

follows:

The behavior of FRCM coupons under tensile load included three distinct stages namely,

uncracked stage (State I), cracking stage (State IIa), and crack growth stage (State IIb) [67,70] as

shown in Figure 2.5. The actual and idealized stress-strain response of an FRCM coupon under

tensile load is also shown Figure 2.5. The first stage represents the uncracked elastic behavior of

FRCM in which the stiffness of the system is approximately equal to that of the cementitious

matrix. In the second stage, the initiation and the formation of cracks within the matrix result in a

significant loss of stiffness. The length of this part of stress-strain response is mainly dependent

on the bond quality at the fabric/matrix interface and the mount of fibers used. In the third stage

19

(State IIb), the matrix cracks continue to grow and widen and the applied load is fully carried by

the fabric until failure takes place due to their rupture or slippage within the matrix. It is important

to note that the tensile behavior of FRCM owns the advantage of pseudo-ductility due to the

combination between micro-cracking and cracking mechanisms [68,69].

Figure 2.5: Actual and idealized stress-strain curve of FRCM coupon in tension [67]

It can be depicted from the actual stress-strain curve (Figure 2.5) that the slope of the cracking

stage (State IIa) is similar to that of the crack growth stage (State IIb). The transition from the

uncracked stage to the cracking stage is called the bend-over-point (BOP) or the transition point

[67,70]. Therefore, the actual tensile stress-strain curve of FRCM coupon could be idealized to a

simple bi-linear curve with a bend-over point [67,70] as shown in Figure 2.5. The tensile behavior

of FRCM systems can be characterized by the following parameters: the tensile modulus of

elasticity of the uncracked specimen, 𝐸𝑓∗; the tensile modulus of elasticity of the cracked specimen,

𝐸𝑓; the ultimate tensile strain, ԑ𝑓𝑢; the tensile strain corresponding to the transition point, ԑ𝑓𝑡; the

ultimate tensile strength, 𝑓𝑓𝑢, and the tensile stress corresponding to the transition point, 𝑓𝑓𝑡 [54].

Table 2.2 gives the characteristics of PBO-FRCM and C-FRCM coupons that were used in this

study.

20

Table 2.2: Tensile characteristics of PBO-FRCM and C-FRCM coupons [65]

FRCM property Symbol

PBO-FRCM Carbon-FRCM

Mean STD Mean STD

Uncracked tensile modulus of

elasticity, GPa

𝐸𝑓∗ 216 65 74 19

Cracked tensile modulus of elasticity,

GPa

𝐸𝑓 18 2 12 3

Tensile stress corresponding to the

transition point, MPa 𝑓𝑓𝑡 54 12 66 7

Tensile strain corresponding to the

transition point, %

ԑ𝑓𝑡

0.0172 0.0044 0.1020 0.0449

Ultimate tensile stress, MPa 𝑓𝑓𝑢

241 11 150 8

Ultimate tensile strain, % ԑ𝑓𝑢

1.7565 0.1338 1.000 0.1405

Note: Coupon tested with 6 in (150 mm) long tabs.

STD: Standard Deviation

2.3.3 Bond Behavior of FRCM Composites

Understanding the stress transfer mechanisms in FRCM/concrete joints is essential to maximize

their potential as strengthening materials. The fact that FRCM composites consists of dry-fiber

fabric embedded in a mortar prevents the matrix from full impregnating and bonding to each fiber.

Therefore, the weakness of FRCM/concrete joints usually occurs at the fabric/matrix interface

rather than at the matrix concrete interface [71,72] as is typically observed in FRP composites [73].

Large slips between the fabric and the matrix are usually observed during the debonding between

the FRCM layers and the concrete substrate. The debonding mechanism is complicated because of

the so-called telescopic failure that occurs due to the differential slip between the external fibers

and the inner fibers, which makes the composite vulnerable to slipping and/or debonding [69].

The failure mechanisms at the fabric/matrix interface can be modeled as a fiber bundle containing

sleeve and core filaments as shown in Figure 2.6 [74] The bond between the external fibers (the

sleeve fibers) and the matrix as well as the frictional bond between the internal fibers (the core

fibers) govern the bond behavior of FRCM composites and constantly control their strengthening

contribution. The failure at the fabric/matrix interface occurs in three different mechanisms: a)

21

debonding between the external fibers and the matrix, b) slippage between the internal core fibers,

or c) a combination of the two mechanisms [70,74].

Figure 2.6: Failure mechanism of a fiber bundle embedded in cementitious matrix [74]

There are some other factors that affect the bond behavior of FRCM composites. Ombers [75]

reported that increasing the number of FRCM layers changed their mode of failure from fiber

slippage to delamination at the fabric/matrix interface. It was reported that the fabric strain

decreased when debonding occurred. These observations were also consistent with the findings of

D’Amberisi et al. [76]. The bond capacity of FRCM composites increased with an increase in the

bond length. The effective bond length was reported to be in the range of 150-330 mm [76–78].

Nevertheless, varying the width of FRCM composites have no impact on their bond capacity [77].

This was attributed to the independent behavior of the longitudinal fiber bundles as mentioned

earlier. Arboleda et al. (2014) [79] investigated the bond and tensile behavior of carbon and PBO

FRCM composites after exposure to freeze/thaw cycles, high temperature, water vapor, and

immersion in sea water. The test results indicated that the exposure to these hash environments

had no significant impact on the bond and the tensile behaviors of FRCM composites.

2.4 FRCM-Strengthened Beams Under Monotonic Load

Many studies have been conducted to investigate the flexural behavior of undamaged RC beams

strengthened with different FRCM composite systems. Nevertheless, only one study investigated

the feasibly of using FRCM to strengthen corrosion-damaged RC beams. Some of the previous

studies are summarised in the following sections:

Brückner et al. (2006) [58] conducted an experimental work to investigate the feasibility of using

FRCM as externally bonded strengthening technique. RC slabs of 100 mm depth and 1600 mm

22

span were tested under four-point load configuration. Different numbers of FRCM layers were

applied. Improvements in the load carrying capacity, shear loading, and ductility were observed

by increasing the number of FRCM layers used.

Täljsten and Blanksvärd (2007) [59] investigated the behavior of RC slabs strengthened with C-

FRCM and C-FRP systems. Six specimens of 100 mm depth, 1000 mm width, and 4000 mm length

were tested until failure under four-point load configuration. The test results showed the

effectiveness of using FRCM composites as a strengthening technique not only to improve the

ultimate strength of the tested slabs but also to enhance their ductility in comparison to those

strengthened with FRP systems. In addition, small RC beams were tested in bending to investigate

the effect of using different types of cementitious matrices on the bond and ultimate strengths of

the specimens. It was reported that using polymer-modified cementitious mortars enhanced the

bond between the fabric and the matrix as well as the bond at the FRCM/concrete interface, which

caused an increase in the load-carrying capacities of the beams.

Ombres (2011) [80] studied the structural behavior of RC beams strengthened with PBO-FRCM.

Twelve RC beams of 150×250×3000 mm were tested under four-point load configuration up to

failure. The test results showed the efficiency of the PBO-FRCM in increasing both the yielding

and ultimate strengths of the beams. The flexural capacities of the tested beams increased between

10% and 44% in comparison to those of the control beams depending on the amount of FRCM

used. The number of applied plies of FRCM significantly changed the mode of failure from

concrete crushing in case of 1 ply of FRCM to delamination at the concrete/FRCM interface when

2 and 3 plies of FRCM were used. The assumption of strain compatibility between the fabric and

the mortar was verified until 70 to 80% of the failure load, while significant fabric slip in the

mortar was observed at failure. The ductility of the strengthened specimens with 1 ply of FRCM

increased (ductile failure was reported) while it decreased when large number of plies were used

(brittle failure was reported).

D’Ambrisi and Focacci (2011) tested RC beams of 250 mm depth and 400 mm width with

different span lengths (2200 mm and 1600 mm) [81]. The short beams were tested under three-

point bending configuration while the long beams were tested under four-point load configuration.

The effects of using different fiber materials (PBO-FRCM, C-FRCM and CFRP), two different

cement-based matrices (M50 and M75), and three different strengthening configurations

23

(continuous U-shaped strip, U-shaped strip only at the beam ends, and no wrapping) on the

structural performance of the beams were investigated. The experimental results showed that:

a) Using 2 plies of PBO-FRCM with axial stiffness, Af Efu, of about 83% of that of 1 ply of

C-FRP improved the ultimate capacities of the strengthened beams by 30% of that of the

unstrengthened beam (all beams had the same strengthening scheme).

b) Using transverse strengthening changed the mode of failure from sudden detachment of

the strengthening material from concrete to debonding of fibers from the matrix.

Consequently, enhancements in the ultimate capacity and ductility were observed.

Specimens with continuous U-shaped strips showed 18% higher load-carrying capacities

than those with no wrapping.

c) The M50 matrix made of pozzolanic cement, selected silica aggregates, polycarboxylic

water-reducing admixtures, and organic adhesion promoter showed less bond strength

between the fiber and the matrix and between the matrix and concrete than that of the M750

matrix, which was made of composite high-fineness cement binder, adhesion promoter,

inorganic nanoparticles, micro-aggregates, and a new-generation of high-effectiveness

polycarboxylic water-reducing admixtures.

Schladitz et al. (2012) [55] reported on the structural behavior of large scale RC slabs of 230 mm

depth and 6750 mm span strengthened with different number of layers of C-FRCM systems. All

specimens were subjected to four-point load configuration until failure. The test results showed a

significant increase in the ultimate loads as the number of layers increased. It was reported that

using 1 layer and 4 layers of FRCM increased the ultimate loads by 67% and 245%, respectively.

Fabric rupture was observed at failure in all of the strengthened specimens even when high amount

of FRCM layers was used. Significant decrease in deflection was reported when high amount of

FRCM was used.

Si Larbi et al. (2012) [82] conducted an experimental and analytical study to investigate the

effectiveness of alkali-resistant AR glass-FRCM as a strengthening system for RC structures. Five

beams were strengthened and tested under four-point load configuration. The results showed that

FRCM increased the ductility of the strengthened specimens in comparison to the FRP-

24

strengthened ones. On the other hand, using AR glass-FRCM and C-FRP increased the ultimate

capacity by 30% and 80%, respectively.

Elsanadedy et al. (2013) [83] conducted an experimental and numerical investigation to study

the structural behavior of RC beams strengthened with basalt-FRCM systems. Six specimens of

150×200×200 mm were tested under four-point load configuration up to failure. The specimens

were strengthened with continuous U-shaped strips of FRCM. Cementitious and polymer-

modified cementitious mortars were used. The test results were compared with specimens

strengthened with FRP composites. The experimental results approved the effectiveness of using

basalt-FRCM in increasing the flexural strengths of the tested specimens slightly less than FRPs.

Using polymer-modified cementitious matrix provided better bond between the fabric and the

matrix and between the FRCM and the concrete substrate. In addition, 3D finite elements models

were developed to predict the flexural behavior of the tested beams. The basalt-FRCM and FRP

composites were modeled by 4-node shell elements. Bond between composites and concrete was

modeled through the tiebreak surface-to-surface contact definition of LS-DYNA to account for

both normal and shear forces at the interface. A bond-stiffness coefficient, defined as the ratio of

the composite material stiffness to its tensile bond strength, was introduced and recommended not

to be less than 225 for basalt-FRCM to avoid premature debonding failure. The predicted results

were in a good agreement with the experimental results.

Babaeidarabad et al. (2014) tested 18 beams of 1829 mm length, 260 mm depth, and 152 mm

width under three-point load configuration [84]. The test parameters included the concrete strength

(low and high strength), and the number of PBO-FRCM plies used (1 and 4 layers). The specimens

strengthened with 1-ply of fabric showed an increase in the ultimate flexural strength by 32% and

13% for low and high strength concrete, respectively. The specimens failed by fabric slippage

within the matrix while FRCM delamination from the concrete substrate governed the failure of

all specimens strengthened with 4-plies. The flexural strength of specimens with 4-plies of FRCM

was enhanced by 92% and 73% for low and high strength concrete, respectively. A decrease in the

pseudo-ductility of the beams was observed when higher amount of FRCM was used. This work

confirmed the findings of Loreto et al. (2013) [85].

Only one study on the effectiveness of using FRCM composites in repairing corrosion-damaged

RC beams was found in the literature. El-Maaddawy and El Refai (2016) carried out an

25

experimental study to investigate the structural behavior of severely corroded (22% steel mass

loss) T-beams strengthened with either carbon or basalt FRCM systems [86]. The corrosion

damage was restricted to the middle third of the beam span. The FRCM systems were internally-

embedded within the corroded-repaired zone or/and externally bonded along the beam span. The

test results indicated that basalt-FRCM system could not restore the original flexural capacity of

the beam whereas the C-FRCM system restored 109% of the capacity. The use of basalt-FRCM

did not change the brittle failure of the corroded beams (steel rupture) whereas the beams

strengthened with C-FRCM failed by steel yielding followed by FRCM delamination. It was also

concluded that the use of a combination of internally-embedded and externally-bonded C-FRCM

layers was more effective in improving the strength and ductility of the beams than the use of the

same amount of FRCM layers internally embedded within the corroded-repaired region.

2.5 Fatigue and Durability of FRCM-strengthened Beams

Little attention has been paid to investigate the fatigue and durability performance of the RC

beams strengthened with FRCM. The fatigue and long-term performance (after exposure to

environmental conditions) of undamaged RC beams strengthened with PBO-FRCM was

investigated by Aljazaeri and Myers (2016) [87]. Eight beams (203×305×2133 mm) were

strengthened with 1 or 4 layers of PBO-FRCM. The beams were divided into two groups. The

beams of the first group were tested immediately after strengthening while those of the second

group were exposed to varying cycles of freezing and thawing, elevated temperatures, and high

relative humidity before being tested. All of the beams were subjected to fatigue load of 2 million

cycles. The applied fatigue load ranged between 35 and 65% of the specimen ultimate flexural

strength. The results indicated that PBO-FRCM composites enhanced the fatigue performance of

RC beams as well as the residual flexural strength after experiencing fatigue loading. The

enhancement in the fatigue and post-fatigue responses was dependent on the number of layers of

FRCM used. All strengthened beams successfully survived two million cycles without any

evidence of FRCM debonding. Moreover, exposing the PBO-FRCM strengthened beams to high

temperature and humidity did not affect their fatigue performance.

Recently, Pino et al (2017) carried out an experimental study to investigate the fatigue

performance of RC beams strengthened with PBO-FRCM [88]. Ten beams (152×305×1829 mm)

26

were strengthened with 1, 3, or 4 layers of PBO-FRCM. All beams were subjected to 2 million

cycles before testing. The minimum fatigue load was 20% of the specimen theoretical yield load

while the maximum fatigue load varied between 75 and 90% of the specimen theoretical yield

load. The results showed that FRCM improved the fatigue performance of the RC beams. The

fatigue life of the tested specimens decreased with the increase in the maximum fatigue load above

75% of the yield strength. The failure mechanism of the strengthened beams was governed by the

fatigue rupture of the steel bars rather than the failure of FRCM.

2.6 Findings of Literature Review

The literature survey presented in this chapter highlighted in detail the different aspects may

affect the behavior of corroded and FRCM strengthened beams under monotonic and fatigue loads

as well as the long-term performance of FRCM strengthened beams. The findings of the conducted

literature survey can be summarized as follow:

• Accelerated corrosion with impressed electrical current density less than 200 μA/cm2 would

be small enough to obtain corrosion-damage similar to that reported in the field.

• Little emphasis had been given to study the viability of using externally bonded FRCM

composites to strengthened/repair corroded beams at various levels of corrosion damage. In

additions, the effect of FRCM type, amount, and configuration on the flexural response of the

strengthened beams are needed to be quantified.

• To date, no date is available in the literature concerning the durability (post-strengthening

performance) of FRCM strengthened beams after exposure to corrosive environmental

conditions. The effect of FRCM composites on the rate of the corrosion activities was not

investigated.

• The behavior corroded RC beam strengthened with FRCM under fatigue load and the potential

of FRCM composites to restore the fatigue life of corroded beams were not investigated.

• Little effort has been devoted to develop numerical models able to predicted the non-linear

flexural response of FRCM strengthened RC beams.

27

3. Chapter 3

Research Objectives and Methodology

3.1 Research Significance

The conducted literature review presented in Chapter 2 revealed that little research has been

devoted to study the viability of using FRCM to strengthen corrosion-damaged RC elements.

Therefore, the structural behavior of such elements after being exposed to corrosion needs to be

investigated before recommending their use in the field. It is also required to quantify the

strengthening effect of FRCM systems when used in such applications. Corrosion-damaged

structures are often vulnerable to the same deterioration mechanism after repair, which may require

further repair during their service life. To date, no data is available in the literature concerning the

post-repair performance of such structures. In addition, no research has investigated the fatigue

behavior of corrosion-damaged RC beams strengthened with FRCM composites. The lack of

understanding these aspects represents significant obstacles to broad applications of the FRCM

composites. The scope of this study consisted of experimental, numerical, and analytical

investigations. The experimental program was designed to provide a thorough understanding of

the behavior of corrosion-damaged RC beams strengthened with FRCM composites through

design, construction, instrumentation, and testing of 30 large-scale beams as will be detailed in the

following sections. It is thought that the information provided in this thesis is of great value to

designers who wish to use the FRCM composites and also for the development of code provisions

and recommendations.

3.2 Research Objectives

The present study aims at providing technical information about the structural performance of

corrosion-damaged RC beams strengthened with FRCM composites under monotonic and fatigue

loading. The research includes experimental investigations, numerical modeling, and analytical

investigation of the current design code provisions. The objectives of this work can be summarized

as follows:

28

1. To investigate experimentally the viability of using different FRCM systems to strengthen

corrosion-damaged RC beams subjected to various degrees of corrosion.

2. To investigate the effect of FRCM material, number of plies, and strengthening schemes on

the behavior of corrosion-damaged FRCM-strengthened beams.

3. To investigate experimentally the post-strengthening (long-term) behavior of corrosion-

damaged FRCM-strengthened beams under monotonic loads.

4. To investigate experimentally the behavior of corrosion-damaged FRCM-strengthened beams

under fatigue loads.

5. To develop numerical models that are able to predict the non-linear flexural response of

corrosion-damaged FRCM-strengthened beams.

6. To examine the validity of the current design provisions used to predict the flexural strength

of RC members strengthened with FRCM composites on corrosion-damaged FRCM-

strengthened members.

3.3 Methodology

3.3.1 Experimental Work Program

The experimental work included casting and testing thirty large-scale reinforced concrete (RC)

beams of 150 × 250 mm cross section and 2800 mm long. Twenty-seven beams were subjected to

an accelerated corrosion process before being repaired, strengthened, and tested. The main test

parameters included the type of FRCM composites (PBO-FRCM and Carbon FRCM), the level of

corrosion damage (10, 20, and 30% tensile steel mass loss), the amount of FRCM used (1, 2, 3,

and 4 plies of FRCM), the FRCM scheme (end-anchored and continuously wrapped), the short-

and long-term exposure to corrosion, and the type of loading (monotonic and fatigue).

The test matrix of the experimental work program is shown in Table 3.1. The test specimens were

divided into four main groups [A], [B], [C], and [D]. Each group was designed to achieve one or

more of the objectives of the experimental program. The beams were labeled following the X-Y-

Z-N format. ‘X’ represents the beam condition (UU, CU, and CS referring to Uncorroded-

Unstrengthened, Corroded-Unstrengthened, and Corroded-Strengthened, respectively) while ‘M’

29

and ‘F’ refer to Monotonic and Fatigue loading, respectively. ‘Y’ refers to the anticipated

percentage of the tensile steel mass loss due to corrosion. ‘Z’ describes the utilized composite

system (PBO, C, and C-FRP referring to PBO-FRCM, C-FRCM, and Carbon FRP, respectively).

Finally, ‘N’ describes the number of layers and the strengthening scheme of the applied composite

system (I and II refer to the end-anchored and the continuous wrapping schemes, respectively).

3.3.1.1 Group [A]: Control Beams under Monotonic Loading

Group [A] consisted of five control beams that served as benchmarks. Two beams were tested in

flexure up to failure to determine the ultimate capacity of the uncorroded-unstrengthened (virgin)

beam. Three other beams were subjected to accelerated corrosion process for 70, 140, 210 days to

obtain medium (10% mass loss), severe (20% mass loss), and very severe (30% mass loss) levels

of corrosion damage. The beams were then tested in flexure without being strengthened.

3.3.1.2 Group [B]: Short-term Monotonic Behavior of FRCM-strengthened Beams

Group [B] consisted of sixteen beams that were subjected to accelerated corrosion process to

obtain 10, 20, and 30% mass loss in the reinforcing steel bars. The beams were then tested

immediately after being strengthened with different FRCM systems and configurations (short-

term). In addition, one beam was strengthened with C-FRP sheets for comparison. Table 1

illustrates the details of the corrosion level, the strengthening system used, the number of plies,

and the strengthening scheme of each beam. The objective of testing the beams of Group [B] was

to investigate the effect of corrosion-damage level on the flexural performance of the FRCM-

strengthened beams. It also aimed at examining the effectiveness of different FRCM composite

systems in strengthening the corrosion-damaged beams. The test results of Group [B] dictated the

selection of the parameters investigated in the following groups. They also provided guidelines for

the optimal design of FRCM-strengthened beams.

30

Table 3.1: Test matrix of the experimental program

No. Beam Exposure time in days

(Expected mass loss %)

Composite

system (No. of

plies)

Strengthening

scheme

Group [A]: Control beams

1 UUMa 0 (0%) - -

2 UUMb 0 (0%) - -

3 CUM-10% 70 (10%) - -

4 CUM-20% 140 (20%) - -

5 CUM-30% 210 (30%) - -

Group [B]: Short-term beams

6 CSM-10%-PBO-1I 70 (10%) PBO-FRCM (1) I



9 CSM-10%-PBO-2II 70 (10%) PBO-FRCM (2) II


11 CSM-10%-C-2II 70 (10%) C-FRCM (2) II

12 CSM-10%-C-3II 70 (10%) C-FRCM (3) II

13 CSM-10%-CFRP-1I 70 (10%) C-FRP (1) I

14 CSM-20% -PBO-2I 140 (20%) PBO-FRCM (2) I

15 CSM-20% -PBO-4I 140 (20%) PBO-FRCM (4) I


17 CSM-20%-C-3II 140 (20%) C-FRCM (3) II




21 CSM-30%-C-3II 210 (30%) C-FRCM (3) II

Group [C]: Long-term beams

22 CSM-10%:30%-PBO-4I 70 (10%):140 (30%) PBO-FRCM (4) I

23 CSM-10%:30%-PBO-4II 70 (10%):140 (30%) PBO-FRCM (4) II

24 CSM-10%:30%-C-3II 70 (10%):140 (30%) C-FRCM (3) II

Group [D]: Beams under fatigue

25 UUF 0 (0%) - -

26 CUF-20% 140 (20%) - -

27 CSF-20%-PBO-2I 140 (20%) PBO-FRCM (2) I

28 CSF-20%-PBO-4I 140 (20%) PBO-FRCM (4) I

29 CSF-20%-PBO-4II 140 (20%) PBO-FRCM (4) II

30 CSF-20%-C-3II 140 (20%) C-FRCM (3) II

31

3.3.1.3 Group [C]: Long-term Behavior of FRCM-strengthened Beams

Group [C] consisted of three specimens that were subjected to accelerated corrosion process for

70 days (10% mass loss in the steel reinforcing bars). The beams were then strengthened with

PBO- or C-FRCM systems using different strengthening schemes as shown in Table 1. After being

repaired, the beams were exposed to further corrosion for 140 days prior to testing (corresponding

to 30% of steel mass loss). The objective of testing this group of beams was to investigate the long-

term behavior of the FRCM-strengthened beams by simulating a strengthened beam in-service.

The effect of wrapping the beams with FRCM on the corrosion rate and the effect of the harsh

environmental conditions on the integrity of the FRCM systems were also investigated. The test

results of this group of beams were compared to those of the corresponding beams of Group B (the

short-term group).

3.3.1.4 Group [D]: Fatigue Behavior of FRCM-strengthened Beams

Group [D] consisted of six specimens. One beam was neither corroded nor repaired and acted as

a control beam. Five beams were subjected to accelerated corrosion process corresponding to 20%

mass loss in the tensile reinforcement. One of the five corroded beams were tested without being

strengthened while the other four corroded beams were strengthened with PBO- or C-FRCM

systems based on the results obtained from the previous tests. The test parameters included the

number of plies and the strengthening scheme as shown in Table 1. All beams of this group were

tested under fatigue loading until failure occurred. The aim of testing this group of beams was to

examine the effectiveness of FRCM systems in restoring the fatigue life of the damaged-then-

strengthened beams.

3.3.2 Numerical Analysis

A nonlinear finite element (FE) analysis was carried out on the investigated beams. The

numerical analysis included the development of three-dimensional FE models to simulate the

flexural behavior of corrosion-damaged RC beams strengthened with C-FRP and PBO-FRCM

composites. The corrosion damage of the steel bars was represented by a reduction in the cross-

section area of the bars based on the actual mass loss obtained during the tests. The C-FRP

composite system was modeled as discrete reinforcement bonded directly to the beam soffit

without binder while the PBO-FRCM composites were modeled using a more detailed approach

32

that involved modeling the fabric and the matrix layers separately. Interfacial bond stress-slip

models were adopted at the CFRP/concrete and the PBO-fabric/matrix interfaces to simulate the

failure mechanisms observed during the tests. The accuracy of the FE models was validated against

the experimental test results. The validated models provided a valuable supplementary to the

laboratory tests and can be used as a numerical platform to predict the performance of RC flexural

members strengthened with FRP and FRCM composites. In addition, the models were used in a

parametric study to investigate the effect of varying the concrete compressive strengths on the

flexural behavior of the FRCM-strengthened beams.

3.3.3 Analytical Investigation

The analytical investigation included the analysis of the test results using the available design

provisions pertinent to RC members strengthened with FRCM composites. The test results of each

beam were compared to the values predicted using the guidelines of ACI 549.4R-13. Comparing

the analytical with the experimental results aimed at enhancing the current design guidelines by

considering new parameters such as the effect of continuous wrapping on the ultimate flexural

strength of the FRCM-strengthened beams.

33

4. Chapter 4

Corrosion-damaged Reinforced Concrete Beams Repaired

with Fabric-Reinforced Cementitious Matrix (FRCM)

Mohammed Elghazy, Ahmed El Refai, Usama Ebead, and Antonio Nanni

Journal of Composites for Construction, ASCE. Submitted in the revised form: Fabraury 10, 2018

Status: Under review

Résumé

La performance structurelle des poutres en béton armé endommagées par la corrosion et réparées

avec la matrice cimentaire renforcée de fibres (MCRF) a été étudiée. Onze poutres à grande échelle

ont été construites et testées en flexion en configuration de charge à quatre points. Neuf poutres

ont été soumises à un processus de corrosion accélérée pendant 70 jours afin d’obtenir une perte

de masse moyenne de 12,6% dans les barres d'armature en acier, tandis que deux autres poutres

ont été testées comme poutres témoins. Une autre poutre corrodée a été réparé avec les polymères

renforcés de fibres de carbone (PRFC) avant d'être testée. Les paramètres d'essai comprenaient le

nombre de couches de fibres (1, 2, 3 et 4), le schéma de réparation de MCRF (couches ancrées aux

extrémités et couches continues sous forme U) et les matériaux MCRF [carbone et

polyparaphénylène benzobisoxazole (PBO)]. Les résultats des tests ont montré que la corrosion

réduisait légèrement les résistances de plastification et ultime des poutres. L'utilisation de MCRF

a augmenté la capacité ultime des poutres corrodées entre 5% et 52% et leur résistance de

plastification entre 6% et 22% de celles de la poutre vierge non-corrodée. Les poutres réparées

avec des bandes de MCRF en U ont montré une capacité et une ductilité plus élevées que celles

réparées avec des couches ancrées aux extrémités ayant un nombre similaire de couches. Un gain

élevé de la capacité de flexion et un faible indice de ductilité ont été rapportés pour les poutres

avec une plus grande quantité de couches MCRF. Un nouveau facteur a été incorporé dans les

équations de conception de l'ACI-549.4R-13 pour tenir compte du schéma MCRF.

Mots-clés des auteurs : Corrosion; Matrice cimentaire renforcée de fibres; Flexion; Béton armé;

Réparation; Renforcement.

34

4.1 Abstract

The structural performance of corrosion-damaged reinforced concrete (RC) beams repaired with

fabric-reinforced cementitious matrix (FRCM) was investigated. Eleven large-scale RC beams

were constructed and tested in flexure under four-point load configuration. Nine beams were

subjected to an accelerated corrosion process for 70 days to obtain an average mass loss of 12.6%

in the tensile steel reinforcing bars while two other beams were tested as controls. One corroded

beam was repaired with carbon fiber-reinforced polymer (C-FRP) before testing for comparison.

The test parameters included the number of fabric plies (1, 2, 3, and 4), the FRCM repair scheme

(end-anchored and continuous U-wrapped strips), and FRCM materials [carbon and

polyparaphenylene benzobisoxazole (PBO)]. Test results showed that corrosion slightly reduced

the yield and ultimate strengths of the beams. The use of FRCM increased the ultimate capacity of

corroded beams between 5% and 52% and their yield strength between 6% and 22% of those of

the uncorroded virgin beam. Beams repaired with U-wrapped FRCM strips showed higher capacity

and higher ductility than those repaired with the end-anchored bottom strips having similar number

of layers. A high gain in the flexural capacity and a low ductility index were reported for specimens

with high amount of FRCM layers. A new factor was incorporated in the design equations of the

ACI-549.4R-13 to account for the FRCM scheme.

Authors’ keywords: Corrosion; Fabric-reinforced cementitious mortars; Flexure; Reinforced

concrete; Repair; Strengthening.

4.2 Introduction and Background

Corrosion of steel reinforcement is one of the main causes of the deterioration of reinforced

concrete (RC) structures. Corroded structures suffer from loss of cross section of the bars, bond

deterioration, and concrete spalling, which can jeopardize the structure’s safety [44,45,48]. Several

techniques had been adopted to repair and strengthen corroded structures, with the use of

externally-bonded steel plates and more recently the epoxy-bonded fiber-reinforced polymers

(FRP) being the most common techniques.

Numerous studies have documented the advantages of using FRP as repair materials for

corrosion-damaged structures [89–91]. However, concerns about the poor fire resistance of epoxy

35

[16], the incompatibility with the concrete substrate [92], and the loss of ductility of the

strengthened structures [93] have been widely reported. In a desire to overcome these drawbacks,

the fabric-reinforced cementitious matrix (FRCM) systems, with their cement-based adhesives,

have been introduced as a promising alternative to the FRP systems [55–59]. FRCM systems are

characterized by their lightweight, high tensile strength, corrosion resistance, and ease of

application. More importantly, the mortars in the FRCM composites act as barriers against chloride

ions penetration, which may protect the reinforcing bars from corrosion. Their mechanical

properties are strongly influenced by the fabric’s material and geometry and the ability of the

cementitious matrix to impregnate the fabric. The bond strength at the fabric/matrix interface and

at the FRCM composite/concrete interface greatly affect the performance of the strengthened

element [74].

Several studies have reported on the use of FRCM in strengthening RC flexural

[16,56,58,92,94,95]. Most of these studies have related the improvement in the performance of the

strengthened elements to the fabric type and the number of layers used. Parameters such as the

FRCM strengthening scheme and the axial stiffness of the FRCM system used were rarely

reported. D’Ambrisi and Focacci (2011) [81] reported 6 to 46% gain in the load-carrying capacity

of RC beams strengthened with carbon and PBO-FRCM systems. Beams strengthened with PBO-

FRCM performed better than those strengthened with C-FRCM due to the superior bond

characteristics of the former system at the fabric/matrix interface. The use of polymer-modified

cementitious matrix in the PBO-FRCM improved the bond of the fabric to the matrix and

consequently increased the ultimate capacity of the strengthened beams.

In another study, Loreto et al. (2013) [85] reported an increase between 35 and 112% in the

flexural capacity of RC slabs strengthened with PBO-FRCM depending on the volume fraction of

the fabric used. The authors reported that increasing the number of FRCM layers reduced the

ductility of the strengthened slabs. Slabs strengthened with one ply failed due to slippage of the

fabric within the matrix whereas those repaired with four plies failed by fabric delamination at the

fabric/matrix interface. These results were also confirmed by Babaeidarabad et al. (2014) [84].

In a comparison between the performance of FRCM- and FRP- strengthened beams, Elsanadedy

et al. (2013) [83] reported that basalt-FRCM systems were less effective than carbon-FRP systems

36

in enhancing the flexural strength of the beams, yet the FRCM-strengthened beams showed more

ductility at ultimate.

On the other hand, the feasibility of using FRCM systems to strengthen corrosion-damaged

concrete structures have received little attention. The challenges in repairing corroded RC elements

are two-fold, namely the absence of a sound concrete substrate due to corrosion and the durability

of the repair system should corrosion resumes. To the authors’ knowledge, only two studies [86,96]

have documented the effectiveness of using FRCM systems to restore the ultimate capacity and

serviceability of corroded beams. El-Maaddawy and El Refai (2016) [86] reported on the flexural

response of T-beams repaired with carbon and basalt FRCM systems after a mass loss of 22% in

their tensile reinforcement due to corrosion. It was concluded that the basalt-FRCM system could

not restore the original flexural capacity of the beam whereas the C-FRCM system restored 109%

of the capacity. The authors reported that the use of a combination of internally-embedded and

externally-bonded C-FRCM layers was more effective in improving the strength and ductility of

the beams than the use of the same amount of FRCM layers internally embedded within the

corroded-repaired region.

This paper reports the results of the flexural tests conducted on corrosion-damaged RC beams

repaired with FRCM systems. The test program included the type of the FRCM used (carbon and

PBO), the FRCM reinforcement ratios (represented by the number of fabric plies bonded to the

concrete substrate, namely 1, 2, 3, and 4 plies), and the FRCM repair scheme (end-anchored and

U-wrapped strips). The paper also reports on the failure modes, the load-carrying capacities, the

ductility, and the straining actions at different stages of loading of the tested beams. Theoretical

formulations are also presented to predict the flexural response of the beams.

4.3 Experimental Program

Eleven large-scale RC beams were constructed and tested as follows: two specimens were neither

corroded nor repaired (UU), one specimen was corroded but not repaired (CU), seven specimens

were corroded then repaired with different FRCM systems, and one specimen was corroded and

repaired with carbon-FRP (CFRP) sheets. The test matrix is shown in Table 4.1.

The beams were labeled following the X-Y-Z format. ‘X’ represents the beam condition (UU,

CU, and CR referring to Uncorroded-Unrepaired, Corroded-Unrepaired, and Corroded- repaired,

37

respectively). ‘Y’ denotes the number and type of the FRCM layers applied (1P, 2P, 4P, 2C, 3C

and 1FRP referring to one layers of PBO-FRCM, two layers of PBO-FRCM, four layers of PBO-

FRCM, two layers of C-FRCM, three layers of C-FRCM, and one laminate of Carbon FRP

respectively). Finally, ‘Z’ describes the FRCM strengthening schemes (I and II) as will be detailed

in the following sections.

4.3.1 Test Specimen

The test specimen was 2800 mm long with a 150×250 mm rectangular cross section. All beams

were reinforced with 2-15M deformed bars at the bottom (As = 400 mm2) and 2-8M deformed

bars at the top (As' = 100 mm2). The tensile reinforcement ratio was kept at 1.067% that represents

a typical ratio associated with under-reinforced RC section. All of the specimens had a constant

moment span of 800 mm and two shear spans of 880 mm. The shear spans were reinforced with

10M deformed stirrups spaced at 100 mm to avoid a premature shear failure. A hollow stainless-

steel tube with external and internal diameters of 9.5 mm and 7 mm, respectively, was placed at

80 mm from the specimen tension face to act as cathode during the accelerated corrosion process.

Typical dimensions and reinforcement details of the test specimen are shown in Figure 4.1.

38

Table 4.1: Test matrix

*SY-CC = Steel Yielding followed by Concrete Crushing; FD = Fabric Delamination; FS

= Fabric Slippage; MC-SFM = Matrix Cracking with Separation of Fabric within the

Matrix; LR = C-FRP Laminate Rupture.

Figure 4.1: Typical dimensions and reinforcement details of the test specimen (all dimensions in

mm)

Specimen Average mass loss

(%)

𝜌𝑓

(%)

𝐾𝑓 = 𝜌𝑓𝐸𝑓

(MPa)

Mode of Failure*

UUa, UUb - - - SY-CC

CU 12.9 - - SY-CC

CR-1P-I 13.7 0.02 24.1 SY-CC

CR-2P-I 13 0.04 48 FD

CR-4P-I 12.6 0.08 95.3 FD

CR-2P-II 13 0.04 48 FS

CR-4P-II 12.3 0.08 95.3 FS

CR-2C-II 12.5 0.125 93.5 MC-SFM

CR-3C-II 12.1 0.185 139.1 MC-SFM

CR-1FRP-I 11.6 0.153 99.7 LR

39

4.3.2 Accelerated Corrosion Aging

Salt (NaCl) measured as 5% of the cement weight was added to the concrete mix used to cast the

middle-bottom of the corroded specimens with a height of 100 mm (Figure 4.1). Corrosion of the

main reinforcement was localized in the middle 1200 mm of the beam’s span. The accelerated

corrosion process was induced by using a power supply to impress a constant electrical current of

380 mA on the tensile steel bars. The applied current corresponded to a current density of 180

µA/cm2. The invention of this process was produce the desired corrosion damaged within

reasonable amount of time but without remarkable change in the structural response that would be

encountered due to natural corrosion [21]. As there is no laboratory testing standards for the

accelerated corrosion of RC specimens, the current density was chosen based on a previous study

by El Maaddawy and Soudki (2003) [20] that limited the maximum impressed current intensity to

200 µA/cm2 to represent well the natural corrosion in terms of its resulting products. It is also

important to note that this current density is larger than the current densities in real structures which

normally ranged between 0.1 and 100 µA/cm2 [29]. Therefore, this area still needs extensive

research and standardization. During the accelerated corrosion process, the bottom reinforcement

acted as anode whereas the stainless-steel tube acted as cathode and the salted concrete acted as

electrolyte.

The test specimens were electrically connected in series to ensure that the induced current was

uniform in all specimens (Figure 4.2). With applying the anodic current, all specimens were

subjected to wet-dry cycles that consisted of 3 days wet followed by 3 days dry in a large

environmental chamber. The wet-dry cycles provided water and oxygen necessary for the

corrosion process. In this study, a 10% mass loss in the reinforcing bars was anticipated to

represent moderate corrosion damage which commonly encountered in real structures. According

to Faraday’s law that relates the mass loss to the electrical current and the exposure duration, the

duration of corrosion exposure required to achieve this mass loss was 70 days.

40

Figure 4.2: Specimens connected in series inside the corrosion chamber

4.3.3 Materials

Two types of ready mix concrete namely, normal and salted with the same water/cement ratio

were used to cast the beams. Standard concrete cylinders (150×300 mm) were prepared from each

concrete batch and were tested in compression on the day of testing at least after 28 days of casting.

Table 4.2 lists the compressive strengths of both mixes. Prior to FRCM application, the corroded

beams were repaired using local commercial cementitious repair mortar (Sikacrete-08SCC) having

a compressive strength of 55.4 MPa (standard deviation of 5 MPa) and flexural strength of 3.4

MPa (standard deviation of 0.3 MPa) as determined by the authors. The yield strengths of the

longitudinal reinforcing steel bars of diameter 15 and 8 mm were 466 MPa (with a standard

deviation of 4.2 MPa) and 573 MPa (with a standard deviation of 17.7 MPa), respectively, as tested

by the authors.

Two commercial FRCM systems (PBO and carbon) in addition to carbon-FRP composites were

used to strengthen the corroded specimens (Figure 4.3). The fabric properties in the primary

direction as reported in the manufacturers’ data sheet are shown in Table 4.3. The PBO fabric

consists of an unbalanced net of spaced fiber rovings organized along two orthogonal directions

as shown in Figure 4.3a. The associated inorganic cementitious matrix had a compressive strength

of 43.9 MPa (standard deviation of 0.4 MPa) and a flexural strength of 3 MPa (standard deviation

of 0.3 MPa) after 28 days as determined by the authors. On the other hand, the C-FRCM composite

consists of unidirectional net made of carbon-fiber strands oriented in one direction (Figure 4.3b)

41

and impregnated in an inorganic cementitious matrix of compressive strength of 42.1 MPa

(standard deviation of 4.3 MPa) and flexural strength of 3.2 MPa (standard deviation of 0.3 MPa)

after 28 days as determined by the authors. Finally, the carbon-FRP composite consists of

unidirectional carbon fiber sheet (Figure 4.3c) and an epoxy resin. According to the manufacturer’s

data sheet, the composite has a tensile strength of 0.89 GPa, a tensile modulus of 65.4 GPa, and an

ultimate elongation of 1.33%. Table 4.4 lists the properties of the FRCM composite systems as

reported by Ebead et al. (2016) [97].

Table 4.2: Concrete compressive strengths

Compressive strength

(MPa)

Standard deviation

(MPa)

Coefficient of variation

(%)

28-day Normal concrete 37.9 0.8 2

Salted concrete 33.5 1.1 3.2

Testing day Normal concrete 41.8 4.8 11.4


Figure 4.3: Strengthening materials: a) unbalanced PBO fabric, b) unidirectional carbon fabric,

and c) unidirectional carbon fabric

42

Table 4.3: Fabric properties in the primary direction as given in the manufactures’ data sheet

Fabric Area per unit width

(𝐴𝑓) (mm2/m) Tensile strength

(GPa)

Elastic modulus

(GPa)

Ultimate strain

(%)

PBO 50 5.8 270 2.15

Carbon 157 4.3 240 1.75

CFRP 381 3.45 230 1.5

Table 4.4: Mechanical properties of FRCM systems [97]

FRCM system Cracked tensile modulus

of elasticity, Ef (GPa)

Ultimate tensile

strength, ffu (GPa)

Ultimate strain,

εfu (%)

PBO-FRCM 121 1.55 1.4

C-FRCM 75 0.97 1.25

4.3.4 FRCM Equivalent Axial Stiffness

According to the ACI 549.4R (2013) provisions [54], the tensile stress-strain curve of the FRCM

coupon can be represented by a simple bilinear curve as shown in Figure 4.4 The first linear

segment represents the behavior of the FRCM composite prior to cracking and is characterized by

the uncracked modulus of elasticity, 𝐸𝑓∗. The second linear segment represents the cracked

behavior with a reduced cracked modulus of elasticity, 𝐸𝑓. An equivalent axial stiffness, Kf, was

utilized to compare between the two FRCM systems used in this study based on their cracked

elastic modulus and the cross-sectional area of the fabric as given by Equation (4.1):

𝐾𝑓 = 𝜌𝑓𝐸𝑓 = [(𝑁𝐴𝑓)/𝑑𝑓]𝐸𝑓 Eq. (4.1)

where

𝜌𝑓 =𝑁𝐴𝑓

𝑑𝑓

𝜌𝑓, Af, and 𝐸𝑓 are listed in Table 4.1, Table 4.3, and Table 4.4, respectively. The equivalent axial

stiffness, Kf, of each repaired specimen is shown in Table 4.1. It is important to note that for beams

strengthened with the continuously wrapped PBO-FRCM layer (scheme II), the fibers located on

the lateral sides of the beams were neglected in estimating 𝜌𝑓 (and consequently 𝐾𝑓) due to their

minor contribution to the flexural strength of the beams.

43

Figure 4.4: Idealized tensile stress-strain curve of FRCM coupon specimen [54]

4.3.5 FRCM Repair Schemes

Two FRCM repair schemes were utilized in this study, as shown in Figure 4.5. Scheme I

consisted of one or more FRCM flexure plies having 150 mm width (equal to the width of the

beam) and applied to the soffit of the beam over a length of 2400 mm. The fabrics were oriented

so that their primary direction was parallel to the longitudinal axis of the beam. The plies were

anchored at each end using one U-shaped transverse strip of 300 mm width and 200 mm height as

shown in Figure 4.5a. Scheme II consisted of bottom flexural strips similar to those of Scheme I

but wrapped with an additional U-shaped continuous ply along the beam’s span (Figure 5b). The

primary direction of the U-wrapped PBO ply was oriented parallel to the longitudinal axis of the

beams. For instance, the beam CR-4P-II consisted of 3 bottom flexural strips plus one U-shaped

layer, with the primary fibers of all 4 layers running parallel to the longitudinal axis of the beam.

Therefore, the 4 layers of the PBO-fabric contributed to the flexural performance of the beam. On

the other hand, the carbon fabric is a unidirectional fabric. Therefore, the bottom strips of the C-

FRCM were oriented parallel to the longitudinal axis of the beams whereas the U-shaped layer

was oriented in the transverse direction and therefore did not contribute to the flexural behavior of

the beam (Figure 4.5b). For example, specimen CR-3C-II was repaired with 3 flexural strips in the

longitudinal direction plus one U-shaped layer in the transverse direction. Only 3 layers of the C-

FRCM were considered later in the analysis of this beam.

44

Figure 4.5: Repair schemes: (a) Scheme I and (b) Scheme II

4.3.6 Repair Technique

Corroded specimens were repaired before applying the FRCM repair system. Figure 4.6 depicts

the repair procedure. The deteriorated concrete was first removed using a hydraulic hammer. The

corroded steel bars were then brushed, and the beams were repaired using Sikacrete-08SCC

mortar. After 7 days of curing in ambient temperature, sandblasting was used to roughen the

concrete substrate. The beam’s substrate was damped in water for 2 hours before applying the first

layer of the cementitious matrix with a thickness of 3 to 4 mm. Then, the fabric was installed and

coated with a second layer of matrix of similar thickness. The procedure was then repeated until

the specified number of layers was attained.

45

Figure 4.6: Repair procedure: a) removing the deteriorated concrete, b) patch repair, c)

roughening the concrete surface with sandblasting, and d) FRCM application

4.3.7 Test Setup and Instrumentation

All beams were instrumented at mid-span with a 60 mm long concrete strain gauge bonded to

the top surface of the beams and 5 mm steel strain gauges bonded to the tensile reinforcing bars.

The repaired specimens were instrumented with 5 mm strain gauges installed directly on the outer

fabric of the FRCM composite and distributed along the beam span as shown in Figure 4.7. The

beams were tested under four-point loading configuration as shown in Figure 4.1. The load was

applied in displacement control at a rate of 2 mm per minute using a MTS actuator. Beam

deflections were measured by means of three linear variable differential transducers (LVDTs)

located at mid-span and under the point loads. A data acquisition captured the readings of strain

gauges and LVDTs at all stages of loading.

Figure 4.7: Positions of the electrical strain gauges along the outer fabric

46

4.4 Test Observations

4.4.1 Corrosion Cracks and Mass Loss

Due to corrosion, continuous longitudinal cracks parallel to the reinforcing bars were observed

as shown in Figure 4.8 for specimen CU. No concrete spalling was observed. All of the corroded

specimens did not meet the ACI 318-14 [98] service requirements that limits the maximum crack

width in service to 0.40 mm ACI 318-14 [98]. The average and maximum measured crack widths

after corrosion were determined as 0.7 and 1 mm, respectively.

Figure 4.8: Corrosion cracks pattern for specimen CU

Visual inspection of the corroded beams revealed the existence of several corrosion pits randomly

dispersed along the surface of the bars. Six steel coupons, 200 mm long each, were extracted from

each corroded bar after testing. The actual mass losses of the examined bars were determined

according to the ASTM G1-03 standards [26]. The average tensile steel mass loss for each

specimen are listed in Table 4.1. The average, minimum, and maximum steel mass loss determined

for all specimens were 12.6, 11.5, and 13.7%, respectively.

4.4.2 Modes of Failure

The modes of failure of the tested specimens are summarized in Table 4.1 and shown in Figure

4.9 for the tested beams. Beams UU (control) and CU (corroded unrepaired) failed by yielding of

the steel bars followed by concrete crushing (SY-CC). A similar mode of failure was observed in

specimen CR-1P-I as shown in Figure 4.9a. No loss of bond was observed between the PBO-

FRCM and the concrete substrate while loading. The PBO fabric remained intact with its matrix

until crushing of concrete occurred at ultimate. For the other repaired specimens, four different

modes of failure were observed:

47

a) FRCM delamination (FD): this type of failure occurred at the fabric/matrix interface with

complete delamination between the fabric and the first layer of the matrix adjacent to the concrete

substrate (Figure 4.9b). The delamination was caused by the propagation of flexural cracks to this

thin layer of the matrix and the relative deformation between the fabric and the matrix. This mode

of failure was reported for specimens CR-2P-I, and CR-4P-I.

b) Fabric slippage (FS): slippage occurred between the PBO U-shaped fabric and its

cementitious matrix (Figure 4.9c). Cracks were first observed in the matrix of the U-shaped FRCM

layer followed by the gradual slippage of the fabric until the FRCM strengthening action was lost.

This mode of failure was observed in specimens CR-2P-II and CR-4P-II. It should be noticed that

the continuous PBO-U-shaped ply mitigated the FRCM delamination. Therefore, specimens that

failed in this category showed a more ductile behaviour compared to that observed in specimens

that failed due to FRCM delamination.

c) Matrix cracking and fabric separation from the matrix [MC-SFM)]: this type of failure was

reported for specimens with C-FRCM namely, CR-2C-II and CR-3C-II, as shown in Figure 4.9d.

As the applied load increased, progressive cracking in the cementitious matrix associated with the

separation of the carbon fabric from the matrix was observed. Matrix cracking took a web pattern

as shown in Figure 4.9d for the bottom of specimen CR-3C-II. This mode of failure was more

brittle than that observed in the PBO-repaired specimens, which can be attributed to the superior

characteristics of the cementitious matrix of the PBO-FRCM compared to those of the C-FRCM

counterparts.

d) C-FRP laminate rupture (LR): this mode of failure was reported for specimen CR-1FRP-I

(Figure 4.9e). A longitudinal crack initiated at mid span at the laminate/concrete interface followed

by the sudden rupture of the laminate. This mode of failure was consistent with the high strains

recorded in the laminate at ultimate.

48

Figure 4.9: Typical modes of failure: (a) SY-CC in beam CR-1P-I, (b) FD in beam CR-2P-I, (c)

FS in beam CR-2P-II, (d) MC-SFM in beam CR-3C-II, and (e) LR in beam CR-1FRP-I

4.4.3 Load-deflection Response

Load-deflection relationships of the tested beams are shown in Figure 4.10 to Figure 4.12. The

flexural response of the virgin beam (UU), the corroded-unrepaired beam (CU), and the FRP-

repaired beam (CR-1FRP-I) are also shown for reference. The load-deflection curve of specimen

CU indicated that corrosion slightly reduced the load-carrying capacity and stiffness of the beam.

The load-deflection curve of the repaired beams consisted of three segments with two turning

Concrete

crushing

FRCM delamination

Fabric slippage

Matrix crushing

Laminate rupture

49

points indicating the concrete cracking and the yielding of the tensile steel. The flexural response

of the repaired beams was highly dependent on the FRCM repair scheme, its type, and the number

of FRCM layers used.

Figure 4.10 shows the load-deflection relationships of the beams repaired with PBO-FRCM using

Scheme I. All of the beams showed similar stiffness prior to yielding of steel reinforcing bars

indicating the slight influence of the FRCM composite on the flexural response at this stage.

Increasing the number of the PBO plies increased the post-yielding stiffness of the repaired

specimens in comparison to the control ones. Specimen CR-1FRP-I (repaired with one layer of C-

FRP fabric) showed higher post-yielding stiffness than that of specimen CR-4P-I (repaired with

four layers of PBO fabric). However, the later specimen showed slightly higher load carrying

capacity with more ductile mode of failure than the former one.

Figure 4.11 shows the effect of the FRCM scheme on the load-deflection response of the PBO-

repaired beams. Specimens repaired with two and four PBO plies in Scheme II showed a slight

enhancement in the pre-yielding and post-yielding stiffness, which can be attributed to the

enlargement of the beam width and the effect of the continuous U-wrapped strips in delaying the

delamination of the FRCM.

Figure 4.12 compares the load-deflection responses of the Carbon- and PBO-FRCM repaired

beams using scheme II. It can be noticed that specimens repaired with C-FRCM showed higher

post-yielding stiffness than that of their PBO-repaired counterparts. The former specimens

exhibited a sudden drop after reaching the ultimate load whereas specimens repaired with PBO-

FRCM showed a gradual decreasing trend after reaching the ultimate. This can be related to the

brittle mode of failure reported for specimens repaired with C-FRCM.

4.4.4 Strength Analysis

Table 4.5 gives the strength results of the tested beams. The experimental yield, 𝑃𝑦𝑒𝑥𝑝

, and

ultimate, 𝑃𝑢𝑒𝑥𝑝

, loads of all specimens were normalized to those of the virgin specimen. It can be

noticed that corrosion of the main reinforcement reduced the yield and ultimate loads by 8% and

5%, respectively. The reduction in the load-carrying capacity due to corrosion was smaller than

the steel mass loss due to the good anchorage of the bars within the shear zone, which allowed a

tied-arch action to be developed when the specimen approached failure [10].

50

Figure 4.10: Effect of number of PBO-FRCM plies on the load-deflection curves

Figure 4.11: Effect of the repair scheme on the load-deflection curves

51

Figure 4.12: Effect of FRCM materials on the load-deflection curves

Table 4.5: Strength results

Specimen 𝑃𝑦

𝑒𝑥𝑝

(KN)

𝑃𝑢𝑒𝑥𝑝

(KN)

Normalized loads** 𝑃𝑢𝑡ℎ

(KN)


𝑃𝑢𝑡ℎ

ϕ𝑚𝑃𝑢𝑡ℎ

(KN)


ϕ𝑚𝑃𝑢𝑡ℎ

𝑃𝑦𝑒𝑥𝑝


UUa, UUb* 75.1 79.7 1 1 78.5 1.02 70.65 1.13

CU 69.5 76.1 0.92 0.95 70.6 1.08 63.54 1.1

CR-1P-I 71.1 82.9 0.95 1.05 74.9 1.10 67.4 1.23

CR-2P-I 79.5 86.4 1.06 1.08 81.1 1.06 73 1.18

CR-4P-I 83.3 99.6 1.11 1.25 93.3 1.06 84 1.19

CR-2P-II 85.4 102.2 1.14 1.28 81.1 1.26 73 1.4

CR-4P-II 91.3 114.4 1.22 1.44 93.3 1.23 84 1.36

CR-2C-II 79.8 104 1.06 1.30 93.4 1.11 84 1.25

CR-3C-II 90 120.6 1.16 1.52 106 1.14 95.3 1.27

CR-1FRP-I 77.9 96.5 1.04 1.21 - - - -

* Average values reported **Normalized with respect to the yield and ultimate loads of the virgin beam

4.4.4.1 Effect of Number of FRCM Plies on Strength

The use of a single PBO-FRCM layer in specimen CR-1P-I restored 95 and 105% of the yield

and ultimate loads of the virgin beam, respectively. Increasing the number of PBO-FRCM layers

52

further increased the yield and ultimate loads (specimen CR-2P-I restored 106 and 108% and

specimen CR-4P-I restored 111 and 125% of the yield and ultimate loads, respectively). However,

the strength enhancement was not linearly proportional to the added number of FRCM layers.

A similar trend was encountered in specimens repaired with Scheme II. Increasing the number

of FRCM layers enhanced the yield and ultimate strengths of the repaired beams. Specimen CR-

4P-II showed an increase of 22 and 44% of the yield and ultimate loads, respectively, compared

to 14 and 28% for specimen CR-2P-II. Similarly, the use of two layers of C-FRCM in specimen

CR-2C-II increased the yield and ultimate strengths by 6 and 30%, respectively, compared to 16

and 52% for specimen CR-3C-II.

4.4.4.2 Effect of FRCM Repair Scheme on Strength

Scheme II was more effective than Scheme I in restoring the yield and load-carrying capacity of

the repaired beams. This was depicted from the results of the beams repaired with two and four

PBO-FRCM layers. The enhancement in yield load was 6 and 14% for specimens CR-2P-I and

CR-2P-II, respectively. Their corresponding ultimate strengths increased by 8 and 28%,

respectively. The use of four layers of PBO-FRCM with Scheme II in specimen CR-4P-II

increased the yield and ultimate loads by 22 and 44%, respectively, in comparison to 11 and 25%

for specimen CR-4P-I having the same number of PBO-fabric layers.

4.4.4.3 Effect of Axial Stiffness on Strength

Figure 4.13 shows the effect of changing the axial stiffness of the strengthening system, Kf, on

the normalized ultimate load of the tested specimens. Specimens with similar axial stiffness of

their repair system didn’t show similar ultimate capacities. This can be depicted from the results

of specimens CR-2P-I and CR-2P-II having the same axial stiffness of their FRCM system but

with different repair schemes. The former specimen showed a load-carrying capacity of 86.4 KN

versus 102.2 KN for the later one. Similarly, specimens CR-4P-I and CR-4P-II, also having the

same axial stiffness, showed 99.6 KN and 114.4 KN, respectively. This finding was also

demonstrated in specimens CR-4P-II and CR-2C-II having almost similar axial stiffness but

repaired with two different FRCM systems. Specimen CR-4P-II showed a load carrying capacity

of 114.4 KN whereas the specimen CR-2C-II showed a load carrying capacity of 104 KN. This

finding indicates that the axial stiffness of repair system, Kf, should not be used solely to compare

53

the strengthening actions of different FRCM systems without taking into account the material

properties, the fabric architecture, and the repair scheme used. The equivalent axial stiffness of

FRCM systems needs to be further calibrated with more experimental tests to implement other

parameters which is out of scope of this study.

Figure 4.13: Normalized ultimate load versus the equivalent stiffness

4.4.5 Ductility Performance

The ductility index, ΔI, defined as the ratio of the midspan deflection of the beam at ultimate, δu,

to its midspan deflection at yielding, δy, was used to quantify the ductility of the tested specimens.

In general, a higher ductility index means a higher ability of the beam to redistribute moment and

to exhibit large overall deformation and energy dissipation [84]. Table 4.6 lists the deflections at

yielding and ultimate and the ductility indices for all of the tested beams normalized to that of the

virgin beam.

It can be noticed that corrosion of the steel bars increased the ductility index of the corroded

beam by 15%. All beams repaired with PBO in Scheme I restored the ductility of the virgin beam

except beam CR-4P-I that showed a ductility index 13% less than that of the virgin beam. For this

set of beams, increasing the number of PBO plies decreased the ductility of the repaired beam. The

ductility indices of beams CR-4P-I, CR-2P-I, and CR-1P-I were 2.4, 2.8, and 3.0, respectively. On

the other hand, the CFRP-repaired specimen (CR-1FRP-I) did not restore the ductility of the virgin

54

beam and had a similar ductility index of its FRCM-repaired counterpart (CR-4P-I) having similar

axial stiffness.

The set of beams repaired with PBO in Scheme II restored the ductility of the virgin beam. In

fact, these beams showed 2 to 13% increase in their ductility indices as compared to their

counterparts repaired with scheme I. However, increasing the number of the PBO plies in Scheme

II had a less pronounced effect on the ductility index than in Scheme I. Beams CR-4P-II and CR-

2P-II had ductility indices of 2.8 and 2.9, respectively, which indicates that doubling the number

of plies in Scheme II resulted in only 3.5% reduction in the ductility index of the beam.

The ductility indices of the beams repaired with C-FRCM (CR-3C-II and CR-2C-II) were lower

than that of the beams repaired with PBO-FRCM having same repair scheme. Both beams couldn’t

restore the ductility of the virgin beam. Their ductility index was 14 and 22% less than that of the

virgin beam, respectively. This reduction in ductility was attributed to their brittle mode of failure

that was due to the rapid loss of the strengthening action at the fabric/matrix interface. It is

important to note that increasing the number of carbon plies in this set of beams increased the

ductility index of the beam, which is contrary to what has been noticed in the PBO-repaired beams.

This increase was attributed to the increase in the ultimate load of the C-FRCM repaired beams

with similar yielding deflections in comparison to their PBO-counterparts.

4.4.6 Strain Response

Table 4.6 lists the strains measured at midspan in both concrete and the outer fabric at ultimate.

Figure 4.14 and Figure 4.15 show the load-strain curves for specimens repaired with Scheme I and

Scheme II, respectively. Similar to the load-deflection responses of the repaired beams, the load-

strain curves consisted of three segments with two turning points that indicated the concrete

cracking and the yielding of the tensile steel.

55

Table 4.6: Ductility indices and strains at ultimate

Specimen

Midspan

deflection (mm) Ductility index Concrete strains

at ultimate (µ𝜖)

Fiber strains at

ultimate (µ𝜖) δy δu ΔI ΔInorm

**

UUa, UUb* 11.7 32.9 2.8 1.0 -3311 -

CU 10.9 35.4 3.2 1.15 -2992 -

CR-1P-I 11.6 35.2 3.0 1.08 -2711 14921

CR-2P-I 11.8 33.0 2.8 1.0 -2342 8670

CR-4P-I 12.9 31.5 2.4 0.87 -2421 9541

CR-2P-II 11.2 32.0 2.9 1.02 -3491 11261

CR-4P-II 12.7 35.5 2.8 1.0 -2761 9598

CR-2C-II 12.6 27.6 2.2 0.78 -2370 5753

CR-3C-II 12.4 30.1 2.4 0.86 -2262 5991

CR-1FRP-I 12.3 30.4 2.5 0.88 -2351 13772


Figure 4.14: Load-strain curves for specimens with repair Scheme I

56

Figure 4.15: Load-strain curves for specimens with repair Scheme II

Figure 4.14 shows that, prior to yielding, all repaired specimens showed a similar increase in

concrete strains as the applied load increased. This increase in concrete strains continued after

yielding but at different rates depending on the repair system used. Specimen CR-1FRP-I showed

the highest rate of increase in concrete strains when compared to the PBO-repaired ones. On the

other hand, specimen CR-1P-I recorded the maximum tensile strains in the outer fabric of FRCM

(14921 μɛ) as no fabric delamination was observed for this specimen until failure. Specimens CR-

2P-I and CR-4P-I, repaired with two and four plies, failed by FRCM delamination and therefore

recorded low tensile strains in the PBO fabric (8670 μɛ and 9541 μɛ, respectively).

As shown in Figure 4.15, the concrete strains measured in the PBO-repaired specimens were

higher than those recorded in their C-FRCM counterparts. For instance, specimens CR-2P-II and

CR-2C-II recorded concrete strains of 3491 and 2370 μɛ, respectively. It was observed that

concrete strains of the C-FRCM specimens increased at higher rate than that of strains of the PBO-

FRCM specimens. This can be depicted from the strains recorded for specimens CR-2C-II and

CR-3C-II in Figure 4.15. On the other hand, the tensile strains in the C-FRCM at failure was lower

than those in PBO-FRCM. Specimens CR-2C-II and CR-3C-II recorded tensile strains in the outer

fabric at failure of 5753 μɛ and 5991 μɛ, respectively, whereas their counterparts CR-2P-II and

CR-4P-II recorded tensile strains of 11262 μɛ and 9598 μɛ, respectively. These findings were

consistent with the mode of failure of the C-FRCM repaired specimens where premature matrix

57

cracking and fabric separation were encountered. They were also consistent with the measured

ductility indices for these beams.

The distribution of the outer fabric tensile strains along the beam axis are plotted in Figure 4.16

to Figure 4.18 for specimens CR-4P-I, CR-4P-II, and CR-3C-II, respectively, at a service load

equal to 60% of ultimate (0.6 Pu), at the yielding load (Py), and at two post-yielding loads equal

to 0.9 Pu, and Pu. It can be noticed that the strains in the fabric increased with the increase of the

applied load until yielding occurred. Post yielding, a significant increase in fabric strains were

observed, with the maximum increase occurring in the constant moment zone. This finding

indicates that the FRCM system became more effective in resisting the applied loads after yielding

of the steel bars. The repair scheme had marginal effect on the fabric strain profiles as can be

depicted from Figure 4.16 and Figure 4.17.

It is important to note that relying solely on the effective strains in the fabric may be misleading

in predicted computations based on prefect bond between the FRCM and the concrete substrate.

In fact, the slip of the fabric gives origin to a pseudo-strain that can capture the effectiveness of

FRCM strengthening in design. The assumption of prefect bond suggested by the ACI 549.4R-13

committee [54] is a simplification that appears justifiable and easy to implement by engineers.

Figure 4.16: Strain profile in the PBO fabric for specimen CR-4P-I

58

Figure 4.17: Strain profile in the PBO fabric for specimen CR-4P-II

Figure 4.18: Strain profile in the carbon fabric for specimen CR-3C-II

4.5 Theoretical Predictions

The flexural behavior of the tested beams were predicted according to the provisions of the ACI

318-14 building code [98] and the ACI 549.4R-13 committee [54]. Perfect bond was assumed

between the fabric and the cementitious matrix and between the FRCM and the concrete substrate.

A bilinear-elastic behavior of the FRCM repair system was presumed up to failure. The cracked

tensile modulus of elasticity, Ef, of the FRCM system was used after cracking,

59

The FRCM effective tensile strain at failure, 휀𝑓𝑒 , was limited to the FRCM design tensile

strain, 휀𝑓𝑑, as given in Equation (4.2a) (ACI 549.4R [54]). The effective tensile stress in the FRCM

at failure, 𝑓𝑓𝑒 , was calculated in accordance with Equation (4.2b)

휀𝑓𝑒 = 휀𝑓𝑑 ≤ 0.012 Eq. (4.2a)

𝑓𝑓𝑒 = 𝐸𝑓휀𝑓𝑒 Eq. (4.2b)

Strains in concrete, steel reinforcing bars, and FRCM systems were computed in accordance with

Equation (4.3) using the strain compatibility principle as shown in Figure 4.19.

𝜀𝑓𝑒

𝑑𝑓−𝑐𝑢=

𝜀𝑠

𝑑−𝑐𝑢 =

𝜀𝑠′

𝑐𝑢−𝑑′=

𝜀𝑐

𝑐𝑢 Eq. (4.3)

Figure 4.19: Stress and strain distribution at ultimate stage

The nominal flexural strength, Mn, was calculated in accordance with Equations (4.4) as follows:

𝑀𝑛 = 𝑀𝑠 + 𝑀𝑓 + 𝑀𝑠′ Eq. (4.4)

where,

𝑀𝑠 = 𝑇𝑠 ( d − 𝛽1 𝐶𝑢

2 ) Eq. (4.4a)

𝑀𝑓 = 𝑇𝑓 ( d − 𝛽1 𝐶𝑢

2 ) Eq. (4.4b)

𝑀𝑠′ = 𝐶𝑠′ (𝛽1 𝐶𝑢

2− 𝑑′) Eq. (4.4b)

60

𝑀s, 𝑀f, and 𝑀s were calculated with respect to the centroid of the equivalent rectangular stress

block as shown in Figure 4.19. The concrete stress block factors, 𝛽1 and 𝛼1, and the modulus of

elasticity of concrete, Ec, were calculated as follows (ACI 318-14 [98]):

β1 = (4εc

′ −εc(Cu)

6εc′ −2εc(Cu)

) Eq. (4.5)

α1 = (3εc

′ εc(Cu)−[εc(Cu)]2

3β1(Cu)εc′2 ) Eq. (4.6)

𝐸𝑐 = 4700√𝑓c′ Eq. (4.7)

휀𝑐′ = 1.7𝑓𝑐

′/𝐸𝑐 Eq. (4.8)

The force equilibrium was satisfied in accordance with Equations (4.9) and as shown Figure 4.19:

𝑇𝑠 + 𝑇𝑓 = 𝐶 + 𝐶𝑠′ Eq. (4.9a)

Where,

𝑇𝑠 = 𝑅𝑐𝑜𝑟 𝐴𝑠𝑓𝑦 Eq. (4.9b)

𝑇𝑓 = 𝑁𝐴𝑓𝑏𝑓𝑓𝑒 Eq. (4.9c)

𝐶 = 𝛼1𝑓𝑐′𝛽1𝑐𝑢𝑏 Eq. (4.9d)

𝐶𝑠′ = 𝐴𝑠′ 𝐸𝑠휀𝑠

′ Eq. (4.9e)

Where, 𝑅𝑐𝑜𝑟= 1 – average tensile steel mass loss.

Table 4.5 lists the theoretical ultimate loads, 𝑃𝑢𝑡ℎ, for all of the tested specimens. Good agreement

between the experimental and theoretical values was obtained especially for specimens repaired

in Scheme I. However, the capacities of specimens repaired with Scheme II were under-estimated.

The theoretical formulations adopted do not account for the effect of the U-shaped FRCM layers

on the flexural response of the beams. The obtained results suggested the increase of the nominal

capacity, Mn, of FRCM-repaired beams with U-wrapped layers by 10% to account for the scheme

of the FRCM used.

61

4.5.1 Design Provision

According to the provisions of the ACI 549.4R [54], the design flexural strength, 𝑀𝐷, is

calculated in accordance with Equation (4.10). The strength reduction factor, 𝜙𝑚, is given by

Equation (4.11). In addition, the ACI-549 committee limits the increase in the nominal flexural

strength provided by the FRCM reinforcement by 50% of the flexural capacity of the structure

prior to repair. Table 4.5 lists the theoretical design load ϕ𝑚𝑃𝑢𝑡ℎ and the ratio 𝑃𝑢

𝑒𝑥𝑝/ϕ𝑚𝑃𝑢𝑡ℎ. It can

be noticed that applying both the flexural strength reduction factor and the 50% increase limitation

makes the gap between the experimental and design values lager, especially for the specimens

repaired with Scheme II.

𝑀𝐷 = 𝜙𝑚𝑀𝑛 Eq. (4.10)

𝜙𝑚 = {

0.90 for ɛ𝑡 ≥ 0.005

0.65 +0.25(ɛ𝑡−ɛ𝑠𝑦)

0.005−ɛ𝑡−ɛ𝑠𝑦

0.65 for ɛ𝑡 ≤ ɛ𝑠𝑦

for ɛ𝑠𝑦 < ɛ𝑡 < 0.005 Eq. (4.11)

4.6 Conclusions

This study investigated experimentally and analytically the structural performance of corrosion-

damaged RC beams repaired with PBO- and C-FRCM systems. The following conclusions can be

drawn:

• An average mass loss of 12.9% in the tensile steel reduced the yield and the ultimate loads of

the beam by 8% and 5%, respectively. The corroded-unrepaired specimens failed to meet the

provisions of the ACI-318 standards for crack width criteria.

• PBO-FRCM repaired specimen showed slightly higher ultimate load carrying capacities with

more ductile mode of failure than those of C-FRP repaired specimen with similar axial

stiffness and repair scheme.

• Repairing corrosion-damaged RC beams with PBO- and C-FRCM restored 105 to 144% and

130 to 152%, respectively, of the original load-carrying capacity of the virgin uncorroded

beam. The gain in capacity was highly dependent on the number of fabric layers, their

material, and the scheme used.

62

• Beams repaired with PBO-FRCM systems failed in a ductile mode due to either fabric

delamination (repair Scheme I) or fabric slippage within the matrix (repair Scheme II),

whereas beams repaired with U-wrapped C-FRCM systems showed a more brittle failure due

to matrix cracking and complete separation of the fabric.

• Beams repaired with C-FRCM showed higher post-yielding stiffness than that of their PBO-

repaired counterparts. The former beams exhibited a sudden drop after reaching the ultimate

load whereas the later beams showed a gradual decrease after reaching the ultimate.

• Increasing the number of FRCM layers increased the yielding and ultimate loads of the

repaired beams. However, specimens with similar axial stiffness didn’t show similar ultimate

capacities. More tests are required to calibrate the axial stiffness, Kf, to implement parameters

such as the material properties, the fabric architecture, and the repair scheme used.

• U-wrapped FRCM scheme was more efficient than the bottom end-anchored scheme in

increasing the ultimate capacity of the repaired beams. The PBO-repaired beams with scheme

II showed ultimate strengths 15 to 18% more than those repaired with scheme I.

• Beams repaired with PBO-FRCM systems showed a more ductile behavior than their

counterparts repaired with C-FRCM. Most of the PBO-repaired beams restored the original

ductility whereas the C-FRCM repaired beams showed lower ductility than that of the virgin

beam.

• Strain values recorded during the tests indicated that the assumption of prefect bond suggested

by the ACI-549.4R-13 committee [54] is a justifiable simplification for easy implementation

by practicing engineers.

• The theoretical formulations of the ACI-549.4R-13 committee reasonably predicted the

ultimate strengths of the FRCM-repaired beams with Scheme I but underestimated those

repaired with Scheme II. A scheme factor of 1.1 is then proposed while calculating the

nominal strength of beams repaired with U-shaped FRCM.

63

4.7 Notation

The following symbols are used in this paper:

Af = equivalent area of fabric per unit width (mm2/mm)

As = cross-sectional area of tension steel reinforcement (mm2)

𝐴𝑠′ = cross-sectional area of compression steel reinforcement (mm2)

b = width of the beam (mm)

C = compression force provided by concrete (kN)

Cs’ = compression force provided by the compression reinforcement (kN)

cu = distance from extreme compression fiber to neutral axis (mm)

d = distance from top of the beam to the centroid of tension steel (mm)

d’ = distance from top of the beam to the centroid of compression steel (mm)

df = distance from top of the beam to the centroid of fabric reinforcement (mm)

Ef = cracked elastic modulus of the FRCM composite (GPa)

Es = elastic modulus of steel reinforcement (GPa)

Ec = elastic modulus of concrete (MPa)

𝑓𝑐′ = compressive strength of concrete (MPa)

𝑓𝑓𝑒= effective tensile stress in FRCM composite at failure (MPa)

ffu = ultimate tensile strength of FRCM composite (MPa)

fy = yield strength of steel reinforcement (MPa)

MD = design flexural strength (kN-m)

Mf = moment contribution of FRCM reinforcement to flexural strength (kN-m)

Ms = moment contribution of the tensile steel reinforcement to flexural strength (kN-m)

Ms’ = moment contribution of the compression steel reinforcement to flexural strength (kN-m)

Mn = nominal flexural strength (kN-m)

N = number of fabric layers

Rcor = corrosion reduction factor

Ts = tension force in steel reinforcement (kN)

Tf = tension force in FRCM reinforcement (kN)

휀𝑐 = compression strain in concrete (mm/mm)

휀𝑐′ = compression strain of unconfined concrete corresponding to 𝑓𝑐

′ (mm/mm)

64

εcu = concrete strain at ultimate (mm/mm)

휀𝑠′ = tensile strain in compression steel reinforcement (mm/mm)

휀𝑠𝑦 = tensile yield strain of steel reinforcement (mm/mm)

휀𝑡 = the net tensile strain in extreme tensile steel reinforcement at the nominal strength (mm/mm)

휀𝑓𝑑 = FRCM design tensile strain (mm/mm)

휀𝑓𝑒 = effective tensile strain in FRCM composite at failure (mm/mm)

εfu = ultimate tensile strain of FRCM composite (mm/mm)

𝜌𝑓 = fabric reinforcement ratio

𝜅𝑓 = equivalent axial stiffness (MPa)

∆𝐼 = ductility index

β1 = ratio of depth of equivalent rectangular stress block to depth to neutral axis

α1 = multiplier of 𝑓𝑐′ to determine intensity of the equivalent block stress for concrete

𝜙𝑚= strength reduction factor

65

5. Chapter 5

Effect of Corrosion-damage on the Flexural Performance

of RC Beams Strengthened with FRCM Composites


Journal of Composites Structures. Date of acceptance: August 16,2017

( https://doi.org/10.1016/j.compstruct.2017.08.069 )

Résumé

Cet article rend compte du comportement en flexion des poutres en béton armé endommagées

par la corrosion et renforcées par différents systèmes de matrice cimentaire renforcée de fibres

(MCRF). Trois groupes de poutres ont été soumis à une corrosion accélérée pendant 70, 140 et

210 jours pour obtenir une perte de masse théorique dans les barres d'acier de traction de 10%,

20% et 30%, respectivement. Les paramètres d'essai comprenaient le type de fibres (PBO et

carbone), le nombre de couches de MCRF (deux, trois et quatre), et le schéma de renforcement

(couches ancrées aux extrémités et couches continues sous forme U). Les résultats des tests ont

montré que les composites MCRF gouvernaient la défaillance des poutres renforcées plutôt que le

niveau de dommage auquel les poutres étaient soumises en raison de la corrosion. Les résultats des

tests sur les poutres endommagées par la corrosion ont confirmé que la contribution des composites

MCRF compensait de manière significative l'impact de la corrosion sur la résistance. Les poutres

renforcées par MCRF présentaient une augmentation de la résistance qui variait entre 7 et 55% de

celle de la poutre vierge selon le type, la rigidité axiale et le schéma de la MCRF utilisé. Les

poutres renforcées ont montré des indices d'absorption d'énergie qui se situaient entre 111 et 153%

de celui de la poutre vierge. Les formulations théoriques de l'ACI-549.4R-13 ont raisonnablement

prédit les résistances ultimes des poutres renforcées ancrées à l'extrémité, mais ont sous-estimé

celles des poutres renforcées par des couches continues sous forme U.



https://doi.org/10.1016/j.compstruct.2017.08.069

66

5.1 Abstract

This paper reports on the flexural behavior of corrosion-damaged reinforced concrete (RC)

beams strengthened with different fabric-reinforced cementitious matrix (FRCM) composites.

Three groups of beams were subjected to accelerated corrosion for 70, 140, and 210 days to obtain

theoretical mass loss in their tensile steel bars of 10%, 20%, and 30%, respectively. The test

parameters included the fabric type (PBO and carbon), the number of FRCM layers (two, three,

and four), and the strengthening scheme (end-anchored and continuously wrapped). Test results

showed that FRCM composites governed the failure of the strengthened beams rather than the

damage level to which the beam was subjected due to corrosion. The reported load-carrying

capacities of the corrosion-damaged beams confirmed that the contribution of FRCM composites

significantly offset the impact of corrosion damage on strength. FRCM-strengthened beams

exhibited an increase in strength that ranged between 7 and 55% of that of the virgin beam based

on the type, the axial stiffness, and the scheme of the FRCM used. The strengthened beams showed

energy absorption indices that ranged between 111 and 153% of that of the virgin beam. The

theoretical formulations of ACI-549.4R-13 reasonably predicted the ultimate strengths of the end-

anchored strengthened beams but underestimated those continuously anchored beams.

Authors’ keywords: Corrosion; Fabric-reinforced cementitious matrix; Flexure; Reinforced

concrete; Repair; Strengthening.


Corrosion of steel reinforcing bars is inevitable. Despite the stringent provisions specified by

most of existing building codes, corrosion is still being reported in reinforced concrete (RC)

structures due to the continuous exposure to harsh environments, proximity to sea-shores, and the

use of de-icing salt. The transfer of water, oxygen, and aggressive agents such as carbon dioxide

and chloride into concrete leads to corrosion and consequently to concrete cracking, spalling, and

deterioration. Steel corrosion results in reduction in the cross-sectional areas of the bars and the

loss of bond at the steel/concrete interface. These phenomena significantly reduce the structural

integrity of the RC member and may lead to its premature collapse [44,45,48,53]. As a result,

engineers face a big challenge not only to assess the corrosion-damaged structure but also to select

the appropriate strengthening technique.

67

The superior properties of epoxy-bonded fiber-reinforced polymer (FRP) products and their non-

corrosive characteristics have inspired engineers to adopt them in strengthening applications.

Several studies have reported on the effectiveness of using FRP composites in strengthening

corrosion-damaged RC structures [10,90,91]. Although the epoxy bonding agents used with FRPs

are commonly durable and resistant to the harsh environmental conditions, various problems

associated with their performance at high temperature have been reported [16]. In addition, epoxies

have low compatibility with the concrete substrate [92] and can’t be applied on wet surfaces or at

low temperatures. In order to overcome such drawbacks, fabric-reinforced cementitious matrix

(FRCM) systems were introduced as promising alternatives.

FRCM composites consist of one or more fabric mesh made of long dry-woven embedded in a

cement-based matrix that serves as a binder. The fabric may be made of carbon (C), glass (G), or

Polyparaphenylene benzobisoxazole (PBO) while the matrix is an inorganic hydraulic or non-

hydraulic cementitious mortar that holds in place the reinforcing meshes. Many studies have

demonstrated the effectiveness of FRCM composites in enhancing the flexural and shear

performances of damage-free RC beams [57,87,99]. Babaeidarabad et al. (2014) [84] investigated

the flexural performance of RC beams strengthened with PBO-FRCM. The results showed that the

strength gain ranged from 13 to 93% of that of the unstrengthened beams depending on the amount

of FRCM layers used. Fabric slippage within the matrix and FRCM delamination were reported

as two distinct modes of failure. Similar results were reported by Loreto et al. (2013) [85].

Schladitz et al. (2012) [55] also reported that increasing the volume fraction of the fabric used in

strengthening RC slabs increased their load-carrying capacities. However, this gain in strength was

accompanied by a decrease in the ductility of the strengthened slabs. D’Ambrisi and Focacci

(2011) [81] demonstrated that the strengthening effectiveness of FRCM composites was highly

dependent on the type of fabric and the bond between the matrix and the fabric.

The use of FRCM in strengthening corrosion-damaged RC members have been rarely reported.

The challenge in strengthening such members arises from the potential loss of integrity of the

member following the corrosion of its reinforcement. It also arises from the potential loss of bond

between the deteriorated concrete and the new mortar applied. The study conducted by El-

Maaddawy and El Refai (2016) [86] evidenced the feasibility of using basalt and carbon FRCM

systems to strengthen severely corrosion-damaged RC beams that suffered 22% mass loss of their

68

steel bars. It was reported that basalt-FRCM system could not restore the original flexural capacity

of the beams whereas the C-FRCM system restored 109% of the capacity. El-Maaddawy and El

Refai (2016) [86] also reported that the beams strengthened with a combination of internally-

embedded and externally-bonded C-FRCM layers have restored both their strength and ductility.

No results were reported on the effect of the FRCM strengthening scheme or the degree of

corrosion damage on the flexural behavior of the strengthened beams.

The present work is part of a large experimental program that aims at filling the gap in knowledge

on the effectiveness of FRCM composites in strengthening corrosion-damaged RC structures. The

tested beams presented herein were subjected to three levels of corrosion damage prior to

strengthening. FRCM composites having different amounts of fibers, different mechanical

properties, and different schemes were used to strengthen the corrosion-damaged beams. The test

results report on the gain in the yield strengths, the ultimate strengths, and the ductility of the

strengthened beams. The influence of the damage degrees, in addition to other parameters, on the

flexural performance of the strengthened beams is presented and discussed.


The experimental program is summarized in Table 5.1. Data of some of the tested beams were

previously reported in Elghazy et al. [100] and are presented herein for comparison purposes. The

test parameters included the corrosion level, the strengthening scheme, and the number of FRCM

layers used. The beams were subdivided into three groups (A, B, and C) and were subjected to

accelerated corrosion process for 70, 140, and 210 days, respectively. At the end of the corrosion

process, one beam in each group was not strengthened and was used as a benchmark while other

beams were strengthened with the designated FRCM systems and configurations. In addition, two

virgin beams (i.e. not corroded nor strengthened) were used as controls.

The beams were labeled following the X-Y-Z format. ‘X’ represents the beam condition (UU,

CU, and CS referring to Uncorroded-Unstrengthened, Corroded-Unstrengthened, and Corroded-

Strengthened, respectively) and the beam’s group (A, B, and C). ‘Y’ denotes the number and type

of the FRCM layers applied (2P, 4P, and 3C referring to two layers of PBO-FRCM, four layers of

PBO-FRCM, and three layers of C-FRCM, respectively). Finally, ‘Z’ describes the FRCM

strengthening schemes (I and II) as will be detailed in the following sections.

69

5.3.1 Test Specimen and Materials

Figure 5.1 shows the beam geometry and the reinforcement details. All beams were designed

according to ACI 318-14 provisions [98]. All beams were 2.8 m long with 150×250 mm

rectangular cross section. The bottom and top reinforcement consisted of two 15M (diameter 15

mm) and 8M (diameter 8 mm) deformed bars, respectively. The shear spans were reinforced with

10M (diameter 10 mm) deformed stirrups spaced at 100 mm to avoid shear failure. To accelerate

the corrosion process of the bottom steel bars, a hollow stainless-steel tube with an external

diameter of 9.5 mm and a wall thickness of 2.5 mm was placed at 80 mm from the beam soffit to

act as cathode. Salt (NaCl) weighed as 5% of the cement weight was added to the concrete mix

used to cast the middle-bottom of the corroded beams with a height of 100 mm (Figure 5.1). Details

about the accelerated corrosion process are presented in the following section.

Six standard concrete cylinders (150×300 mm) were prepared from the normal and salted

concrete. Table 5.2 lists the compressive strengths of both mixes after 28 days and on the day of

testing. Prior to FRCM application, the corroded beams were repaired using a commercial

cementitious repair mortar (Sikacrete-08SCC) having a compressive strength of 55.4 MPa

(standard deviation of 5 MPa) and a flexural strength of 3.4 MPa (standard deviation of 0.3 MPa)

as tested by the authors. The yield strengths of the longitudinal reinforcing steel bars of diameter

15 and 8 mm were 466 MPa (with a standard deviation of 4.2 MPa) and 573 MPa (with a standard

deviation of 17.7 MPa), respectively, as tested by the authors.

70

Table 5.1: Summary of the test results

*Average values reported **

SY-CC = Steel Yielding followed by Concrete Crushing; FD = FRCM Delamination; FS-PED = Fabric Slippage followed by Partial Debonding; MC-FS =

Matrix Cracking followed by Fabric Slippage

Specimen wavg

(mm)

Avg. Mass

loss (%)

𝜌𝑆

(%)

𝜌𝑓

(%)

𝐾𝑓

(MPa)

𝛽𝑓

(%)

𝑃𝑦

(kN)

𝑃𝑢

(kN) 𝑃𝑦 𝑁𝑜𝑟𝑚 𝑃𝑢 𝑁𝑜𝑟𝑚

𝑃𝑢𝑡ℎ

(kN)

𝑃𝑢

𝑃𝑢𝑡ℎ

εfu

(µϵ)

Mode of

failure**

Virgin beams

UUa, UUb* - - 1.07 - - - 75.1 79.7 1 1 81.9 0.97 - SY-CC

Group A: Theoretical mass loss of 10%

CUA 0.65 12.9 0.93 - - - 69.5 76.1 0.93 0.95 72.31 1.05 - SY-CC

CSA-2P-I 0.72 12.6 0.93 0.04 48 2.58 79.5 86.4 1.06 1.08 83.64 1.03 7743 FD

CSA-4P-I 0.75 12.6 0.93 0.08 95.3 5.11 83.3 99.6 1.11 1.25 94.56 1.05 9442 FD

CSA-4P-II 0.68 12.3 0.94 0.08 95.3 5.1 91.3 114.4 1.22 1.44 100.1 1.14 9526 FS-PFD

CSA-3C-II 0.58 12.1 0.94 0.185 139.1 7.41 87 120.6 1.16 1.51 104.78 1.15 6000 MC-FS

Group B: Theoretical mass loss of 20%

CUB 1 18 0.88 - - 64.5 74.2 0.86 0.93 68.27 1.09 - SY-CC

CSB-2P-I 1.1 19.6 0.86 0.04 48 2.8 71.8 85.6 0.96 1.07 78.2 1.09 8180 FD

CSB-4P-I 1.15 19.4 0.86 0.08 95.3 5.54 79.6 102.6 1.06 1.29 89.38 1.15 10659 FD

CSB-4P-II 0.95 19.5 0.86 0.08 95.3 5.55 80.7 102.9 1.07 1.29 94.6 1.09 8253 FS-PFD

CSB-3C-II 1.15 18.6 0.87 0.185 139.1 8.01 78.8 123.3 1.05 1.55 99.95 1.23 5530 MC-FS

Group C: Theoretical mass loss of 30%

CUC 1.65 22.5 0.83 - - - 64.1 72.2 0.85 0.91 64.68 1.12 - SY-CC

CSC-4P-I 2 22.7 0.83 0.08 95.3 5.78 82.5 102.8 1.1 1.29 86.85 1.18 7339 FD

CSC-4P-II 1.6 21.1 0.84 0.08 95.3 5.66 80.4 111.1 1.1 1.39 93.1 1.19 9653 FS-PFD

CSC-3C-II 1.5 21.5 0.84 0.185 139.1 8.31 75.2 109.3 1 1.37 97.77 1.12 4885 MC-FS

71

Figure 5.1: Test specimen geometry and reinforcement details. (All dimensions in mm)

Table 5.2: Concrete compressive strengths

Compressive strength

(MPa)

Standard deviation

(MPa)


(%)






A DC galvanostatic power supply was used to impress a constant electrical current of 380

milliamps (mA) with an approximate density of 180 µA/cm2. The density level was chosen less

than 200 µA/cm2 to represent natural corrosion encountered in the field and to avoid the bond loss

at the steel/concrete interface as recommended in [20]. All beams were connected in series in a

large environmental chamber as shown in Figure 5.2. Therefore, the reinforcing bars acted as

anodes, the stainless-steel bars acted as cathodes, and the salted concrete acted as electrolyte.

During the corrosion process, the beams were subjected to consecutive wet-dry cycles that

consisted of 3 days wet followed by 3 days dry each.

72

Figure 5.2: Specimens inside the environmental chamber during a dry cycle

The theoretical mass loss of the steel bars due to corrosion was calculated using Faraday’s law

that relates the mass loss to the electrical current and the exposure duration as follows:

𝑚 =𝐼𝑡𝑎

𝑛𝐹 Eq. (5.1)

where m = mass loss (in grams), I = impressed current (in Ampere), t = corrosion duration (in

seconds), a = the atomic mass of iron (55.847grams), n = the number of electrons transferred

during the corrosion reaction (n = 2 for iron), and F = Faraday’s constant (96.500 C/equivalent).

As previously mentioned, three mass losses were considered in this study namely, 10% for beams

of group A, 20% for beams of group B, and 30% for beams of group C, which represented

moderate, severe, and very severe degrees of corrosion damage, respectively.

5.3.3 FRCM Systems

In this study, two commercially available FRCM systems (PBO-FRCM and C-FRCM) were used

to strengthen the corrosion-damaged beams. The PBO fabric consisted of an unbalanced net of

spaced fiber rovings running along two orthogonal directions as shown in Figure 5.3a. The

associated inorganic cementitious matrix had a compressive strength of 43.9 MPa (standard

deviation of 0.4 MPa) and a flexural strength of 3 MPa (standard deviation of 0.3 MPa) as

determined by the authors. The C-FRCM composite consisted of unidirectional net made of

carbon-fiber strands that were oriented in one direction as shown in Figure 5.3b. The associated

matrix had a compressive strength of 42.1 MPa (standard deviation of 4.3 MPa) and a flexural

73

strength of 3.2 MPa (standard deviation of 0.3 MPa). As noted in Table 5.3, the area per unit width

of the carbon fabric was a little more than three times that of the PBO fabric.

Figure 5.3: FRCM systems: a) PBO-FRCM (Unbalanced PBO fabric) and b) C-FRCM

(Unidirectional carbon fabric)

In a previous study conducted by Ebead et al. [97], direct tensile tests were performed on PBO-

and C-FRCM coupons having one layer of embedded fabric in order to characterize their

mechanical properties. The tests were conducted according to AC-434 provisions [65]. Test results

indicated that the stress-strain relationship of FRCM consisted of three distinct stages as shown in

Figure 5.4. In the uncracked stage, the PBO-FRCM coupons showed higher initial stiffness than

their C-FRCM counterparts. After cracking (second stage), the fabric transferred the load back to

the matrix until the tensile strength of the fabric was reached at ultimate (third stage). According

to ACI-549.4R-13 provisions [54], the tensile behavior of each FRCM system could be

characterized by its cracked tensile modulus of elasticity, Ef, its tensile strength, ffu, and its ultimate

strain, εfu. Table 5.4 lists the results obtained from the direct tensile tests for both FRCM systems

that were used in this study [97].

74

Table 5.3: Fabric properties as given in the manufacturers’ data sheets

Fabric Area per unit

width

(𝐴𝑓) (mm2/m)

Tensile

strength (GPa)

Elastic

modulus

(GPa)

Ultimate

strain

(%)

PBO (primary direction) 50 5.8 270 2.15

PBO (secondary direction) 15 5.8 270 2.15

Carbon 157 4.3 240 1.75

Figure 5.4: Stress-strain relationships for FRCM-tensile coupons [97]

Table 5.4 : Mechanical properties of FRCM systems [97]



Ultimate tensile

strength, ffu (GPa)

Ultimate strain,

εfu (%)

PBO-FRCM 121 1.55 1.4

C-FRCM 75 0.97 1.25

75

5.3.4 Strengthening Schemes

The two strengthening schemes used in this study are illustrated in the schematics shown in

Figure 5.5a and Figure 5.5b. Scheme I consisted of one or more FRCM flexure plies having 150

mm width (equal to the beam width) and applied to the soffit of the beam over the middle 2.4 m.

The fabric was oriented so that its primary direction was parallel to the longitudinal axis of the

beam. The flexural plies were anchored at each end using one U-shaped transverse strip of 300

mm width as shown in Figure 5.5a. Scheme II consisted of one or more flexural plies similar to

those used in Scheme I but enclosed with one U-shaped continuous ply running along the clear

span of the beam (Figure 5.5b). For the PBO-FRCM composite used in Scheme II, the primary

direction of the U-wrapped ply was oriented parallel to the longitudinal axis of the beams and was

counted as an additional flexural ply. However, while the bottom flexural plies of the C-FRCM in

Scheme II were oriented parallel to the longitudinal axis of the beams, the U-shaped continuous

ply was oriented in the transverse direction and therefore did not contribute to the flexural

resistance (Figure 5.5b).

For comparison purposes, the equivalent axial stiffness coefficient, 𝐾𝑓, for each FRCM system

was determined based on their cracked modulus and the cross-sectional area of the fabric

embedded within the FRCM composite as follows:

𝐾𝑓= 𝜌𝑓𝐸𝑓 Eq. (5.2)

𝜌𝑓 = 𝑁 𝐴𝑓

𝑑𝑓 Eq. (5.3)

where N is number of the fabric layers in the matrix (shown in each beam’s label in Table 5.1),

b is beam width (150 mm), Af is the fabric area per unit width (Table 5.3), df is the effective depth

of the fabric, Ef is the cracked modulus of the FRCM composite in N/mm2 (Table 5.4), and 𝜌𝑓 is

the fiber reinforcement ratio. Values of 𝜌𝑓 and 𝐾𝑓 are listed in Table 5.1 for all the tested beams.

It is important to note that for beams strengthened with the continuously wrapped PBO-FRCM

layer (Scheme II), the fibers located on the lateral sides of the beams were neglected in estimating

𝜌𝑓 (and consequently 𝐾𝑓). However, their contribution to the flexural strengths of the beams was

considered in the analysis as will be detailed later.

76

The ratio of contribution of FRCM composites to that of steel reinforcement was expressed by

the stiffness factor, 𝛽𝑓, as given in Equation (5.4).

𝛽𝑓 =𝜌𝑓𝐸𝑓

𝜌𝑠𝐸𝑠=

𝐾𝑓

𝐾𝑠 Eq. (5.3)

Where, 𝜌𝑠 is the tensile steel reinforcement ratio and Es is the steel elastic modulus. The steel

reinforcement ratio, 𝜌𝑠, of the corrosion-damaged beams was determined after considering the

actual average steel mass loss due to corrosion. The computed values of 𝜌𝑠 and 𝛽𝑓 are listed in

Table 5.1 for each of the tested beams.

Figure 5.5: Strengthening schemes: a) Scheme I and b) Scheme II

77

5.3.5 FRCM Installation

Prior to FRCM application, all layers of unsound deteriorated concrete were removed as shown

in Figure 5.6a. The corroded steel bars were then cleaned, and a commercial repair mortar was

used to repair the damaged zone. After curing, the beam surface was sandblasted as shown in

Figure 5.6b. The hand lay-up method recommended by ACI-549.4R-13 provisions [54] and the

FRCM manufacturers was adopted during installation. The beam substrate was first soaked in

water for 2 hours before applying the first layer of the cementitious matrix with a thickness of 3 to

5 mm. The fabric was then installed in place and was gently impregnated into the cementitious

matrix. A second layer of matrix having the same thickness was immediately applied as shown in

Figure 5.6c. The procedure was then repeated until the specified number of layers was achieved.

All strengthened beams were left for curing for 28 days in laboratory conditions before being

tested.

Figure 5.6: FRCM installation procedure: a) removing the deteriorated concrete, b) patch

repairing and sandblasting, and c) installation of PBO-FRCM composite

5.3.6 Instrumentation and Test Setup

All beams were instrumented at mid-span with 60 mm long strain gauges on the top concrete

surface and with 5 mm strain gauges bonded to the tensile steel bars. The FRCM-strengthened

78

beams were instrumented with 5 mm strain gauges installed directly on the outer fabric layer of

the FRCM composite at mid-span and at the loading points.

The beams were tested to failure in a four-point bending configuration as shown in Figure 5.1.

All tests were conducted under displacement control at a rate of 2 mm/minute using a MTS

actuator. Deflections at midspan and under the loading points were monitored using three linear

variable differential transducers (LVDTs). All gauges and LVDTs were connected to a 20-channel

data acquisition system that captured the readings at a rate of 5 readings/sec.

5.4 Test Results

The following sections report on the corrosion observations and the test results of both

unstrengthened and strengthened beams in terms of their modes of failure, strength response,

strains in FRCM, and ductility. This is followed by a discussion on the effect of corrosion damage

on the flexural performance of the tested beams.

5.4.1 Corrosion Observations

At the end of the corrosion process, rust stains and longitudinal cracks running parallel to the

corroded steel bars were observed. Corrosion-induced cracks were also observed at the bottom of

the beams and/or on their sides at the level of the tensile steel bars. Table 5.1 lists the average

corrosion crack width, wavg, for each beam. The maximum crack widths were 1.5, 2.8, and 3.5 mm

for groups A, B, and C, respectively. Wider cracks indicated higher steel mass loss due to

corrosion. All the corrosion-damaged beams did not meet the service requirements of ACI 318-14

[98] that limits the maximum crack width to 0.40 mm.

After testing the beams, the corroded steel bars were carefully removed and cut into coupons of

200 mm length each. Visual inspection of the corroded bars revealed the existence of several

corrosion pits that were randomly dispersed on the surface of the bars within the corrosion zone

(Figure 5.7). The mass loss of the corroded bars were determined for each beam according to

ASTM G1-03 provisions [26]. Table 5.1 lists the actual steel mass loss for each beam. Average

mass losses of 12.5, 19, and 22% were determined for groups A, B, and C, respectively,

corresponding to theoretical mass losses of 10, 20, and 30% as predicted from Equation (5.1). The

discrepancy between the theoretical and actual mass losses in the steel bars is shown in Figure 5.8.

79

As reported in [101], this discrepancy was attributed to the reduction in the rate of rust production

as corrosion progresses.

Figure 5.7: Profile of steel bars: a) uncorroded bar, b) corroded bar extracted from CSA-4P-I

(12.6% mass loss), c) corroded bar extracted from CSB-3C-II (18.6% mass loss), and d)

corroded bar extracted from CUC (22.5% mass loss)

Figure 5.8: Actual and theoretical mass loss versus the duration of corrosion process

a)

b)

c)

d)

80

5.4.2 Failure Mechanisms

The failure mechanisms of all the tested beams are summarized in Table 5.1. The virgin beams

(UU) and the corroded unstrengthened beams (CUA, CUB, and CUC) showed classical modes of

failure of under-reinforced beams in which large flexural cracks appeared within the constant

moment zone after steel yielding followed by concrete crushing. On the other hand, three distinct

failure mechanisms were observed in the strengthened beams as shown in Figure 5.9 and were

described as follows:

1) FRCM delamination (FD): this failure mode was encountered in all beams strengthened with

PBO-FRCM systems in Scheme I. The delamination occurred at the fabric/matrix interface

adjacent to the concrete substrate due to the propagation of flexural cracks. This mode of failure

is shown in Figure 5.9a for beam CSA-4P-I.

2) Fabric slippage with partial fabric debonding within the matrix (FS-PFD): this failure mode

was observed in all beams strengthened with PBO-FRCM systems in Scheme II. Vertical flexural

cracks were first observed in the FRCM matrix in the moment zone while inclined cracks extended

in the shear spans. As the applied load increased, gradual fabric slippage within the matrix was

observed until failure occurred. At failure, partial debonding was noticed in the U-shaped ply at

locations where several cracks formed as shown in Figure 5.9b for beam CSC-4P-II. As will be

detailed later, this mode of failure resulted in a more ductile behavior than that observed when the

first mode of failure occurred. This was attributed to the wrapping effect of the continuous U-

shaped FRCM layer that delayed the delamination of the bottom flexural plies.

3) Matrix cracking with extensive fabric slippage (MC-FS): this mode of failure was observed in

all beams strengthened with C-FRCM in Scheme II. Vertical flexural cracks formed in the

cementitious matrix of the U-shaped FRCM layer after steel yielding. As the load increased,

progressive cracking accompanied with large fabric slippage was observed at the beam’s soffit

(Figure 5.9c for beam CSA-3C-II). It should be noticed that this mode of failure was more brittle

than that reported in the PBO-strengthened beams.

81

Figure 5.9: Modes of failure: a) FRCM delamination, b) fabric slippage with partial fabric

debonding within the matrix, and c) matrix cracking with extensive fabric slippage

5.4.3 Strength Response

The load-deflection relationships for the corroded-unstrengthened (CU) beams are shown in

Figure 5.10. Corrosion of steel bars slightly affected the flexure response of the beams with slightly

notable impact on the beams’ stiffness. Due to corrosion, the yield and ultimate loads of beams of

groups A, B, and C were reduced by 7 and 5%, 14 and 7%, and 15 and 9%, respectively, of the

corresponding values of the virgin beams. The yield and ultimate strengths of the corrosion-

damaged beams decreased at average rates of 0.66 and 0.40%, respectively, per 1% of mass loss.

The insignificant effect of corrosion of the tensile steel bars on the beam’s strength could be

attributed to a transition in the beam behavior from a pure flexural action to a tied-arch action due

to bond deterioration within the flexure (corroded) zone while remaining anchored in the shear

(uncorroded) zones. It might also be attributed to the fact that corroded bars lost their lugs, which

increased the measured mass loss without affecting their effective cross section area as shown in

Figure 5.7.

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Figure 5.10: Load-deflection relationships for corroded-unstrengthened beams

Figure 5.11 shows the load versus the midspan deflection curves for the FRCM-strengthened

beams. All the strengthened beams of groups A, B, and C showed similar flexural response. The

load-deflection curves consisted of three stages with two turning points indicating the concrete

cracking in tension and the yielding of the steel bars. FRCM systems slightly enhanced the stiffness

of the strengthened beams prior to steel yielding, which indicated their marginal contribution in

the beam’s behavior. However, the post-yielding response was significantly dependent upon the

number and the type of the FRCM plies used.

(a)

83

(b)

(c)

Figure 5.11: Load-deflection relationships for corrosion-damaged FRCM-strengthened beams: a)

beams of Group A, b) beams of Group B, and c) beams of Group C

The yield load, 𝑃𝑦, ultimate load, 𝑃𝑢, and normalized yield and ultimate loads to those of the

control specimen, 𝑃𝑦 𝑁𝑜𝑟 and 𝑃𝑢 𝑁𝑜𝑟, respectively, are presented in Table 5.1. It can be noticed that

all the FRCM-strengthened beams fully restored the yield and ultimate strengths of the virgin beam

except beam CSB-2P-I that restored only 96% of the yield capacity. The gain in strengths was

highly dependent upon the amount of fabric, the FRCM strengthening scheme, and the type of the

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FRCM used. The use of two layers of PBO-FRCM in beams CSA-2P-I and CSB-2P-I increased

their ultimate strengths by 8 and 7%, respectively, of that of the virgin beam while using four PBO-

FRCM plies in beams CSA-4P-I, CSB-4P-I, and CSC-4P-I increased their ultimate strengths by

25, 29, and 29%, respectively.

The use of a U-shaped fabric layer in Scheme II was more efficient in increasing the load-carrying

capacities of the strengthened beams than the use of end anchors in Scheme I. For instance, the

use of four plies of PBO-FRCM with Scheme II in beams CSA-4P-II, CSB-4P-II, and CSC-4P-II

increased their ultimate strengths by 44, 29, and 39% compared to 25, 29, and 29% increase in

beams CSA-4P-I, CSB-4P-I, and CSC-4P-I, respectively. This increase in strength was associated

with the contribution of the longitudinal PBO fibers attached to the lateral surfaces of the beams.

Moreover, wrapping the beam with a continuous U-shaped PBO layer delayed the premature

delamination of the fabric and consequently increased the strengthening effectiveness of the

FRCM system.

The use of three layers of C-FRCM in Scheme II increased the load-carrying capacities of beams

CSA-3C-II, CSB-3C-II, and CSC-3C-II by 51, 55, and 37%, respectively. It was obvious from the

test results that, despite the various levels of corrosion, the difference in gain in strengths observed

in most of the tested beams strengthened with similar FRCM systems was marginal. This finding

will be discussed in the following sections.

The equivalent axial stiffness of the FRCM composites, Kf, was utilized to compare between the

load-carrying capacities of the strengthened beams (normalized with respect to that of the virgin

beam UU). These results are shown in Figure 5.12. It can be noticed that almost doubling the

equivalent axial stiffness of PBO-FRCM systems from 48 MPa (2 plies in beam CSA-2P-I) to 95.3

MPa (4 plies in beam CSA-4P-I) increased the gain in the load-carrying capacities from 10.3 kN

(for beam CSA-2P-I) to 23.5 kN (for beam CSA-4P-I). A similar trend was observed for specimens

CSB-2P-I and CSB-4P-I. This finding revealed that the gain in ultimate strengths was

approximately proportional to the axial stiffness of the PBO-FRCM systems represented by the

coefficient Kf. This finding was also confirmed by the similar ultimate fiber strains recorded for

specimens CSA-2P-I, CSA-4P-I, CSB-2P-I, and CSB-4P-I (Figure 5.13a).

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Figure 5.12: Normalized strength versus the FRCM equivalent axial stiffness, Kf

On the other hand, the beams strengthened with three layers of C-FRCM in Scheme II (beams

CSA-3C-II, CSB-3C-II, and CSC-3C-II) showed average load-carrying capacities only 7.5%

higher than those of their counterparts strengthened with four layers of PBO-FRCM in the same

scheme whereas the equivalent axial stiffness of three layers of C-FRCM was 146% higher than

that of four layers of PBO-FRCM. This finding was attributed to the premature cracking of the

matrix associated with the C-FRCM system, which limited the strengthening effectiveness of the

system. This was also confirmed by the strain results shown in Figure 5.13b. The ultimate strains

measured in the fibers of beams strengthened with C-FRCM were obviously lower than those

measured in the fibers of beams strengthened with PBO-FRCM.

FRCM configuration had a notable effect on the flexural response of the beams. Beams

strengthened with four plies of PBO-FRCM systems using Scheme II (beams CSA-4P-II, CSB-

4P-II, and CSC-4P-II) showed an increase in their ultimate strengths that ranged between 29 and

44% with an average gain of about 37%. Their counterpart beams CSA-4P-I, CSB-4P-I, and CSC-

4P-I strengthened with the same amount of FRCM layers (similar FRCM equivalent stiffness) but

in Scheme I showed an average increase in their strengths of about 28%. As previously noted, this

increase in strength was associated with the contribution of the longitudinal fibers attached to the

lateral surfaces of the beams and with the wrapping effect of the U-shaped fabric on delaying the

delamination of the PBO-FRCM fabrics.

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5.4.4 Fabric Strains

Table 5.1 lists the strains recorded in the outer fabric layer at ultimate load, εfu, for all the

strengthened beams. Figure 5.13a and 5.13b shows the fabric strains in the beams strengthened in

Scheme I and II, respectively.

(a)

(b)

Figure 5.13: Load versus outer fabric strain for beams strengthened in a) Scheme I and b)

Scheme II

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It is important to note that the ultimate fiber strains measured during the tests were lower than

those obtained from the direct tensile tests conducted on FRCM coupons listed in Table 5.4.

Measured strains ranged between 4885 μɛ for beam CSC-3C-II and 10659 μɛ for beam CSB-4P-I

while the tested coupons showed strains of 12500 μɛ for C-FRCM and 14000 μɛ for PBO-FRCM

[97]. This discrepancy in the strain results between the flexural and tensile specimens was

attributed to the different failure mechanisms observed in both the strengthened beams and the

coupons. While three distinct modes of failure were encountered in the strengthened beams, all the

tested FRCM coupons failed due to fabric slippage in tension.

5.4.5 Ductility and Energy Absorption

The ductility index, ΔI, was determined for all beams as the ratio of deflections at ultimate, 𝛿𝑢,

to the deflection at yielding, 𝛿𝑦. The energy absorption index, ψ, represented by the area under the

load-deflection curve up to the ultimate point was also determined. Both indices were utilized to

evaluate the ductility of the tested beams. Table 5.5 lists the deflections, the ductility indices, and

the energy absorption indices for each beam. It was noticed that the corrosion-damaged beams

showed higher ductility indices than that of the virgin beams. The ductility index increased with

the increase of the corrosion level. The unstrengthened beams of groups A, B, and C (average mass

loss of 12.9, 18, and 22.5%, respectively) showed ductility indices of 3.2, 3.3, and 3.4 representing

114, 118, and 121%, respectively, of those of the virgin beams. The same beams showed energy

absorption indices of 118, 104, and 88% of that of the virgin beams, respectively. On the other

hand, beams strengthened with FRCM composites showed ductility indices that ranged between

86 to 118% of that of the virgin beams whereas their energy absorption indices improved by 11 to

55%.

Figure 5.14 and Figure 5.15 show the change in the normalized indices with the stiffness factor

𝛽𝑓. Recall that 𝛽𝑓 represents the ratio of the axial stiffness of the FRCM system to that of the steel

reinforcement and therefore represents the contribution of the FRCM composites to the beam’s

performance. Beams strengthened with two layers of PBO-FRCM in Scheme I showed average

ductility and energy absorption indices of 103.5 and 116.5 % of those of the virgin beam,

respectively, whereas beams strengthened with four layers of PBO-FRCM showed average indices

of 100 and 134.3%, respectively. For beams strengthened with four layers of PBO-FRCM in

Scheme II, the average indices were 93 and 130% of those of the virgin beam, respectively. On

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the other hand, beams strengthened with three layers of C-FRCM had similar average indices to

those of their counterparts strengthened with four layers of PBO-FRCM and Scheme II.

Table 5.5: Ductility and energy absorption of the tested beams

Specimen

Midspan

deflection, mm Ductility index Energy absorption index

δy δu ΔI ΔINor.** ψ ψ Nor.

**

UUa, UUb* 11.7 32.9 2.81 1 1972 1

CUA 10.9 35.4 3.2 1.14 2326 1.18

CSA-2P-I 11.83 33.02 2.8 1 2314 1.17

CSA-4P-I 12.86 31.48 2.4 0.86 2346 1.19

CSA-4P-II 12.64 35.17 2.8 1 3059 1.55

CSA-3C-II 12.55 30.1 2.4 0.86 2491 1.26

CUB 8.6 28.26 3.3 1.18 2059 1.04

CSB-2P-I 11.28 33.72 3 1.07 2292 1.16

CSB-4P-I 12.54 37.41 3 1.07 2933 1.49

CSB-4P-II 11.83 29.48 2.5 0.89 2184 1.11

CSB-3C-II 10.57 34.75 3.3 1.18 3012 1.53

CUC 9.8 32.85 3.4 1.21 1735 0.88

CSC-4P-I 12.04 35.53 3 1.07 2657 1.35

CSC-4P-II 12.69 31.32 2.5 0.89 2426 1.23

CSC-3C-II 12.32 30.16 2.4 0.86 2212 1.12

* Average values reported **Normalized with respect to the virgin beam

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Figure 5.14: Normalized ductility index versus stiffness factor 𝛽𝑓 %

Figure 5.15: Normalized energy absorption index versus stiffness factor 𝛽𝑓 %

5.5 Discussion

The effect of corrosion damage on the flexural response of the strengthened beams is discussed

in the light of the presented test results. As previously mentioned, almost all the strengthened

beams restored the yield and ultimate capacities of the virgin beams. Figure 5.16 shows the effect

of corrosion level on the load-carrying capacities of the strengthened beams.

Figure 5.16 reveals that beams strengthened with four PBO-FRCM layers in Scheme I (beams

CSA-4P-I, CSB-4P-I, and CSC-4P-I) showed almost similar capacities despite the different

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degrees of corrosion in each group of beams. The comparison between the ultimate capacity of

beam CSA-4P-I that suffered 12.6% mass loss due to corrosion and that of beam CSC-4P-I that

suffered almost double of that mass loss (21.1%) revealed the negligible effect of corrosion on the

load-carrying capacities of the strengthened beams. The former beam (beam CSA-4P-I) showed a

load-carrying capacity of 99.6 kN whereas the latter one (beam CSC-4P-I) showed a load-carrying

capacity of 102.8 KN, with a variation in capacity gain not exceeding 3%. This finding was also

confirmed from the strength results of the tested beams having different degrees of corrosion

damage and strengthened with four plies of PBO-FRCM and with three plies of C-FRCM in

Scheme II. The variation in gain in capacities did not exceed 7 and 9% for the former and the latter

beams, respectively.

Figure 5.16: Effect of corrosion damage on the ultimate strength of strengthened beams

The insignificant impact of corrosion levels on the load-carrying capacities of the strengthened

beams was previously confirmed with the capacities obtained for the unstrengthened beams after

corrosion. Beams CUA, CUB, and CUC showed almost similar capacities ranging between 91 and

95% of the capacity of the virgin beams. It was also attributed to the same mode of failure

encountered in all beams that were strengthened with the same amount and same configuration of

FRCM composites. Based on test observations, the three modes of failure reported for the

strengthened beams showed that failure was primarily governed by the FRCM scheme and the

type of the FRCM composite rather than the damage level to which the beam was subjected due

to corrosion. Knowing that the maximum mass loss in the steel reinforcing bars was 22.7%, which

is quite significant, the reported load-carrying capacities of the corrosion-damaged beams

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confirmed that strengthening with FRCM composites had counterbalanced the impact of the

corrosion damage.

Figure 5.17 illustrates the effect of corrosion damage on the load-carrying capacities of the

strengthened beams normalized with respect to those of the virgin beams. Recall that the degree

of corrosion damage in the steel reinforcing bars was represented by a reduction in the axial

stiffness 𝐾s = 𝜌𝑠𝐸𝑠. The contribution of FRCM composites to the beam strength was compared to

that of steel reinforcement using the stiffness factor, 𝛽𝑓 = 𝐾𝑓

𝐾𝑠.

Figure 5.17: Normalized strength versus stiffness factor 𝛽𝑓 %

In Figure 5.17, a best fit curve was plotted to address the variation in the ultimate capacities with

the factor 𝛽𝑓. A linear trend was obtained with a least-square coefficient 𝑅2 = 0.75. As can be seen

in Figure 5.17, a tendency of increase in the ultimate capacities is observed with the increase in

the stiffness factor, 𝛽𝑓, from 2.58% (for beam CSA-2P-I) to 8.31% (for beam CSC-3C-II). Based

on the test results, it can be concluded that the increase in 𝛽𝑓 values was largely affected by the

increase in the equivalent stiffness 𝐾𝑓 rather than the decrease in the equivalent stiffness 𝐾𝑠 that

resulted from the mass loss in the steel bars, which can explain the ascending trend in Figure 5.17.

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5.6 Predicted Strength Results

Analytical calculations were conducted according to ACI 318-14 [98] and ACI 549.4R-13 [54]

provisions to predict the flexural response of the tested beams. The FRCM composites were

assumed to behave linearly up to failure while the steel bars were assumed to behave elastic-

perfectly plastic. The yield strengths of the longitudinal reinforcing steel bars of diameter 15 and

8 mm were taken equal to 466 and 573 MPa, respectively, with elastic modulus 𝐸𝑠 = 200 GPa. The

maximum compression strain in concrete was taken equal to 0.003 mm/mm. Prefect bond between

the concrete substrate and the FRCM system and between the FRCM fabric and the matrix was

assumed during the analysis. However, this assumption was governed by the strain limit of 0.012

mm/mm imposed by ACI 549.4R-13 provisions [54]. Therefore, the design effective tensile strain,

εfe, in FRCM was assumed equal to the experimental ultimate strain, εfu, as obtained from the

coupon test results minus one standard deviation or 0.012 mm/mm, whichever was lower. The

design effective tensile strength, ffe was taken equal to 𝐸𝑓휀𝑓𝑒, where Ef is the cracked tensile

modulus of elasticity of FRCM shown in Table 5.4. The reduction in the cross section of the steel

bars due to corrosion was considered according to the actual average mass loss encountered in

each beam as reported in Table 5.1. The flexural contribution of the longitudinal fibers bonded to

the sides of the beams strengthened in Scheme II was considered in strength calculations.

Following these assumptions, the cracking, yielding, and ultimate load-carrying capacities and the

corresponding deflections were calculated.

The predicted responses were compared to the experimental ones in Figures 5.18a to 5.18c. It

can be depicted that, prior to yielding of steel bars, the predicted and experimental responses of all

beams followed almost a similar trend. However, the predicted responses were lower than the

experimental ones after yielding. The theoretical ultimate capacities, 𝑃𝑢𝑡ℎ, and the ratios of the

experimental to the theoretical values, 𝑃𝑢

𝑃𝑢𝑡ℎ , for all the tested beams are shown in Table 5.1. The

ratios 𝑃𝑢

𝑃𝑢𝑡ℎ ranged between 1.03 for beam CSA-2P-I and 1.23 for beam CSB-3C-II. The ratios

𝑃𝑢

𝑃𝑢𝑡ℎ

indicated that the theoretical formulations of ACI 549.4R-13 [54] reasonably predicted the ultimate

capacities of the corrosion-damaged RC beams strengthened with FRCM composites in Scheme I

but underestimated the capacities of those strengthened in Scheme II. This finding was attributed

to the fact that ACI 549.4R-13 provisions [54] don’t consider the strengthening scheme in their

93

formulations. It was also related to the different modes of failure that occurred in the strengthened

beams. While fabric debonding occurred in beams strengthened with Scheme I and II as previously

explained, the U-shaped FRCM layer in Scheme II played a major role in confining the bottom

fabric layers and in delaying their delamination from the matrix. This explains the enhancement in

the load-carrying capacities of the beams strengthened in Scheme II as compared to their

counterparts strengthened in Scheme I. Since the ACI formulations don’t consider the effect of

scheme in their predictions, they tend to underestimate the capacities of beams strengthened in

Scheme II.

The obtained results suggested increasing the predicted ultimate capacity of the beams

strengthened in Scheme II by 10% to consider the effect of the continuous anchoring. The

comparison between the theoretical and experimental results also indicated that the assumption of

prefect bond suggested by ACI 549.4R-13 [54] while limiting the design effective strain, εfe, in

the FRCM systems is justified.

(a)

94

(b)

(c)

Figure 5.18: Predicted versus experimental flexural response for beams strengthened with a) two

layers of PBO-FRCM in Scheme I, b) four layers of PBO-FRCM in Scheme II, and c) three

layers of C-FRCM in Scheme II

95

5.7 Conclusions

The flexural behavior of RC beams strengthened with PBO- and C-FRCM composites after being

exposed to three different degrees of corrosion damage was presented. The following conclusions

can be drawn from the test results:

• Corrosion of steel reinforcing bars in the moment zone with an average mass loss up to

22.7% had a marginal impact on the flexural response of the tested beams. The maximum

decrease in the yield and ultimate strengths of the corrosion-damaged beams were 15 and

9%, respectively, of those of the virgin beams.

• The use of PBO- and C-FRCM systems enhanced the flexural response of the corrosion-

damaged beams. The type, amount, and anchoring scheme of the applied FRCM composite

governed the failure mechanism and the load-carrying capacities of the strengthened beams.

• Beams strengthened with PBO-FRCM showed strength gain that ranged between 7 and 44%

of that of the virgin beam. Beams strengthened with PBO-FRCM in Scheme I failed by

FRCM delamination while those strengthened in Scheme II failed due to fabric slippage

within the matrix.

• The strength gain in beams strengthened with C-FRCM in scheme II ranged between 39 to

55 % of that of the virgin beams and failed due to excessive premature matrix cracking.

• The mechanical properties of the cementitious matrix and their bond to the fabric layers

played a dominant role in defining the failure mechanism of the PBO- and C-FRCM

strengthened beams and consequently the strengthening efficiency of the FRCM system.

• Although the equivalent stiffness of the FRCM composites, Kf, and the stiffness factor, 𝛽𝑓,

are believed to be good indicators of the strengthening effectiveness of the FRCM systems,

they should not be solely used to compare the strength gain in beams without considering

the mechanical characteristics of the matrix, the number of fabric plies, and the anchoring

scheme used.

96

• Beams strengthened with FRCM composites showed ductility indices and energy absorption

indices that ranged between 86 and 118% and between 111 and 153%, respectively, of those

of the virgin beams.

• The theoretical formulations of ACI-549.4R-13 reasonably predicted the capacities of the

beams strengthened in Scheme I (with end anchors) but underestimated the capacities of

those strengthened in Scheme II (continuous U-shaped). This was attributed to the fact that

ACI-549.4R-13 formulations don’t consider the effect of the continuous anchorage on

delaying the delamination of the FRCM system.

• The theoretical capacities of beams strengthened in scheme II should be increased by 10%

to consider the effect of the continuous anchoring.

97

6. Chapter 6

Post-repair Flexural Performance of Corrosion-Damaged

Beams Rehabilitated with Fabric-Reinforced Cementitious

Matrix (FRCM)


Journal of Construction and Building Materials. Date of acceptance: January 16,2018

https://doi.org/10.1016/j.conbuildmat.2018.01.128

Résumé

Cet article présente les résultats d'un programme de recherche examinant la performance post-

réparation en flexion des poutres en béton armé endommagées par la corrosion et réparées avec

différents systèmes MCRF. Neuf poutres ont été testées, y compris deux poutres qui n'ont été ni

corrodées ni réparées, une poutre corrodée et non réparée et six poutres corrodées et réparées en

deux phases. Les poutres de la phase I ont été soumises à un processus de corrosion accéléré

pendant 210 jours avant d'être réparées alors que les poutres de la phase II ont été soumises à une

corrosion accélérée pendant 70 jours, puis réparées et exposées une autre fois à la corrosion

pendant 140 jours. Les résultats des essais de flexion ont montré que l'exposition des poutres

réparées par MCRF à la corrosion après réparation a entraîné une réduction de 23% de la perte de

masse de l'acier. Les couches MCRF en forme de U étaient plus efficaces à réduire le taux de

corrosion et à augmenter la résistance ultime des poutres réparées que les couches MCRF ancrées

à l'extrémité. Les poutres réparées par la MCRF au PBO ont montré une plus faible rigidité post-

plastification et une plus grande ductilité à l’ultime que celles de leurs contreparties réparées par

MCRF au carbone. Les poutres soumises à un environnement corrosif après réparation présentaient

des capacités de charge qui se situaient entre 14 et 65% au-dessus de celles des poutres vierges.

Les dispositions de l'ACI 549.4R-13 prévoient de façon conservatrice les capacités ultimes des

poutres réparées par MCRF et exposées à un environnement corrosif après réparation.


Réparation; Renforcement; Réhabilitation; Durabilité; Performance à long terme.

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6.1 Abstract

This paper presents the results of a research program examining the post-repair flexural response

of corrosion-damaged reinforced concrete (RC) beams repaired with different FRCM systems. A

total of nine RC beams were tested, including two beams that were neither corroded nor repaired,

one beam that was corroded and not repaired, and six corroded-repaired beams that were prepared

in two phases. Beams of phase I were subjected to an accelerated corrosion process for 210 days

before being repaired whereas beams of phase II were initially subjected to accelerated corrosion

for 70 days, then repaired and exposed to further corrosion for 140 days. Flexural test results

showed that exposing the FRCM-repaired beams to corrosion after repair resulted in 23%

reduction in steel mass loss. The use of U-shaped FRCM layers was more efficient in reducing the

corrosion rate and increasing the ultimate strength of the repaired beams than the end-anchored

FRCM layers. The PBO FRCM-repaired beams showed lower post-yielding stiffness and more

ductility at failure than those of their carbon FRCM-repaired counterparts. Beams that experienced

post-repair corrosive environment showed load-carrying capacities that ranged between 14 and

65% above those of the virgin beam. ACI 549.4R-13 provisions conservatively predict the ultimate

capacities of the FRCM-repaired beams exposed to post-repair corrosive environment.

Authors’ keywords: Corrosion; Fabric-reinforced cementitious matrix; Flexure; Reinforced

concrete; Repair; Strengthening; Rehabilitation; Durability; Long-term performance.


Repair/strengthening of reinforced concrete (RC) structures is motivated by several factors

including aging, change in use, increased loads, code compliance, and environmental damage (e.g.

corrosion). Corrosion of steel reinforcement is one of the major durability concerns for concrete

structures especially in coastal areas and cold regions where de-icing salts are heavily used. Pitting

corrosion reduces the cross-sectional area of the steel bars and may lead to decrease in ductility

of the steel bars [1,34]. The expansion of corrosion products causes concrete cracking and impair

the composite action of steel and concrete. As a result, the load-bearing capacity and the service

life of the corroded member are considerably jeopardized [4,23,102].

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Existing externally-bonded repair technologies based on organic matrices referred to as fiber-

reinforced polymer (FRP) have proven success in restoring the serviceability and strength of RC

structures [10–12]. Using FRP products in repair applications was driven by their non-corrosive

properties, lightweight, and high tensile strength. However, FRP matrices are flammable, prone to

deterioration at high temperatures, and have poor thermal compatibility to the concrete substrate

[57,59,92,93]. In order to overcome these drawbacks, fabric-reinforced cementitious matrix

(FRCM) systems have been introduced as promising alternatives.

FRCM systems consist of one or more layers of textiles made of carbon, glass, or

Polyparaphenylene benzobisoxazole (PBO) grids that are sandwiched between layers of

cementitious mortars. Its lightweight, high tensile strength, and ease of application makes the

system appealing. The technique also surmounts the epoxy-bonded FRP systems that lack fire

resistance as the embedded grid is shielded between the mortar layers thus minimizing its

vulnerability hazard as the organic matrix is no longer present. In addition, the compatibility

between the mortar used and the concrete substrate is inherited since both materials have the

cement as a common “base”. FRCM systems, with their innovative features, ensure the endurance

of the rehabilitation process and consequently the sustainability of the strengthened structure.

Much research has been reported on reinforced and prestressed concrete structures strengthened

with FRCM subjected to monotonic loading [58,84]. Previous research studies have proven the

success of FRCM in enhancing the performance of RC structures [103,104] and masonry structures

[105]. D’Ambrisi and Focacci [81] demonstrated that the performance of FRCM materials was

strongly dependent on the fabric type and the matrix constituents. Schladitz et al. [55] reported an

increase in the load-carrying capacity of RC slabs strengthened with one and four layers of C-

FRCM by 67% and 245%, respectively. Fabric rupture was observed at failure in all of the

strengthened slabs even when high amount of FRCM layers was used. Significant decrease in

deflection was reported when four layers of FRCM were used [55]. Elsanadedy et al. [106]

reported that providing U-anchorage at the FRCM ends was efficient in alleviating the FRCM end

debonding in flexure tests. More recently, Ebead et al. reported that the failure mechanism of

carbon-and PBO-FRCM-strengthened beams was governed by inter-laminar shear [97].

Only one study [86] has documented the effectiveness of using basalt and carbon FRCM systems

to restore the ultimate capacity and serviceability of T-beams after a mass loss of 22% in their

100

steel reinforcement due to corrosion. It was reported that the basalt-FRCM system could not restore

the original flexural capacity of the beam whereas the C-FRCM system restored 109% of the

capacity. The authors reported that the use of a combination of internally-embedded and

externally-bonded C-FRCM layers was more effective in improving the strength and ductility of

the beams than the use of the same amount of FRCM layers internally embedded within the

corroded-repaired region.

Corrosion-damaged structures are often vulnerable to the same deterioration mechanism after

repair, which may require further repair during their service life. To date, no data is available in

the literature concerning the post-repair performance of FRCM-repaired beams that might be

vulnerable to further corrosive environments while in service. This paper aims at filling this gap

by reporting on the flexural performance of post-repair RC beams. The test program investigated

the type of the FRCM used (carbon and PBO) and the FRCM repair scheme (end-anchored and U-

wrapped). The paper also reports on the failure modes, the load-carrying capacities, the ductility,

and the straining actions as observed and quantified during the tests. The ultimate load-carrying

capacities of the tested beams were calculated in accordance with the provisions of the ACI

549.4R-13 [54] and compared to the experimental results.


The test matrix of the experimental program is given in Table 6.1. Nine large-scale RC beams,

including two virgin uncorroded beams (UU), one corroded unrepaired specimen (CU), and six

corroded and repaired beams were prepared in two phases. In phase I, three beams (referred to as

the short-term beams) were subjected to an accelerated corrosion process for 210 days. The beams

were then tested immediately after being repaired with different FRCM systems and

configurations. In phase II, three beams (referred to as the long-term beams) were initially

subjected to an accelerated corrosion process for 70 days. The beams were then repaired with

different FRCM systems. After curing, the beams were exposed to further corrosive environment

for 140 days prior to testing. The schematic shown in Figure 6.1 describes the two phases of the

experimental program.

The beams were labeled following the A-B-C format. ‘A’ represents the beam condition (UU,

CU, and CR referring to Uncorroded-Unrepaired, Corroded-Unrepaired, and Corroded-Repaired,

101

respectively), while ‘S’ and ‘L’ refer to the short-term and long-term testing procedure,

respectively. ‘B’ denotes the number and type of the FRCM layers applied (4P, and 3C referring

to four layers of PBO-FRCM and three layers of C-FRCM, respectively). Finally, ‘C’ describes

the FRCM strengthening schemes (I and II) as will be detailed in the following sections.


* 𝐴𝑠 was determined based on the measured average tensile steel mass loss due to corrosion

Figure 6.1: Schematic of the testing procedure of the short- and long-term beams

6.3.1 Test Specimen and Materials

Specimen Ave. Mass

loss (%)

Max. Mass

loss (%) 𝐴𝑠

*

(mm2)

𝜌𝑓

(%)

𝐾𝑓 = 𝜌𝑓𝐸𝑓

(MPa)

Control specimens

UUa, UUb - - 400 - -

CU 22.5 25.9 310 - -

Phase I: Short-term specimens

CRS-4P-I 22.7 25.6 309.2 0.08 95.3

CRS-4P-II 21.1 23.9 315.6 0.08 95.3

CRS-3C-II 21.5 23.2 314 0.185 139.1

Phase II: Long-term specimens

CRL-4P-I 18.7 20.3 325.2 0.08 95.3

CRL-4P-II 18.1 19.2 327.6 0.08 95.3

CRL-3C-II 17.5 18.7 330 0.185 139.1

102

All of the beams were 2.8 m long with similar cross section of width, b = 150 mm and height, h

= 250 mm. The beams were reinforced with two 8M (diameter 8 mm) and 15M (diameter 15 mm)

deformed rebars at top and bottom, respectively. The test specimen was designed to fail in flexure

under four-point load configuration. Thus, the shear spans were reinforced with 10M (diameter 10

mm) deformed stirrups spaced at 100 mm. A hollow stainless-steel tube with external and internal

diameters of 9.5 mm and 7 mm, respectively, was placed at 80 mm from the specimen tension face

to act as a cathode during the accelerated corrosion process. Typical dimensions and reinforcement

details of the test specimen are shown in Figure 6.2.

Normal and salted ready-mixed concrete batches having similar water/cement ratio were used to

cast the beams. Six standard concrete cylinders (150×300 mm) were prepared from each batch and

were tested in compression after 28 days and on the day of testing. Table 6.2 lists the compressive

strengths of both mixes. Prior to FRCM application, the corroded beams were repaired using local

commercial cementitious repair mortar (Sikacrete-08SCC) having a compressive strength of 55.4

MPa (standard deviation of 5 MPa) and flexural strength of 3.4 MPa (standard deviation of 0.3

MPa) as tested by the authors. The yield strengths of the longitudinal reinforcing steel bars of

diameter 15 and 8 mm were 466 MPa (standard deviation of 4.2 MPa) and 573 MPa (standard

deviation of 17.7 MPa), respectively, as tested by the authors.

Figure 6.2: Typical dimensions and reinforcement details of the test beam (all dimensions in

mm)

103

Table 6.2 : Concrete compressive strengths

Compressive strength,

MPa

Standard deviation,

MPa


%






A potentiostatic technique was used to accelerate the corrosion of the tensile steel reinforcement.

The steel bars were connected to the positive terminal of a DC power supply to work as anode

whereas the stainless-steel tube was connected to the negative terminal of the power supply to act

as cathode. The specimens were electrically connected in series as shown in Figure 6.3 to obtain a

current density of 180 µA/cm2 that was impressed in the reinforcing bars. Corrosion of the steel

bars was localized in the middle 1200 mm of the beam’s span. Salt (NaCl) measured as 5% of the

cement weight was added to the concrete mix used to cast the middle-bottom of the corroded

specimens with a height of 100 mm (Figure 6.2). Similar amount of salt was added to the repair

mortar used to repair the long-term specimens. During the accelerated corrosion process, the

specimens were subjected to wet-dry cycles that consisted of 3 days wet followed by 3 days dry in

a large environmental chamber. The wet-dry cycles provided water and oxygen necessary for the

corrosion process and simulated the environmental conditions of a beam in-service. According to

Faraday’s law, the mass loss in the reinforcing bars was estimated as 30% based on the intensity

of the applied electrical current and the duration of the corrosion process.

104

Figure 6.3: A schematic of the electrical connection

6.3.3 FRCM Composites

Two commercial FRCM systems (PBO and carbon) were used to strengthen the corroded

specimens (Figure 6.4). The fabric properties in the primary direction as reported in the

manufacturers’ data sheet are shown in Table 6.3. The PBO fabric consists of an unbalanced net

of fiber rovings spaced along two orthogonal directions as shown in Figure 6.4a. The associated

inorganic cementitious matrix had a compressive strength of 43.9 MPa (standard deviation of 0.4

MPa) and a flexural strength of 3 MPa (standard deviation of 0.3 MPa) after 28 days as tested by

the authors. On the other hand, the carbon FRCM composite consists of a unidirectional net made

of carbon-fiber strands (Figure 6.4b) and impregnated in an inorganic cementitious matrix of

compressive strength of 42.1 MPa (standard deviation of 4.3 MPa) and flexural strength of 3.2

MPa (standard deviation of 0.3 MPa) after 28 days as tested by the authors.

105

Figure 6.4:FRCM systems; a) PBO-FRCM (unbalanced PBO fabric) and (b) C-FRCM

(unidirectional carbon fabric) - all dimensions in mm

Table 6.3: Fabric properties in the primary direction as given in the manufactures’ data sheet


𝐴𝑓, mm2/m Tensile strength,

GPa

Elastic modulus,

GPa

Ultimate strain,

%

PBO 50 5.8 270 2.15

Carbon 157 4.3 240 1.75

According to ACI 549.4R-13 [54], the tensile stress-strain behavior of a FRCM composite

coupon can be expressed by the bilinear relationship shown in Figure 6.5. The initial linear portion

corresponds to the uncracked behavior of the FRCM matrix while the second linear portion

represents its cracked behavior up to failure. Table 6.4 lists the properties of the FRCM composite

systems as reported in [97].

106

Figure 6.5: Idealized tensile stress-strain curve of FRCM coupon specimen ACI 549.4R-13 [54]

Table 6.4: Mechanical properties of FRCM coupons as reported in [97]


of elasticity Ef, GPa

Ultimate tensile

strength ffu, GPa

Ultimate strain

εfu, %

PBO-FRCM 121 1.55 1.4

Carbon-FRCM 75 0.97 1.25

An equivalent axial stiffness, Kf, given in Equation (6.1), was utilized to compare between the

two FRCM systems based on their cracked elastic modulus, 𝐸𝑓, and the cross-sectional area of the

fabric. The equivalent axial stiffness of each strengthened specimen is shown in Table 6.1. It is

important to mention that for beams repaired with continuous U-shaped PBO-FRCM layer

(Scheme II), the fiber strands on the lateral sides of beams were neglected in estimating 𝜌𝑓 (and

consequently 𝐾𝑓). Their contribution to the flexural strengths of the beams was considered in the

analysis as will be explained later.

𝐾𝑓 = 𝜌𝑓𝐸𝑓 Eq. (6.1)

Where, 𝜌𝑓 and 𝐸𝑓 are the fiber reinforcement ratio and the cracked elastic modulus of the FRCM

composite, respectively.

6.3.4 FRCM Schemes

Two FRCM strengthening schemes were utilized in this study as shown in Figure 6.6. Scheme I

consisted of four PBO-FRCM flexure plies having 150 mm width (equal to the width of the beam)

107

and applied to the soffit of the beam over a length of 2.4 m. The fabric was oriented so that its

primary direction was parallel to the longitudinal axis of the beam. The flexural plies were

anchored at each end using one U-shaped transverse strip of 300 mm width (Figure 6.6a). Scheme

II consisted of bottom flexural plies as in Scheme I but wrapped with one U-shaped continuous

ply along the beam’s clear span (Figure 6.6b). The primary direction of the U-wrapped PBO ply

in Scheme II was oriented parallel to the longitudinal axis of the beams and was counted as an

additional ply. On the other hand, the carbon fabric is a unidirectional fabric. Therefore, the bottom

flexural plies of C-FRCM composite in Scheme II were oriented parallel to the longitudinal axis

of the beams whereas the U-shaped layer was oriented in the transverse direction and therefore did

not contribute to the flexural resistance of the beam (Figure 6.6b).

Figure 6.6: FRCM repair schemes; a) Scheme I and b) Scheme II

108

6.3.5 Repair Methodology

Following the initial corrosion, the corroded beams were placed upside down with the beam soffit

at the top to facilitate the repair process. The deteriorated concrete was first removed using a

hydraulic hammer. The corroded steel bars were totally exposed and cleaned using a steel brush.

The grooved area was then filled with local cementitious repair mortar (Sikacrete-08SCC). After

the curing of the repair mortar, the beam surface was sandblasted prior to FRCM application. The

hand lay-up method proposed by the ACI 549.4R-13 [54] and the manufacturers was followed

while installing the FRCM systems. The beam’s substrate was first damped in water for 2 hours

before applying the first layer of the cementitious matrix with a thickness of 3 to 4 mm. The fabric

was then installed in place and was gently impregnated into the cementitious matrix and covered

with a second layer of matrix with similar thickness. The procedure was then repeated until the

specified number of layers was achieved. All repaired beams were cured for 28 days in laboratory

conditions before being tested or further exposed to corrosion.


The specimens were tested to failure in a four-point bending configuration as shown in Figure

6.1. All tests were conducted under displacement control at a rate of 2 mm/minute using a MTS

actuator. A spreader steel beam was used to spread the load equally to two loading points spaced

800 mm apart. Beam deflections were measured by means of three linear variable differential

transducers (LVDTs) located at mid-span and at the loading points. All beams were instrumented

at mid-span with a 60 mm long strain gauge bonded to the top surface of concrete and 5 mm strain

gauges bonded to the tensile steel bars. The FRCM-repaired beams were instrumented with 5 mm

strain gauges installed directly on the outer fabric of the FRCM composite at mid-span and at the

loading points. The measuring instruments were connected to a 20-channel data acquisition to

capture the readings at all stages of loading.

6.4 Test Results and Discussion

6.4.1 Corrosion Crack Pattern

Crack patterns that resulted from corrosion were reported for all specimens at the end of the

accelerated corrosion process. For specimens in Phase I (short-term), longitudinal cracks extending

109

parallel to the corroded reinforcement were observed on the lateral surface and the soffit of the

beams. Figure 6.7a shows the cracks pattern due to corrosion as observed for the corroded-

unrepaired beam (CU). Those cracks had a maximum and average widths of 3.5 mm and 1.6 mm,

respectively. All specimens in this phase failed to meet the ACI 318-14 [98] service requirements

that limits the maximum crack width during service to 0.40 mm. For the specimens exposed to the

post-repair corrosive environment in Phase II (long-term), longitudinal cracks parallel to the

corroded reinforcement on one or both lateral surfaces of the specimens were observed. The

corrosion cracks had maximum and average crack widths of 1.3 and 0.8 mm for specimen CRL-

4P-I (Figure 6.7b), 0.7 and 0.45 mm for specimen CRL-4P-II (Figure 6.7c), and 0.25 and 0.16 mm

for specimen CRL-3C-II (Figure 6.7d), respectively. These findings imply that the PBO-FRCM

systems were less effective in decreasing the corrosion rate as compared to the C-FRCM system.

However, both systems significantly enhanced the serviceability of the repaired beams in terms of

crack widths.

6.4.2 Steel Mass Loss

Five steel coupons (200 mm long each) were extracted from the corroded steel bars of the

damaged beams after testing. The actual mass loss of the corroded bars were determined according

to the ASTM G1-03 standards [27]. The average and maximum tensile steel mass loss for each

specimen are listed in Table 6.1. It can be noticed that the steel bars of the long-term specimens

exhibited lower mass loss than those of the short-term ones despite the same duration of corrosion

exposure in both cases. This was attributed to the reduction in the amount of the diffused water

and air to the tensile reinforcing bars due to due to the presence of the FRCM layers and the use

of repair mortar with lower permeability than that of concrete, which also explains the reduction

in the corrosion rate in the long-term specimens, which also explains the reduction in the corrosion

rate in the long-term specimens. For instance, specimens CRL-4P-II and CRL-3C-II showed an

average mass loss 18.1 and 17.5 %, respectively. However, their counterparts in Phase I (short-

term) had an average mass loss of 21.1% (beam CRS-4P-II) and 21.5% (beam CRS-3C-II),

respectively. On the other hand, the FRCM repair scheme had a marginal effect on the corrosion

rate in the long-term beams. Beam CRL-4P-II repaired with Scheme II had an average mass loss

of 18.1 %, which was comparable to the steel mass loss of beam CRL-4P-I repaired with Scheme

I. It is important to mention that the mass loss measurements are highly variable due to the random

110

nature of the corrosion phenomena. Therefore, more research is needed to quantify the effect of

using FRCM composites on the rate of corrosion.

Figure 6.7: Corrosion cracks patterns; a) typical corrosion cracks pattern for short-term

specimens (beam CU); b) beam CRL-4P-I; c) beam CRL-4P-II; and d) beam CRL-3C-II

a)

b)

c)

d)

111

6.4.3 Flexural Cracks Pattern and Failure modes

The virgin beams (UU) and the corroded-unrepaired beam (CU) showed typical crack patterns

and failure modes of under-reinforced beams. They failed at ultimate due to the yielding of steel

bars followed by concrete crushing at the top. Concrete spalling was observed in beam CU due to

crossing of the vertical flexural cracks with the longitudinal corrosion cracks (Figure 6.8a). The

crack patterns and failure modes of the FRCM-repaired beams were highly dependent on the type

of the FRCM system and the repair scheme used. All of the repaired beams failed due to loss of

strengthening action followed by concrete crushing. Based on the test observations, the loss of the

strengthening action took place due to one or a combination of the following modes of failure:

a) Mode A: FRCM delamination at the fabric/matrix interface adjacent to the concrete substrate.

This mode of failure is shown in Figure 6.8b and Figure 6.8c for the beams CRS-4P-I and CRL-

4P-I, respectively.

b) Mode B: partial debonding of the fabric from the matrix accompanied by large slip (more than

3 mm) of the fabric as shown in Figure 6.8d for the beam CRL-4P-II.

c) Mode C: wide flexural cracking (up to 4 mm width) in the cementitious matrix with extensive

fabric slippage at the beam’s soffit in the maximum moment zone. This mode of failure was

encountered in the beams reinforced with C-FRCM and it is shown in Figure 6.8e for the beam

CRL-3C-II.

112

Figure 6.8: Failure mode of a) beam CU due to steel yielding and concrete crushing; b) beam

CRS-4P-I due to FRCM delamination; c) beam CRL-4P-I due to premature FRCM delamination;

d) beam CRL-4P-II due to PBO-fabric debonding from matrix; and e) beam CRL-3C-II due to

fabric slippage

a)

d)

c)

b)

e)

TOP

Side

Side

Side

Bottom

Concrete cover

spalling

Concrete

crushing

FRCM

delamination FRCM delamination

FRCM

delamination Corrosion

cracks

FRCM

delamination

Corrosion

cracks

Partial debonding

Matrix cracking

113

The FRCM delamination (mode A) encountered in beam CRS-4P-I (short-term) was attributed

to the propagation of the flexural cracks in the thin layer of the FRCM matrix adjacent to the

concrete substrate. Its counterpart specimen CRL-4P-I (long-term), having similar FRCM system

and configuration but exposed to post-repair corrosive environment, failed in a more brittle manner

due to the premature delamination of the FRCM composite that was initiated by the longitudinal

corrosion cracks (Figure 6.8c). It should be pointed out that the post-repair corrosive environment

resulted in horizontal corrosion cracks parallel to the reinforcing steel, which notably impaired the

concrete cover and consequently limited the strengthening effectiveness of the FRCM system. The

effect of the post-repair exposure to corrosion was also reflected on the load-carrying capacity of

the long-term beam and on its flexural response as will be presented in the following sections.

On the other hand, the short-term beam CRS-4P-II, which was repaired in the U-shaped FRCM

scheme, failed due to the partial debonding of the PBO fabric from the FRCM matrix (mode B).

As the applied load increased, vertical flexural cracks were formed in the FRCM matrix. The

extension of the cracks in the shear spans was followed by gradual slippage of the fabric from

within the matrix until the strengthening action of the FRCM system was lost. This mode of failure

was also reported at failure for the long-term beam CRL-4P-II.

Similarly, both beams repaired with carbon FRCM system (CRS-3C-II and CRL-3C-II) failed in

a similar mode (mode C). Vertical flexural cracks were observed in the matrix of the U-shaped C-

FRCM layer after steel yielding. As the applied load increased, new flexural cracks developed in

the maximum moment zone accompanied with the widening of the existing cracks. At ultimate, it

was observed that wide flexural cracks formed within the FRCM matrix followed by a noticeable

fabric slippage at the beam’s soffit as shown in Figure 6.8e.

Based on these test observations, it can be concluded that the exposure to the post-repair corrosive

environment had a slight impact on the mode of failure of the beams repaired with scheme II

regardless of the type of fabric used. This can be attributed to the wrapping effect of the FRCM

layer that offset the effect of corrosion cracks in weakening the concrete cover and prevented the

premature delamination of FRCM. On the contrary, the long-term beam CRL-4P-I repaired in

scheme I demonstrated large corrosion cracks that significantly affected the strengthening action

of the FRCM system and led to the premature delamination of FRCM (Figure 6.8c).

114

6.4.4 Flexural Response

The load-displacement curves of the repaired specimens are shown in Figure 6.9. The flexural

response of the virgin beam (UU) and the corroded-unrepaired beam (CU) are also shown for

comparison. Based on the test observations, corrosion of the main reinforcement by an average

mass loss of 22.5% did not have a notable effect on the beam stiffness. For all of the tested beams,

the load-deflection curves consisted of three segments with two turning points indicating the initial

cracking in concrete and the yielding of the tensile steel bars. The flexural response of the repaired

beams was highly dependent on the FRCM repair scheme and its fabric type. It is important to

note that the equivalent axial stiffness of the C-FRCM (3 layers) was approximately 1.5 times of

that of PBO-FRCM (4 layers).

Figure 6.9a and 6.9b illustrate the load versus the mid-span deflection for the short-term and

long-term beams, respectively. The use of FRCM had a slight influence on the stiffness of the

repaired beams prior to steel yielding. However, the post-yielding stiffness significantly increased

in the repaired beams in comparison to the control ones. It can be noticed that beams repaired with

PBO-FRCM (CRS-4P-I and CRS-4P-II) showed lower post-yielding stiffness than that of their C-

FRCM repaired counterparts (Figure 6.9a).

Exposing the beams to corrosion after repair did not affect their load-deflection response as

shown in Figure 6.9c and 6.9d. This can be depicted by comparing the response of the short-term

beams and their long-term counterparts. Both beams repaired with C-FRCM (CRS-3C-II and CRL-

3C-II) showed a sudden drop at ultimate, which indicated their brittle mode failure due to the

sudden cracking in the FRCM matrix and the large fabric slippage within the matrix. The PBO-

repaired beams exhibited different trends that were characterized by a gradual declining branch

after reaching the ultimate load as shown for beams CRS-4P-I and CRS-4P-II and their long-term

counterparts. This was attributed to the different modes of failure of the PBO and C-FRCM

repaired beams as previously described.

115

a)

b)

116

c)

d)

Figure 6.9: Load-deflection relationships of a) short-term beams; b) long-term beams; c) beams

repaired with PBO-FRCM (short-term and long-term); and d) beams repaired with C-FRCM

(short-term and long-term)

6.4.5 Load-carrying Capacities

The load-carrying capacities of the tested beams are shown in Table 6.5. The test results indicate

that an average mass loss of 22.5 % due to corrosion resulted in reduction in the yield and ultimate

strengths by 15% and 10%, respectively. The insignificant effect of corrosion on the yield and

ultimate strengths of RC beams might be attributed to the fact that the corroded steel bars lost their

117

lugs, which increased the measured mass loss without notable loss of the effective cross section

area of the bars. Repairing the corroded RC beams with externally bonded FRCM composites had

restored, and in some cases exceeded, the yield and ultimate strengths of the virgin beam. The

enhancement in the ultimate load of the repaired beams ranged between 14% and 65% of that of

the virgin beam. The experimental results of the yield load, 𝑃𝑦𝑒𝑥𝑝

, and the ultimate load, 𝑃𝑢𝑒𝑥𝑝

, were

normalized with respect to those of the virgin specimens as presented in Table 5.

It is important to note that the efficiency of the FRCM systems in strengthening the damaged

beams was evaluated based on their equivalent axial stiffness and the gain in the ultimate capacity

of the repaired beams. Four layers of PBO-FRCM had an equivalent stiffness of 95.3 MPa

compared to 139.1 MPa for three layers of C-FRCM. For the short-term beams, the use of four

PBO-FRCM layers with repair Scheme I in beam CRS-4P-I restored 110 and 129% of the yield

and ultimate strengths of the virgin beam, respectively. The use of similar number of PBO layers

in Scheme II restored 108 and 139% of the yield and ultimate strengths of the virgin beam,

respectively. The enhancement in the ultimate capacity when Scheme II was used was attributed

to the effect of the U-shaped layer on delaying the delamination of the flexural FRCM plies and

the contribution of the PBO strands on the lateral sides. Consequently, the load-carrying capacity

of the beam CRS-4P-II exceeded that of its counterpart CRS-4P-I.

Table 6.5: Strength results of the tested beams

Specimen 𝑃𝑦

𝑒𝑥𝑝

KN


KN

Normalized loads** 𝑃𝑢𝑝𝑟𝑒𝑑

KN


𝑃𝑢𝑝𝑟𝑒𝑑

𝑃𝑦𝑒𝑥𝑝


UUa, UUb* 75.1 79.7 1 1 81.9 0.97

CU 64.19 72.2 0.85 0.90 64.7 1.11

CRS-4P-I 82.47 102.8 1.1 1.29 86.9 1.18

CRS-4P-II 80.87 111.1 1.08 1.39 93.1 1.19

CRS-3C-II 75.86 109.3 1 1.37 97.8 1.11

CRL-4P-I 79.52 91.1 1.06 1.14 89.9 1.01

CRL-4P-II 83.15 108.1 1.11 1.36 94.7 1.14

CRL-3C-II 81.87 131.9 1.09 1.65 100.8 1.3


118

Exposing the beam CRL-4P-I to further corrosion after repair decreased its yield and ultimate

strengths compared to those of beam CRS-4P-I. It is worth noting that these beams had average

steel mass losses of 18.7 and 22.7%, respectively. The beam CRL-4P-I restored 106 and 114% of

the yield and ultimate strengths of the virgin beam, respectively (Table 6.5 and Figure 6.8c). This

reduction in the strength gain can be attributed to the presence of longitudinal corrosion cracks

with widths up to 1.3 mm in the beam as observed during the test. These wide corrosion cracks

weakened the concrete substrate and caused premature delamination of the FRCM layer, which

significantly limited the strengthening contribution of the FRCM system.

On the contrary, exposing the PBO-repaired beam in Scheme II to post-repair corrosion had no

effect on its yield and ultimate strengths. The use of four layers of PBO-FRCM with Scheme II in

specimen CRL-4P-II increased its yield and ultimate strengths by 11 and 36%, respectively, in

comparison to 8 and 39% for its counterpart short-term specimen CRS-4P-II. A similar trend was

reported for the specimens repaired with CFRCM in Scheme II. Specimen CRL-3C-II showed an

increase of 9 and 65% of its yield and ultimate strengths, respectively, compared to 0 and 37% for

its counterpart specimen CRS-3C-II. The higher enhancement in the ultimate load-carrying

capacity reported in the former specimen (beam CRL-3C-II) can be attributed to the lower mass

loss reported in the steel bars compared to the mass loss in specimen CRS-3C-II (17.5% mass loss

versus 21.5%, respectively). It can also be attributed to the long curing period of the long-term

beam during the post-repair corrosion exposure.

6.4.6 Ductility and Energy Absorption

The ductility index, ΔI, and the energy absorption index, ψ, for each beam are listed in Table 6.6.

The ductility index is defined as the ratio of the mid-span deflection at ultimate, δu, to its mid-span

deflection at yielding, δy, whereas the energy absorption index, ψ, is defined as the area under the

load-deflection curve up to the ultimate load.

119

Table 6.6: Ductility and energy absorption of the tested beams

Specimen

Midspan

deflection, mm Ductility index

Energy absorption

index, KN.mm

δy δu ΔI ΔInorm** ψ ψnorm

**

UUa, UUb* 11.7 32.9 2.81 1 1736 1

CU 9.8 27.06 2.8 1 1735 1

CRS-4P-I 12.04 33.53 2.78 0.99 2657 1.53

CRS-4P-II 12.06 31.32 2.6 0.94 2426 1.4

CRS-3C-II 12.11 30.16 2.49 0.89 2207 1.27

CRL-4P-I 12.74 26.07 2.05 0.73 1780 1.03

CRL-4P-II 12.04 39.99 3.32 1.18 3360 1.94

CRL-3C-II 10.03 31.81 3.17 1.13 2895 1.67

* Average values reported **Normalized with respect to the virgin beam

The corroded-unrepaired specimen (CU) had a ductility index similar to that of the control

specimen. The short-term beams repaired with PBO-FRCM (CRS-4P-I and CRS-4P-II) showed a

ductility index almost similar to that of the virgin beam. The use of C-FRCM in specimen CRS-

3C-II reduced the ductility index by 11% of that of the virgin specimen. On the other hand, long-

term specimens (CRL-4P-II and CRL-3C-II) showed an increase in their ductility indices by 18%

and 13%, respectively, in comparison to that of the control specimen. This was attributed to the

long curing period of the repaired specimens in Phase II. On the contrary, a significant reduction

in the ductility index (27% of that of the control specimen) was determined for the long-term

specimen CRL-4P-I. This was attributed to the premature FRCM delamination as previously

described.

The control and the corroded-unrepaired specimens had similar energy absorption indices. For

the short-term specimens, those repaired with PBO-FRCM (CRS-4P-I and CRS-4P-II) exhibited

an average energy absorption index 15% higher than that of specimens repaired with C-FRCM

(CRS-3C-II). For specimens exposed to post-repair corrosion (long-term specimens), the energy

absorption indices of specimens CRL-4P-II and CRL-3C-II were higher than those reported for

their short-term counterparts by 39 and 31%, respectively. However, specimen CRL-4P-I showed

120

a notable reduction in the energy absorption index by 33% of that of its short-term counterpart

CRS-4P-I. These results are consistent with the modes of failure reported as previously discussed.


Table 6.7 lists the strains measured at mid-span in the fiber, the steel bars, and the compressed

concrete at ultimate for all of the tested specimens. At failure, the strains in the steel bars were

2428 μɛ for the control specimen (UU) while strains ranged from 1946 to 3490 μɛ for the beams

repaired with PBO-FRCM. Higher tensile steel strains (3402 to 4906 μɛ) were reported for the

specimens repaired with C-FRCM.

Figure 6.10a and Figure 6.10b show the load versus the recorded strains in both the concrete and

fiber at mid-span for specimens repaired with PBO-FRCM and C-FRCM, respectively. Prior to

yielding, the concrete strain increased with the same rate in all of the tested specimens. Following

the steel yielding, the virgin beam (UU) and the corroded-unrepaired beam (CU) showed an almost

plastic response until failure occurred by concrete crushing. This response was indicated by the

increase in the recorded strains in concrete in both beams. The use of FRCM composite caused a

notable strain hardening in concrete of the repaired specimens as depicted in Figure 6.10. Concrete

strain hardening in specimens repaired with C-FRCM was higher than that in those repaired with

PBO-FRCM. This finding was consistent with the increase in the post-yielding stiffness

encountered in the former specimens, as discussed earlier. The highest absolute concrete

compressive strain was reported in specimen UU (-3318 μɛ) while the lowest absolute strain was

observed in specimen CRL-4P-I (-1520 μɛ). This observation might be attributed to the premature

FRCM delamination initiated from the post-repair corrosion cracks that resulted in a sudden failure

in specimen CRL-4P-I. The concrete strains recorded in other specimens ranged between -2291 to

-3239 μɛ at ultimate.

121

Table 6.7: Strains at ultimate

Specimen Fiber strains,

µ𝜖

Concrete strains,

µ𝜖

Steel strains,

µ𝜖

UUa, UUb* -3318 2428

CU - -2338 -

CRS-4P-I 8446 -3239 2422

CRS-4P-II 9653 -2992 2928

CRS-3C-II 9777 -2469 4906

CRL-4P-I 7409 -1520 1946

CRL-4P-II 8928 -2296 3490

CRL-3C-II 13876 -2291 3402

* Average values reported

On the other hand, the outer fabric tensile strain increased as the applied load increased. Once

the tensile steel bars reached the yielding point, the fiber strains increased at higher rate than that

observed prior to yielding (Figure 6.10). This result was consistent with the limited effect of the

FRCM system on the beam performance prior to yielding. Exposing the beams to post-repair

corrosion had a marginal effect on the ultimate tensile fiber strain reported for specimens repaired

with PBO-FRCM. Specimens CRS-4P-II and CRL-4P-II showed ultimate fiber strains of 9653 and

8928 μɛ, respectively. Also, the fiber strains measured in specimens CRS-4P-I and CRL-4P-I were

8446 and 7409 μɛ, respectively. The low strain that was encountered in beam CRL-4P-I was

attributed to the premature FRCM delamination initiated from the post-repair corrosion cracks,

which negatively affected the FRCM system strengthening potency. On the other hand, specimens

CRS-3C-II and CRL-3C-II showed ultimate fiber strains of 9777 and 13879 μɛ, respectively. The

increase in the measured fiber strains in the long-term beam CRL-3C-II was consistent with its

higher ultimate load than its short-term counterpart.

122

a)

b)

Figure 6.10: Load-strains relationships for a) beams repaired with PBO-FRCM and b) beams

repaired with C-FRCM

123

It is important to note that the ultimate fiber strain, 휀𝑓𝑢, obtained from the direct tensile tests

(Table 4) conducted on coupons made of the same FRCM used in this study was 14000 μɛ for

PBO-FRCM and 12500 μɛ for C-FRCM [97]. These fiber strains are obviously higher than those

measured during the beam tests. Therefore, relying on the ultimate strains in the fabric may be

misleading in predicting the performance of the repaired beams while assuming perfect bond

between the concrete substrate and the FRCM system. Unlike the fabric slippage that occurred in

the FRCM tensile coupons at ultimate, different failure mechanisms were encountered in the

FRCM-repaired beams. Therefore, the assumption of prefect bond suggested by the ACI 549.4R-

13 [54] while limiting the design effective strain, 휀𝑓𝑒, in the FRCM system is a simplification that

appears justifiable and easy to implement by engineers.

6.5 Predicted Strengths

The predicted ultimate loads, 𝑃𝑢𝑝𝑟𝑒𝑑

, were determined according to the provisions of ACI 318-14

[98] and ACI 549.4R-13 [54]. Perfect bond between FRCM composites and the concrete substrate

was assumed while the fiber strain was limited to 0.012 mm/mm as recommended by ACI 549.4R-

13 [54]. Therefore, the design effective tensile strain, 휀𝑓𝑒, in FRCM was assumed equal to the

experimental ultimate strain, 휀𝑓𝑢, as obtained from the coupon test results minus one standard

deviation or 0.012 mm/mm, whichever was lower. The design effective tensile strength, 𝑓𝑓𝑒, was

taken equal to 𝐸𝑓휀𝑓𝑒, where 𝐸𝑓 is the cracked tensile modulus of FRCM as listed in Table 6.4. For

the beams repaired with PBO-FRCM in Scheme II, the longitudinal fiber strands on the lateral

sides of the U-shaped layer were considered in estimating the flexural strength. The corrosion

damage of the bottom steel bars was presented by a reduction in the cross-section areas of the bars,

calculated based on the measured average mass loss for each beam as given in Table 6.1. The yield

strengths of the longitudinal reinforcing steel bars of diameter 15 and 8 mm were taken equal to

466 and 573 MPa, respectively, with elastic modulus 𝐸𝑠 = 200 GPa. The concrete compressive

strength was taken equal to 41.8 MPa with maximum compression strain equal to 0.0035 mm/mm.

The predicted ultimate load, 𝑃𝑢𝑝𝑟𝑒𝑑

and the ratios of the experimental to predicted ultimate

loads 𝑃𝑢𝑒𝑥𝑝/𝑃𝑢

𝑝𝑟𝑒𝑑, for all of the tested beams are presented in Table 6.5. The ratio 𝑃𝑢

𝑒𝑥𝑝/𝑃𝑢𝑝𝑟𝑒𝑑

for all repaired specimens ranged between 1.01 and 1.3 indicating a very good agreement between

124

the experimental and predicted values. This finding revealed that the load-carrying capacities of

the repaired beams in-service, i.e. those exposed to further corrosive environment after repair, can

be conservatively estimated using ACI 549.4R-13 provisions [54].

6.6 Conclusions

The short and long-term flexural performances of concrete beams repaired with FRCM systems

are presented in this study. The test results have evidenced the following conclusions:

• An average mass loss of 22.5% in the tensile steel bars reduced the yield and the ultimate loads

of the corroded beams by 15% and 10%, respectively, without a notable impact on the beam’s

stiffness or mode of failure.

• The use of FRCM systems reduced the corrosion rate in the steel bars with no evidence on the

effect of the repair scheme on such rate. Exposing the FRCM-repaired beams to post-repair

corrosion resulted in 23% reduction in the steel mass loss.

• Most of the corroded-repaired specimens that were exposed to post-repair corrosive

environment failed to meet the serviceability provisions of the ACI-318-14 for crack widths.

• Repairing the corroded beams with FRCM systems enhanced their flexural behavior and

increased their load-carrying capacities between 14 to 65% of that of the virgin beam.

• The U-wrapped scheme was more efficient than the end-anchoring scheme in delaying the

delamination of the FRCM plies in the short-term repaired beams. It also mitigated the effect

of the longitudinal corrosion cracks and consequently increased the post-repair strengthening

effectiveness of FRCM systems.

• Short-term beams repaired with PBO-FRCM exhibited lower post-yielding stiffness and more

ductility at failure than those of their carbon-repaired counterparts. Average ductility indices

were 97 and 89% of that of the control specimen with average energy absorption indices of

147 and 127%, respectively.

• Long-term beams repaired with scheme II demonstrated higher ductility and energy absorption

indices than those of their short-term counterparts. The short-term beam repaired with scheme

I (CRL-4P-I) failed prematurely due to premature FRCM delamination.

125

• Strain values recorded on the FRCM layer indicated that the assumption of prefect bond

suggested by the ACI 549.4R-13 while limiting the design effective strain in the FRCM system

is justifiable.

• The ACI 549.4R-13 provisions conservatively predict the ultimate capacities of the FRCM-

repaired beams exposed to post-repair corrosive environment.

It is important to note that the results of this study are only applicable to the FRCM systems used

and should not be extrapolated to other systems. Further experimental and analytical studies using

other commercially available FRCM systems are recommended. Tests on RC beams subjected to

higher corrosion levels corrosion levels are also needed to assess the proposed repair technique in

more severe environments.

126

7. Chapter 7

Fatigue and Monotonic Behavior of Corrosion-damaged

Reinforced Concrete Beams Strengthened with FRCM

composites


Journal of Composites for Construction, ASCE. Submitted in the revised form: Fabraury 12, 2018


Résumé

Cet article fournit un compte-rendu complet de l'utilisation de composites à matrice cimentaire

renforcés de fibres (MCRF) pour renforcer les structures en béton armé endommagées par la

corrosion et soumises à la fatigue. Douze poutres ont été construites et testées à la rupture dans

une configuration de chargement à quatre points. Avant les essais, dix poutres ont été soumises à

une corrosion accélérée pendant 140 jours, ce qui a entraîné une perte de masse moyenne de 19%

dans le renforcement en acier. Huit poutres endommagées par la corrosion ont été renforcées et

testées tandis que les deux autres poutres n'ont pas été renforcées. Deux autres poutres vierges qui

n'ont pas été soumises à la corrosion ont été utilisées comme témoins. Les paramètres d'essai

comprenaient le matériau de fibre (PBO et carbone), le nombre de couches MCRF, la configuration

de renforcement et le type de chargement (monotone et fatigue). Les résultats des tests ont montré

que la corrosion des barres d'acier diminuait considérablement la durée de vie en fatigue des

poutres. Après le renforcement, les poutres endommagées par la corrosion ont entièrement restauré

la capacité de charge des poutres vierges. Les poutres renforcées par MCRF ont subi plus de cycles

de fatigue que leurs contreparties non renforcées, mais n'ont pas pu restaurer la durée de vie en

fatigue des poutres vierges. L'effet de la configuration de MCRF était plus prononcé en fatigue

que dans les tests monotones. Le PBO-MCRF était plus efficace que le C-FRCM à améliorer la

performance en fatigue des poutres endommagées par la corrosion.

Mots clés des auteurs : Corrosion; Matrice cimentaire renforcée de fibres; Fatigue; Flexion;


127

7.1 Abstract

This paper provides a comprehensive account of using fabric-reinforced cementitious matrix

(FRCM) composites to strengthen corrosion-damaged reinforced concrete (RC) structures

subjected to fatigue. Twelve beams were constructed and tested to failure under four-point loading

configuration. Prior to testing, ten beams were subjected to accelerated corrosion for 140 days

leading to an average mass loss in the steel reinforcement of 19%. Eight corrosion-damaged beams

were strengthened and tested while the other two beams remained unstrengthened. Two other

virgin beams that were not subjected to corrosion were used as benchmarks. The test parameters

included the fabric material (PBO and Carbon), the number of FRCM plies, the strengthening

configuration, and the type of loading (monotonic and fatigue). Test results showed that corrosion

of steel bars dramatically decreased the fatigue life of the beams. After strengthening, the

corrosion-damaged beams fully restored the load-carrying capacity of the virgin beam. The

FRCM-strengthened beams endured more load cycles than their unstrengthened counterpart, but

could not restore the original fatigue life of the virgin beam. The effect of FRCM configuration

was more pronounced in fatigue than in monotonic tests. PBO-FRCM was more effective than the

C-FRCM composite in enhancing the fatigue performance of the corrosion-damaged beams.

Authors’ keywords: Cementitious Materials; Corrosion; Fabric-Reinforced Cementitious

Matrix; Fatigue; Flexure; Repair; Strengthening.


Transportation infrastructures such as bridges are prone to corrosion of their steel reinforcement

due to the harsh environment and the heavy use of the de-icing chemicals in cold regions. These

structures are continuously subjected to cyclic loads, which may cause fatigue distress and

consequently reduce their anticipated life. It has been established that the fatigue life of concrete

structures is usually governed by the endurance of the steel reinforcing bars under repetitive loads

and is rarely controlled by concrete [50–52]. While fatigue failures are not common in concrete

structures, the corrosion of steel reinforcement combined with fatigue stresses significantly

128

reduces the fatigue life of the structure [53,107], which necessitates immediate intervention with

appropriate strengthening measures.

In the last decades, externally-bonded strengthening technologies based on organic matrices and

referred to as fiber-reinforced polymers (FRP) have become a common practice. Numerous studies

have investigated the behavior of corrosion-damaged RC beams strengthened with FRP

composites under monotonic and fatigue loads [90,108]. Test results demonstrated that the use of

FRP composites successfully restored the yield and ultimate strengths of the beams, delayed the

cracks propagation and reduced their widths, and increased the fatigue life of the strengthened

beams. The enhancement in fatigue life of the FRP-strengthened beams was attributed to the ability

of FRP composites to decrease the stresses in the steel reinforcing bars [10].

Recently, fabric-reinforced cementitious matrix (FRCM) composites were proposed as

promising alternatives to FRP composites. FRCM composites consist of one or more layers of

carbon, glass, or Polyparaphenylene benzobisoxazole (PBO) fabrics that are sandwiched between

layers of cementitious mortars. Unlike FRPs, FRCM composites are believed to have

environmental acceptability, good thermal and fire resistance, and the advantage of application on

wet surfaces and at low temperature. They are also compatible with the concrete substrate and are

characterized by their ease of installation and long-term durability [57,92,93,106].

Multiple studies have documented the effectiveness of FRCM composites in enhancing the

flexural response of undamaged RC structures under monotonic loads [55,58,59,82,85]. However,

very few studies have been devoted to investigate the fatigue performance of the FRCM-

strengthened structures. Aljazaeri and Myers (2015) [87] reported that PBO-FRCM composites

not only improved the fatigue performance of RC beams but also enhanced their residual flexural

strength after fatigue. Moreover, exposing the PBO-FRCM strengthened beams to high

temperature and humidity did not affect their fatigue performance. Yin et al. (2014) [109]

demonstrated that the use of FRCM composite with hybrid (carbon and E-glass) fabric

significantly increased the fatigue life of undamaged RC beams. The enhancement in the fatigue

life was dependent on the fiber reinforcement ratio. More recently, Pino1a et al. (2016) [88]

demonstrated the effectiveness of PBO-FRCM composites in improving the fatigue life of

undamaged RC beams. However, the fatigue life of the strengthened beams decreased by

increasing the maximum applied loads due to the rupture of the steel reinforcement.

129

The feasibility of using FRCM composites to strengthen corrosion-damaged RC structures has

received little attention. Surface deterioration due to corrosion and the absence of a sound concrete

substrate represent major challenges to the application and the long-term durability of the FRCM

composites. Therefore, the flexural behavior of corrosion-damaged structures strengthened with

FRCM composites has rarely been investigated, not to mention their behavior when subjected to

repetitive loads. In a recent study, El-maaddawy and Refai (2016) [86] used basalt and C-FRCM

composites to restore the flexural capacity of RC beams after being subjected to corrosion. The

FRCM composites were either internally embedded within the corrosion-damaged region or

externally-bonded along the beam span. It was reported that externally-bonded composites were

more effective in restoring the ultimate capacities of the corrosion-damaged beams. Nevertheless,

the authors reported that other investigations should be conducted before field applications could

be recommended.

The present study is part of a large research program that aims at investigating the behavior of

corrosion-damaged RC structures after being strengthened with FRCM composites. It provides

insight into the monotonic and fatigue performance of corrosion-damaged beams strengthened

with FRCM composites having different materials and configurations. The test results reported

herein provide unique experimental data and represent the first research work on the use of FRCM

in strengthening corrosion-damaged beams subjected to fatigue.


Twelve large-scale RC beams were fabricated and divided into two groups. Group ‘M’ consisted

of six beams that were tested under monotonic loading up to failure whereas group ‘F’ consisted

of six beams that were tested in fatigue. The beams of each group included four corrosion-damaged

beams strengthened with different FRCM composites and configurations. The other two beams

included one beam that was corroded but not strengthened while the other beam was neither

corroded nor strengthened and acted as control.

Table 7.1 and Table 7.2 list the matrix of both the monotonic and fatigue tests, respectively. The

beams were labeled following the A-B-C format. ‘A’ represents the loading regime (F for fatigue

and M for monotonic) and the beam condition (UU, CU, and CS referring to Uncorroded-

Unstrengthened, Corroded-Unstrengthened, and Corroded-Strengthened, respectively). ‘B’

130

denotes the number and type of the FRCM layers applied (2P, 4P, and 3C referring to two layers

of PBO-FRCM, four layers of PBO-FRCM, and three layers of C-FRCM, respectively). Finally,

‘C’ describes the FRCM strengthening schemes (I and II) as will be detailed in the following

sections. For instance, the label FCS-4P-II refers to a corrosion-damaged beam (C) strengthened

(S) with four layers of PBO-FRCM composites (4P) in scheme (II) and tested under fatigue loading

(F).

7.3.1 Test Specimen

The geometry and the reinforcement details of the test specimen are shown in Figure 7.1. All

beams were 2.8 m long with a rectangular cross section 150 mm in width and 250 mm high. The

beams were reinforced with 2-15M deformed steel bars placed at a cover distance of 25 mm from

the soffit. At the top, 2-8M deformed bars were used as stirrups hangers. All beams were provided

with hollow stainless-steel tubes with external and internal diameters of 9.5 mm and 7 mm,

respectively, placed at 80 mm from the soffit of the beam as shown in Figure 7.1. The shear spans

were reinforced with double-leg steel stirrups of 10 mm diameter each spaced at 100 mm in order

to prevent premature shear failure.

Normal and salted concrete mixes having the same water/cement ratio were used to cast the

beams. Concrete cylinders (150×300 mm) were also cast from both mixes to evaluate their

compressive strengths. The average compressive strengths were 41.8 MPa (standard of deviation

of 4.8 MPa) for the normal mix and 41.2 MPa (standard deviation of 0.6 MPa) for the salted one.

The reinforcing steel bars had a nominal yield strength of 400 MPa and an elastic modulus of 200

GPa as reported by the manufacturer. The yield strengths of the longitudinal reinforcing steel bars

of diameter 15 and 8 mm were 466 MPa (with a standard deviation of 4.2 MPa) and 573 MPa

(with a standard deviation of 17.7 MPa), respectively, as tested by the authors.

131

Table 7.1: Monotonic test matrix and results

* Nornalized to the control specimen (MUU)

** SY-CC = Steel Yielding followed by Concrete Crushing; FD = FRCM Delamination; FS-PED = Fabric Slippage followed by Partial Debonding; MC-FS = Matrix Cracking followed by Fabric Slippage.

Table 7.2: Fatigue test matrix and results

* SY-CC = Steel Yielding followed by Concrete Crushing; FRS = Fatigue Rupture of Steel bars

Specimen Avg. mass

loss (%)

𝜌𝑓

(%)

𝐾𝑓

(MPa)

𝜌𝑆

(%)

𝛽𝑓

(%)

𝑃𝑦

(KN)

𝑃𝑢 (KN)

𝑃𝑦

Nor.*

𝑃𝑢

Nor.* εfu

(µϵ)

Mode of

Failure**

MUU - - - 1.07 - 75.1 79.7 1 1 - SY-CC

MCU 18 - - 0.87 - 64.5 74.2 0.86 0.93 - SY-CC

MCS-2P-I 19.6 0.04 48 0.86 2.8 71.8 85.6 0.96 1.07 8180 FD

MCS-4P-I 19.4 0.08 95.3 0.86 5.54 79.6 102.6 1.06 1.29 10659 FD

MCS-4P-II 19.5 0.08 95.3 0.86 5.55 80.7 102.9 1.07 1.29 8253 FS-PFD

MCS-3C-II 18.6 0.19 139.1 0.87 8.01 78.8 123.3 1.05 1.55 5530 MC-FS

Specimen Avg. mass

loss (%)

𝜌𝑓

(%)

𝐾𝑓

(MPa)

𝜌𝑆

(%) 𝛽𝑓 (%)

Fatigue life

(cycles)

Mode of

Failure*

FUU - - - 1.07 - Over 2

million SY-CC

FCU 19.8 - - 0.86 - 396,000 FRS

FCS-2P-I 18.4 0.04 48 0.87 2.76 545,600 FRS

FCS-4P-I 19.3 0.08 95.3 0.86 5.54 984,800 FRS

FCS-4P-II 18.1 0.08 95.3 0.87 5.48 1,493,300 FRS

FCS-3C-II 18.6 0.19 139.1 0.87 7.99 834,500 FRS

132

Figure 7.1: Geometry and reinforcement details of the test specimen (all dimensions in mm)

7.3.2 Accelerated Corrosion Technique

An accelerated corrosion technique was used to corrode the tensile steel reinforcement. A

constant electric current of 380 milliamps was impressed through the steel bars. The current

density was of 180 µA/cm2 to obtain corrosion products that resemble to those found in natural

corrosion process [20]. Salt (NaCl) measured as 5% of the cement weight was added to the

concrete mix used to cast the middle-bottom of the beams at a height of 100 mm as shown in

Figure 7.1. Corrosion of steel bars was restricted to the middle 1200 mm of the beam’s span. The

steel bars and the stirrups outside the salted zone were coated with anti-corrosion epoxy for

protection.

During the corrosion process, the salted concrete acted as electrolyte while the bottom steel bars

and the hollow stainless-steel tube acted as anode and cathode, respectively. The beams were

placed in a large environmental chamber and were connected in series with a DC galvanostatic

power supply to ensure that the induced current is uniform in all specimens (Figure 7.2). The

beams were subjected to wet and dry cycles that consisted of 3 days wet followed by 3 days dry.

The wet and dry cycles provided water and oxygen necessary for the corrosion process. The

accelerated corrosion process lasted for 140 days to obtain a theoretical steel mass loss 20% in the

main reinforcement according to Faraday’s law.

133

Figure 7.2: Specimens connected in series inside the corrosion chamber

7.3.3 FRCM Composites

Two different FRCM composites (PBO and carbon) were used. PBO-FRCM consisted of an

unbalanced PBO-fabric impregnated in a cementitious matrix with a low dosage of dry polymer.

The fabric was made of spaced fiber rovings prearranged along two orthogonal directions as shown

in Figure 7.3a. The width of rovings was 5 mm and 2.5 mm in the main and secondary directions,

respectively and the nominal thickness is 0.046 mm in the main direction, and 0.011mm in the

secondary direction. The gap opening between the rovings was approximately 5 mm in the main

direction and 15 mm in the secondary direction. Table 7.3 lists the properties of the fabric as

reported in the manufacturers’ data sheet. The cementitious matrix of the PBO-FRCM composite

had a compressive strength of 43.9 MPa (standard deviation of 0.4 MPa) and a flexural strength

of 3 MPa (standard deviation of 0.3 MPa) after 28 days as determined by the authors.

The C-FRCM composite consisted of carbon fabric impregnated in an inorganic cementitious

matrix. The fabric was a unidirectional net made of carbon strands oriented in one direction as

shown in Figure 7.3b. The strands were uniformly distributed at a density of 59 strands per meter

width. The properties of the carbon fabric according to the data sheet provided by the manufacturer

are given in Table 7.3. The cementitious matrix had a compressive strength of 42.1 MPa (standard

deviation of 4.3 MPa) and flexural strength of 3.2 MPa (standard deviation of 0.3 MPa) after 28

days as tested by the authors.

134

Figure 7.3: a) Unbalanced PBO fabric and b) Unidirectional carbon fabric.

Table 7.3: Fabric properties in the primary direction as given in the manufacturers’ data sheet


𝐴𝑓 (mm2/m) Tensile strength

(GPa)

Elastic modulus

(GPa)

Ultimate strain

(%)

PBO 50 5.8 270 2.15

Carbon 157 4.3 240 1.75

Table 7.4 lists the mechanical properties of the PBO- and C-FRCM composites used in this study

as obtained from direct tensile tests that were conducted on FRCM coupons by Ebead et al. (2016)

[97]. An axial stiffness coefficient, Kf, given by Equation (7.1), characterized the FRCM

composites based on their cracked elastic modulus, 𝐸𝑓, as listed in Table 7.4, and the fiber

reinforcement ratio, 𝜌𝑓 . Accordingly, the coefficient Kf varied with the number of fabric layers

used. Table 7.1 and Table 7.2 list the fiber reinforcement ratio, 𝜌𝑓, and the axial stiffness

coefficient, Kf, for each of the strengthened beams.

𝐾𝑓 = 𝜌𝑓𝐸𝑓 (MPa) Eq. (7.1)

Similarly, the axial stiffness coefficient for steel reinforcement was calculated as follows:

𝐾𝑠 = 𝜌𝑠𝐸𝑠 (MPa) Eq. (7.2)

Where, 𝜌𝑠 is the tensile steel reinforcement ratio and 𝐸𝑠 is the elastic modulus of steel bars (200

GPa). For corrosion-damaged beams, Table 7.1 and Table 7.2 list the values of 𝜌𝑠 for each beam

taking into account the actual mass loss in the steel bars after corrosion. The contribution of FRCM

135

composites to that of steel reinforcement was expressed by the stiffness factor, 𝛽𝑓, as given in

Equation (7.3). Table 7.1 and Table 7.2 list the values of 𝛽𝑓 for each of the strengthened beams.

𝛽𝑓 = 𝐾𝑓/𝐾𝑠 Eq. (7.3)

Table 7.4: Mechanical properties of FRCM coupons (Ebead et al. 2016) [97]

FRCM

composite

Cracked tensile modulus


Ultimate tensile

strength ffu (GPa)

Ultimate strain

εfu (%)

PBO-FRCM 121 1.55 1.4

Carbon-FRCM 75 0.97 1.25

7.3.4 FRCM Strengthening Configuration

Two schemes of FRCM were used in strengthening the corrosion-damaged beams. In both

schemes, one or more flexural plies of FRCM composites were applied to the beam’s soffit over

the middle 2.4 m of its span. The fabric was oriented such as its main direction was parallel to the

longitudinal axis of the beam. In scheme I, U-shaped transverse strips of FRCM of 300 mm width

were used as end anchors for the flexural plies as shown in Figure 6.4a. According to Ombres

(2011) [80] and Hashemi and Al-Mahaidi (2012) [93], the end anchors avoided the premature

delamination of the concrete cover at the beam ends. In scheme II, the beams were wrapped with

a continuous U-shaped FRCM ply as shown in Figure 6.4b. D’Ambrisi and Focacci (2011) [81]

reported that wrapping the flexural plies with a continuous U-shaped ply delayed the premature

delamination of FRCM composite. The main direction of the continuous U-shaped layer of the

PBO fabric was oriented parallel to the longitudinal axis of the beams and therefore contributed to

the flexural capacity of the strengthened beams MCS-4P-II and FCS-4P-II. On the contrary, the

carbon fabric was a unidirectional fabric and therefore, the continuous U-shaped layer was oriented

in the transverse direction and did not contribute to the flexural capacities of the beams MCS-3C-

II and FCS-3C-II.

136

Figure 7.4: FRCM strengthening schemes: a) Scheme I and b) Scheme II

`

a)

b)

137

7.3.5 Strengthening Procedure

Figure 7.5 shows the strengthening procedure of the corrosion-damaged beams with FRCM

composites. At the end of the corrosion process, the deteriorated concrete was first removed using

a hydraulic hammer as shown in Figure 7.5a. The corroded steel bars were then brushed, and the

beams were repaired using a commercially available cementitious mortar (Figure 7.5b). The repair

mortar used had an average compressive strength of 55.4 MPa (standard deviation of 5 MPa) and

flexural strength of 3.4 MPa (standard deviation of 0.3 MPa) as determined by the authors. After

7 days of curing at ambient temperature, the beam surface was roughened by sandblasting. The

FRCM composites were then installed using the hand lay-up method as proposed by ACI 549.4R-

13 [54]. As recommended by the FRCM manufacturers, the beam substrate was soaked in water

for 2 hours before applying the first layer of the cementitious matrix with a thickness of 3 to 5 mm.

The fabric was then installed and coated with a second layer of matrix of similar thickness (Figure

7.5c and 7.5d). The procedure was then repeated until the specified number of layers was achieved.

All of the strengthened beams were cured for 28 days in laboratory conditions before being tested.

Figure 7.5: FRCM strengthening procedure: a) removing the deteriorated concrete after

corrosion, b) patch repair, c) PBO-FRCM application, and d) C-FRCM application

138


The beams were tested in a four-point bending configuration using a MTS actuator as shown in

Figure 7.6. A steel beam was used to apply the load equally to two loading points placed 800 mm

apart. The beams had an effective span of 2.8 m and shear spans of 880 mm. Additional supports

were used to prevent the out-of-plane movement of the beam during cycling (Figure 7.6). All

beams were instrumented with 60 mm strain gauges bonded to the top surface of concrete in the

compression zone and 5 mm strain gauges bonded to the tensile steel bars at mid-span. Moreover,

the strengthened beams were instrumented with 5 mm strain gauges installed directly on the outer

fabric of the FRCM composite at mid-span and under the point loads. Beam deflections were

measured by means of three linear variable differential transducers (LVDTs) located at mid-span

and under the point loads. A 20-channel data acquisition system was used to capture the readings

of the measuring gauges.

All beams of group M were tested to failure under monotonic loading under displacement control

at a rate of 2 mm/min. The beams of group F were subjected to cyclic loading at a frequency of 2

cycles per second. The minimum and maximum fatigue loads were 17 and 48 KN, respectively,

representing 21 and 60% of the load carrying capacity of the control beam. This load range was

not only chosen to simulate the actual service loading in the field, but also to challenge the fatigue

strength of the corroded steel reinforcement and FRCM composites. The beams were subjected to

fatigue until failure occurred or 2 million cycles were reached. Beams that lasted 2 million cycles

were tested to failure under monotonic loading to determine their residual strengths.

Figure 7.6: Test setup

Spreader beam

Auxiliary support LVDTs

139

7.4 Test Results and Discussion

7.4.1 Corrosion Crack Patterns and Actual Steel Mass Loss

Figure 7.7 shows the crack patterns at the end of the corrosion process for Beam FCU.

Longitudinal cracks formed on the beam sides along the corroded bars and on the beam soffit. The

widths of the corrosion cracks were measured and recorded for all of the damaged beams. The

maximum and average crack widths were 2.8 and 1.28 mm, respectively. It was therefore

concluded that all beams failed to meet ACI 318-14 [98] serviceability requirements that limit the

maximum crack width for a beam in service to 0.40 mm.

Figure 7.7: Corrosion crack pattern for specimen FCU

To determine the actual steel mass loss, the corroded steel bars were extracted after testing.

Several corrosion pits were observed to be randomly distributed on the bars’ surfaces as shown in

Figure 7.8. A steel brush was used to remove the adhered mortar around the bar. The bars were

then cut into coupons of 200 mm length each. The coupons were chemically cleaned and the mass

loss was determined according to ASTM G1-03 provisions [26]. Table 7.1 and Table 7.2 list the

average mass loss as determined for all of the corrosion-damaged beams. An average mass loss of

19% was reported, which was 1% less than that predicted by Faraday’s law.

140

Figure 7.8: Profile of the steel bars: a) Uncorroded bar, b) corroded steel bar extracted from

beam MCS-4P-II, and c) fatigue rupture of a corroded steel bar extracted from FCS-3C-II

7.4.2 Monotonic Test Results

7.4.2.1 Modes of Failure

Figure 7.9 illustrates the modes of failure of the beams tested under monotonic loads. The control

Beams MUU and MCU failed due to yielding of the steel bars followed by concrete crushing in

compression (Figure 7.9a). On the other hand, PBO-strengthened Beams MCS-2P-I and MCS-4P-

I failed due to the delamination of the FRCM composites at the fabric/matrix interface. This mode

of failure can be depicted in Figure 7.9b for Beam MCS-4P-I. From the test observations, FRCM

delamination was attributed to the propagation of the flexural cracks to the matrix layers adjacent

to the concrete substrate.

The use of continuous U-shaped layer of PBO-FRCM in Scheme II changed the mode of failure

of Beam MCS-4P-II. As the load increased, cracks formed in the matrix of the continuous U-

shaped layer followed by gradual slippage of the fabric from the surrounding matrix. Close to

ultimate, partial debonding of the fabric was observed at the fabric/matrix interface (Figure 7.9c).

On the other hand, Beam MCS-3C-II strengthened with three layers of C-FRCM in Scheme II

failed due extensive cracking in the matrix of the FRCM composite as the applied load increased.

Matrix cracking resulted in slippage of the carbon fabric from the matrix as shown in Figure 7.9d.

c)

a)

b)

141

This mode of failure explained the sudden failure of Beam MCS-3C-II when compared to that of

Beam MCS-4P-II.

As previously stated, fabric strains were measured by means of strain gauges located at midspan.

Table 7.1 lists the ultimate fabric strain, εfu, for all of the strengthened beams. Strain values should

be looked at based on the ultimate capacity and the mode of failure encountered for each beam.

The ultimate fabric strains reported for the PBO-strengthened beams ranged between 8180 and

10659 μɛ, while the strains reported at ultimate for Beam MCS-3C-II was only 5530 μɛ, which

was consistent with the mode of failure of the Beam MCS-3C-II where premature matrix cracking

and fabric separation were observed.

Figure 7.9: Failure modes of the monotonically tested beams: a) Steel yielding followed by

concrete crushing; b) FRCM delamination; c) Fabric slippage and partial debonding; and d)

Matrix cracking followed by fabric slippage

7.4.2.2 Load-deflection and Ultimate Strengths

The load-deflection response of the monotonically tested beams (Group M) are shown in Figure

7.10. The yield load, 𝑃𝑦, and the ultimate load, 𝑃𝑢, of the tested beams are listed in Table 7.1.

Corrosion of steel bars had a marginal effect on the load-deflection response of the unstrengthened

Beam MCU. However, the yield and ultimate strengths after corrosion were reduced by 14 and

7%, respectively.

142

Figure 7.10: Load-deflection relationships of the monotonically tested beams

The use of FRCM composites enhanced the flexural response of the corrosion-damaged beams

with a notable increase in their load-carrying capacities. The strengthened beams showed post-

yielding stiffness that was higher than that of the unstrengthened beams. This finding was

concluded from the slope of the load-deflection curves of Beams MCS-2P-I, MCS-4P-I, MCS-4P-

II, and MCS-3C-II (Figure 7.10). The enhancement in the post-strengthening flexural response

was highly dependent on the type and amount of the applied FRCM composites as follows.

Beam MCS-2P-I, with 𝛽𝑓 = 2.8%, reached an ultimate capacity of 85.6 kN, which corresponds

to a strength gain of only 7% when compared to the ultimate capacity of the control beam. On the

other hand, Beam MCS-4P-I with 𝛽𝑓 = 5.54% (almost double that of Beam MCS-2P-I), reached

an ultimate capacity of 102.6 KN, which was 29% higher than that of the control specimen. These

results indicated that the contribution of FRCM composites increased with an increase in the

number of fabric layers. However, the increase in strength was not linearly proportional to the

increase in the stiffness factor, 𝛽𝑓. This finding was confirmed by comparing the results of Beams

MCS-4P-II and MCS-4P-I having similar 𝛽𝑓 of 5.54%. Both beams showed similar ultimate loads

despite the different schemes used.

The stiffness factor, 𝛽𝑓, should not be used solely to compare the strengthening effectiveness of

different FRCM composites without considering their cementitious matrix bond characteristics.

For instance, Beam MCS-3C-II with 𝛽𝑓 = 8.01% that is 144 % of that of Beam MCS-4P-II, reached

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an ultimate strength of 123.3 KN compared to 102.9 kN for Beam MCS-4P-II. This variation in

the ultimate strength of Beam MCS-4P-II from that of Beam MCS-3C-II (approximately 20%)

compared to the large difference between their stiffness factors (𝛽𝑓) was attributed to the inferior

properties of the cementitious matrix of the C-FRCM with respect to those of the PBO-FRCM

counterparts.

7.4.3 Fatigue Test Results

7.4.3.1 Fatigue Life

The fatigue test results of beams of Group F are shown in Table 7.2. The virgin Beam FUU

successfully survived 2 million cycles without failure. The test was then halted and the beam was

tested monotonically up to failure to examine its residual strength as will be detailed later.

Corrosion of steel bars reduced the fatigue life of the unstrengthened Beam FCU to 396,000 cycles,

that is only 20% of the fatigue life of Beam FUU.

The use of FRCM composites increased the fatigue life of the corrosion-damaged beams.

However, none of the FRCM-strengthened beams could restore the fatigue life of the virgin beam.

Beam FCS-2P-I had the shortest fatigue life among the FRCM-strengthened beams with 545,600

cycles survived. Beams FCS-4P-I and FCS-4P-II strengthened with the same amount of PBO-

FRCM (same 𝛽𝑓) but with different strengthening schemes survived 984,800 and 1,493,300 cycles,

receptively. On the other hand, Beam FCS-3C-II survived 834,500 cycles before failure.

Figure 7.11 shows the variation of the number of fatigue cycles survived and the stiffness factor,

𝛽𝑓 , for the tested beams. It was noted that increasing 𝛽𝑓 from 2.76 to 5.54 % for beams FCS-2P-I

and FCS-4P-I, respectively, increased the fatigue life by 80% (545,600 cycles for the former beam

versus 984,800 cycles for the later one). It is worth mentioning that both beams were strengthened

with PBO-FRCM in Scheme I. On the contrary, Beam FCS-3C-II had 𝛽𝑓 of 7.99, which was 146

that of Beam FCS-4P-II. However, the fatigue life of the former was 80% shorter than that of the

later (834,500 cycles for Beam FCS-3C-II versus 1,493,300 cycles for Beam FCS-4P-II). While

this finding can be attributed to the scatter that commonly encountered in fatigue tests, it matched

well the monotonic results of both beams and confirmed the conclusion previously obtained that

the stiffness factor, 𝛽𝑓 , should not be used solely to compare the strengthening effectiveness of

different FRCM composites without taking into account their matrix properties. It also confirmed

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the superior characteristics of the PBO-FRCM composite compared to those of C-FRCM

counterpart. Beam FCS-4P-II demonstrated exceptional bond characteristics between the fabric

and the surrounding matrix and also between the matrix and the concrete substrate. Good bond

ensured the transfer of stress to the FRCM composite thus reducing the stresses in the reinforcing

bars, which allowed the beam to restore about 75% of the fatigue life of the virgin beam and to

exceed that of the corrosion-damaged beam by 277%.

PBO-strengthened beams having similar values of 𝛽𝑓 but with different configuration of their

FRCM composite showed significant scatter between the number of cycles survived. This was

shown from the test results of beams FCS-4P-I and FCS-4P-II having 𝛽𝑓 of 5.54. The use of U-

shaped continuous layer of FRCM in Beam FCS-4P-II improved its fatigue life by nearly 52% of

that of Beam FCS-4P-I. The enhancement in fatigue life of Beam FCS-4P-II was ascribed to the

confinement effect of the U-shaped layer and its contribution in reducing the stress level in the

steel bars. Recall that both beams showed similar load-carrying capacities when tested under

monotonic loading, which suggested that the effectiveness of the U-shaped continuous layer of

Scheme II was more pronounced in fatigue rather than in monotonic tests.

Figure 7.11: Variation of fatigue life of the strengthened beams with the stiffness factor 𝛽𝑓

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7.4.3.2 Fatigue Behavior and Modes of Failure

The control Beam FUU was tested monotonically to failure after surviving 2 million cycles. The

beam failed due to steel yielding followed by concrete crushing. The beam exhibited a load-

deflection response similar to that of reference MUU specimen with residual deflections at the

onset of loading and reduced stiffness as a result of the cracks formed during the cyclic loading.

The residual yield and ultimate load capacities of Beam FUU were 73.4 and 79.3 KN, respectively,

compared to 75.1 and 79.7 kN for Beam MUU.

Figure 7.12a to Figure 7.12d present the hysteresis loops for the unstrengthened Beam FCU and

for the strengthened Beams FCS-4P-I, FCS-4P-II, and FCS-3C-II, respectively. Test observations

indicate that the fatigue response of all tested beams consisted of three stages. The first stage is

characterized by the initiation of vertical cracks accompanied by the increase in midspan

deflections. In this early stage of the fatigue life, flexural cracks tend to increase in height and

number before they stabilize.

The second stage describes most of the fatigue life of the beams. In this stage, the growth of

cracks is remarkable due to the reduction in modulus of elasticity of concrete as reported by several

authors [52]. As the number of cycles increase, a primary flexural crack (referred to as the fatigue

crack) appeared in the region of maximum moment. Experimentally, the fatigue crack steadily

propagate towards the compression zone until failure occurs at its location. This phenomenon was

common in beams strengthened with both schemes. However, the fatigue crack in beams

strengthened with Scheme II was smaller in width than that observed in beams strengthened with

Scheme I. Fatigue cracks for Beams FCS-4P-I, FCS-4P-II and FCS-3C-II are depicted in Figure

7.13. Small dispersed cracks were also observed in the FRCM cementitious matrix without any

evidence of fabric slippage or FRCM delamination. It was also notable that FRCM-strengthened

beams showed more gradual and steady degradation in their fatigue strength during this stage as

compared to that of the unstrengthened Beam FCU. This might be related to the potential of FRCM

composites to mitigate the crack propagation during cycling.

As per Figures 7.12a through 7.12d, the final stage of the fatigue life is characterized by a sudden

degradation in the beams’ stiffness and a considerable increase in deflection. Failure occurred

suddenly due to the abrupt rupture of one of the bottom steel bars. All of the strengthened beams

failed by fatigue rupture of the corroded steel bars as shown in Figure 7.14 for Beams FCU and

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FCS-4P-I. Rupture of the steel bars in the FRCM-strengthened beams led to the complete falling-

out of the FRCM composite due to the excessive deformation occurred. The visual inspection of

the extracted corroded bars after failure revealed that rupture of steel bars occurred at the location

of the corrosion pits where the greatest mass loss occurred (Figure 7.8c).

a)

b)

0

10

20

30

40

50

60

0 2 4 6 8 10

Load

(K

N)

Deflection (mm)

0 cycles (0%) 7200 cycle (2%)

100800 cycles (25%) 302400 cycles (75%)

396000 cycle (100%)

0

10

20

30

40

50

60

0 5 10 15

Lo

ad (

KN

)

Deflection (mm)

0 cycles (0%) 7200 cycle (0.7%)201600 cycles (20%) 604800 cycles (60%)806400 cycle (80%) 984800 cycles (100%)

147

c)

d)

Figure 7.12: Load-deflection Hysteresis for a) beam FCU; b) beam FCS-4P-I; c) beam FCS-4P-

II; and d) beam FCS-3C-II

0

10

20

30

40

50

60

0 5 10 15

Lo

ad (

KN

)

Deflection (mm)

0 cycles (0%) 7200 cycles (0.5%)201600 cycles (14%) 806400 cycles (54%)1,209600 cycles (81%) 1,493000 cycles (100%)

0

10

20

30

40

50

60

0 5 10 15

Load

(K

N)

Deflection (mm)

0 cycles (0%) 7200 cycle (0.9%)201600 cycles (24%) 403200 cycles (48%)720000 cycles (86%) 834500 cycle (100%)

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Figure 7.13: Fatigue cracks at midspan of the strengthened beams (Side view)

Figure 7.14: Fatigue rupture of steel bars in a) beam FCU and b) beam FCS-4P-I

7.4.3.3 Stiffness Degradation

Figure 7.15 shows the variation in stiffness of the tested beams with the number of fatigue cycles

survived. The flexural stiffness of the tested beams is calculated as the ratio of the maximum

fatigue load to the corresponding midspan deflection. The initial stiffness for all tested beams were

measured after 7,200 cycles. Test results show that the virgin Beam FUU lost 5% of its initial

stiffness after 200,000 cycles. The beam suffered from gradual and steady stiffness degradation as

the number of cycles increased. After 2 million cycles, the beam lost about 10% of its initial

stiffness. Beam FCU exhibited a similar trend of stiffness degradation before failing suddenly after

396,000 cycles with flexural stiffness about 84% of its initial stiffness.

Fatigue cracks (0.25 mm) after 0.98M

cycles

Fatigue cracks (0.12mm) after 1.49M

cycles

Fatigue cracks (0.15mm) after 0.83M

cycles

a) Specimen FCS-4P-I b) Specimen FCS-4P-II c) Specimen FCS-3C-II

Bar rupture Bar rupture

149

Figure 7.15: Effect of fatigue cycles on the stiffness of the tested beams

FRCM-strengthened beams showed initial stiffness that was comparatively higher than that of

the unstrengthened ones. As the number of cycles increased, different rates of stiffness degradation

were encountered. The average rate of stiffness degradation in the beams strengthened with PBO-

FRCM was 3.33% per 100,000 cycles for the first 200,000 cycles of fatigue loading. This rate then

decreased to 0.64% per 100,000 cycles until a significant loss of stiffness was observed following

the rupture of one of the steel bars.

Test results indicate that the type of the FRCM composite rather than the number of fabric layers

and the strengthening scheme have a notable effect on the rate of stiffness degradation. This was

depicted in the comparable plots shown in Figure 7.15 for Beams FCS-2P-I, FCS-4P-I, and FCS-

4P-II, having different number of fabric layers and/or different schemes. On the other side, beam

FCS-3C-II demonstrated a higher rate of stiffness degradation than that recorded for the PBO-

FRCM strengthened beams. This finding might be attributed to the superior characteristics of the

PBO-FRCM composite as previously demonstrated in the monotonic test results. It also suggests

that the type of FRCM composite played a major role in defining the stiffness degradation during

fatigue.

150

7.4.3.4 Strain Response

Figure 7.16 shows the variation of strains in concrete and in the outer fabric with the number of

cycles survived. All of the tested beams showed a gradual increase in the measured concrete strains

with an increase in the number of cycles. This is attributed to the concrete softening that results in

an increase in the beam deflections as illustrated in Figure 7.12. The maximum compressive

concrete strains ranged between -763 to -961 μɛ for Specimen FCU while the maximum strains

recorded for the strengthened beams ranged between -603 to -984 μɛ. Similarly, the outer fabric

strains increased during the fatigue life due to the propagation of cracks until failure occurred.

Figure 7.16: Effect of fatigue cycles on the concrete and fabric strains

The maximum fabric strains measured for Beam FCS-2P-I increased from 2707 μɛ at 7200 cycle

to 2827 μɛ at 500,000 cycles and reached 4051 μɛ at failure. Beam FCS-4P-I showed maximum

fabric strains that grown from 2350 μɛ at 7200 cycle to 3875 μɛ at failure. Beam FCS-4P-II

exhibited relatively lower initial fabric strains than that of Beam FCS-4P-I (1972 versus 2350 μɛ)

while the strains reached 2791 μɛ after 1,493,000 cycles. The maximum carbon fabric strains in

Specimen FCS-3C-II were lower than those recorded for the PBO-FRCM and ranged between

1480 to 1850 μɛ. It is important to note that the measured stains in concrete and FRCM did not

exceed their ultimate strain values. These results indicate that the residual fatigue strength of the

steel bars after corrosion control the fatigue strength of the strengthened RC beams rather than the

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fatigue strength of concrete or FRCM composites. This assumption is valid as long as the

externally bonded composites are able to reduce the stress level in the corroded steel bars to that

level of stress prior to corrosion.

7.5 Conclusions

In this study, the monotonic and fatigue performances of corrosion-damaged RC beams

strengthened with PBO- and C-FRCM composites were experimentally investigated. The

following conclusions can be drawn from the test results:

Monotonic Tests

• An average mass loss of 19% in the steel reinforcing bars due to corrosion reduced the

yield and ultimate strengths of the beams by 14 and 7%, respectively, without changing its

ductile mode of failure.

• Strengthening corrosion-damaged beams with PBO- and C-FRCM composites enhanced

the flexural response and restored/exceeded the load capacities of the virgin beam. The

ultimate loads of the PBO-strengthened beams ranged between 107 and 129% of that of

the control beam and was 155% for the carbon-strengthened beam.

• The modes of failure of the FRCM-strengthened beams varied with the type and scheme

of the FRCM composite used. PBO-strengthened beams failed by fabric delamination in

scheme I and by fabric slippage and debonding in scheme II. Carbon-strengthened beams

failed by the premature cracking of the matrix followed by the fabric slippage.

• The contribution of FRCM composites, expressed by the stiffness factor 𝛽𝑓, increased with

an increase in the number of fabric layers. However, the associated gain in capacity was

not linearly proportional to the increase in stiffness factor, 𝛽𝑓.

• The stiffness factor 𝛽𝑓 should not be used solely to compare the strengthening effectiveness

of different FRCM composites without considering the mechanical properties of its

constituents. Corrosion-damaged beams strengthened with PBO- and C-FRCM composites

having same 𝛽𝑓 showed different load-carrying capacities and different modes of failure.

152

Fatigue Tests

• Corrosion of steel bars dramatically decreased the fatigue life of the RC beam. While the

virgin beam sustained 2 million fatigue cycles, the corrosion-damaged beam failed after

396,000 cycles. Therefore, corrosion shortened fatigue life to 20% of that of the virgin

beam.

• Rupture of steel bars at the locations of the highest corrosion pits intensity was the

governing mode of failure for all of the unstrengthened and strengthened beams.

• Strengthening with FRCM composites increased the fatigue life of the corrosion-damaged

beams by 38 to 377% of that of the unstrengthened beam depending on the type and

configuration of the FRCM composite used. However, strengthening did not restore the

fatigue life of the virgin beam.

• PBO-FRCM composite was more effective than the C-FRCM counterpart in restoring the

fatigue life of the corrosion-damaged beams. Beam FCS-3C-II (with 𝛽𝑓 = 7.99) survived

80% less cycles than those survived by beam FCS-4P-II (with 𝛽𝑓 =5.48).

• The effect of FRCM configuration was more pronounced in fatigue than in monotonic tests.

Having the same number of PBO fabric layers, beam FCS-4P-II strengthened with scheme

II (U-shaped continuous layer) exhibited longer fatigue life than beam FCS-4P-I

strengthened with scheme I (U-shaped end-anchors). Both beams showed similar

monotonic load-carrying capacities.

• The type of the FRCM composite had a notable effect on the rate of stiffness degradation

of the strengthened beams in fatigue rather than the number of fabric layers and the

strengthening scheme applied. The beam strengthened with C-FRCM demonstrated a

higher rate of stiffness degradation with cycling than the beam strengthened with PBO-

FRCM.

The test results presented in this paper demonstrate that strengthening with FRCM composites is

an effective technique to restore the fatigue life of corrosion-damaged concrete beams. However,

the outcome of this study should not be extrapolated to other FRCM composites before being

153

validated with experimental tests. Future work should investigate the fatigue performance of other

FRCM composites and for various levels of corrosion of the reinforcing steel bars.

154

8. Chapter 8

Finite Element Modeling and Experimental Results of Corroded Concrete Beams Strengthened with Externally-

bonded Composites


Journal of Engineering Structures. Date of submission: November 4, 2017


Résumé

Cet article présente les résultats de la modélisation en éléments finis 3D (ÉF) de poutres en béton

armé endommagées par la corrosion et renforcées en flexion avec des composites collés à

l'extérieur. Les modèles ont été validés par rapport aux résultats d'essais expérimentaux effectués

sur dix poutres non renforcées et renforcées. Les paramètres étudiés comprenaient les niveaux de

corrosion (10% et 20% de perte de masse de l'armature en acier), le type de composite (matrice

cimentaire renforcée de fibre, MCRF et polymère renforcé de fibre, PRF) et le nombre de couches

de fibres (une, deux et quatre couches). Les résultats prédits ont montré un bon accord avec ceux

des tests expérimentaux. Les modèles ÉF ont pu capturer le comportement non linéaire des poutres.

Les modèles de lien-glissement aux interfaces des MCRF/matrice et PRF/béton et le nombre de

couches de fibres ont eu l'impact le plus significatif sur la réponse prédite des poutres renforcées

alors que le niveau de corrosion, modélisé comme une réduction de la section des armatures, a

montré un léger effet sur leur performance. Une étude paramétrique a examiné l'effet de la

variation de la résistance du béton à la compression sur la performance des poutres renforcées.

L'abaissement de la résistance à la compression du béton a diminué les capacités ultimes des

poutres renforcées par les composites, quel que soit le système de renforcement utilisé (MCRF ou

PRF). Le mode de défaillance des poutres renforcées par MCRF était indépendant de la résistance

à la compression du béton et était gouverné par le glissement des fibres de la matrice. Cependant,

pour les poutres renforcées en PRF, l'abaissement de la résistance à la compression du béton a

modifié leur mode de rupture, de la rupture de PRF au décollement à l'interface PRF/béton.

Mots-clés des auteurs : Corrosion; Matrice cimentaire renforcée de fibres; Polymères renforcés

de fibres; Analyse des éléments finis; Flexion; Béton armé; Réparation; Renforcement.

155

8.1 Abstract

This paper presents results of 3D finite element (FE) modeling of corrosion-damaged reinforced

concrete (RC) beams strengthened in flexure with externally-bonded composites. The models were

validated against the results of experimental tests conducted on ten unstrengthened and

strengthened beams. The investigated parameters included the corrosion levels (10% and 20%

mass loss of steel reinforcement), the type of composite (fabric-reinforced cementitious matrix

(FRCM) and fiber-reinforced polymers (FRP)), and the number of fabric layers (one, two, and four

layers). The predicted results showed good agreement with those of the experimental tests. The FE

models were able to capture the non-linear behavior of the beams. The interfacial bond stress-slip

models at the FRCM/matrix and FRP/concrete interfaces and the number of fabric layers had the

most significant impact on the predicted response of the strengthened beams whereas the corrosion

level, modeled as a reduction in the reinforcement cross-section, showed a slight effect on their

performance. The validified models were used in parametric studies to investigate the effect of

varying the compressive strength of concrete substrate and the thickness of concrete cover on the

flexural performance of the strengthened beams. Lowering the compressive strength of the

concrete substrate or increasing the thickness of the concrete cover decreased the load-carrying

capacities of the strengthened beams regardless of the strengthening system used (FRCM or FRP).

Failure of FRCM-strengthened beams was independent of the compressive strength of concrete or

the thickness of concrete cover and was governed by fabric slippage within the matrix, unlike in

the case of FRP-strengthened beams.

Authors’ keywords: Corrosion; Fabric-reinforced cementitious matrix; Fiber reinforced

polymers; Finite element analysis; Flexure; Reinforced concrete; Repair; Strengthening.


Strengthening of corrosion-damaged reinforced concrete (RC) structures has become one of the

most imperative activities in the construction industry. Corrosion impairs the structural integrity

and the serviceability of the structures and can lead to unexpected collapses [44,45,53]. In the past

decades, previous research has documented the effectiveness of fiber-reinforced polymers (FRP)

as reliable strengthening materials for concrete structures [10,11,111]. More recently, cement

mortars reinforced with fabrics made of carbon, glass, or polyparaphenylene benzobisoxazole

156

(PBO), known as fabric-reinforced cementitious matrix (FRCM) or textile-reinforced mortar

(TRM), have been introduced as promising, sustainable, and durable alternatives to FRP

composites. FRCM systems own all the merits of FRPs in terms of corrosion resistivity, light

weight, and ease of installation but with the use of inorganic cementitious matrices as bonding

materials in lieu of the epoxy-bonding agents for FRPs to overcome some of the drawbacks

associated with epoxies [16,59,103,112,113].

Finite element packages have been used to examine the structural performance of strengthened

concrete beams and slabs [114–116] with the majority of the literature body focusing on modelling

the behavior of beams strengthened using FRP sheets or plates. The bond behavior between

concrete and FRP have been developed through rigorous numerical modelling of concrete-FRP

joints in direct shear tests, a basic application that provides insight into FRP-concrete interfacial

behavior [117,118]. The motivation for such existing numerical work was the fact that, despite the

large amount of experimental data available on FRP strengthening of concrete structures, a full

understanding of the various load-deformation behaviors and debonding phenomenon was still

lacking.

On the other hand, recent experimental studies involving FRCM systems have focused on

strengthening undamaged RC members. Test results have demonstrated their efficiency in

restoring the capacities and serviceability of the deficient members [55,81,119]. Elsanadedy et al.

(2013) [83] developed a FE model to predict the flexural behavior of six beams strengthened with

basalt FRCM (B-FRCM) and carbon-FRP systems (C-FRP) using LS-DYNA software. Bond

between FRCM and concrete was modeled through the tiebreak surface-to-surface contact

definition of LS-DYNA to account for both normal and shear forces at the interface. A parametric

study was also conducted by altering the type of mortar and the number of B-FRCM layers used.

A bond-stiffness coefficient, defined as the ratio of the B-FRCM stiffness to its tensile bond

strength, was introduced and recommended not to be less than 225 to avoid premature debonding

failure.

Al-Salloum et al. (2012) [92] evaluated both experimentally and numerically the efficacy of B-

FRCM system for shear strengthening of deficient concrete beams using long woven, knitted or

even unwoven fiber rovings in two orthogonal directions. A prefect bond between FRCM

composites and concrete was assumed in the FE model. The number of textile layers and the

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orientation of the textile material were altered through a parametric study using FE models. The

models were able to accurately capture the shear strengths of the strengthened beams.

It was observed that most of the previous studies have reported that debonding of FRCM

composites governs the failure of the flexural strengthened RC members [80,84,85], which

highlights the significance of bond of FRCM to the concrete substrate. Debonding of FRCM

usually occurs at the fabric/matrix interface rather than the matrix/concrete interface, unlike what

is typically reported for FRPs, in which debonding occurs at the epoxy/concrete interface

[72,76,117]. D’Ambrisi et al. [120] developed bond-slip models to describe bond between FRCM

fabric and the surrounding matrix. These models were then calibrated with data obtained from

experimental tests in which RC elements were strengthened with PBO-FRCM layers. This model

will be described later in the FE model developed in this study. Ombers [75] proposed another

bond-slip relationship for PBO-FRCM composites bonded to concrete. However, Ombres’ model

was reported to be more conservative than that introduced by D’Ambrisi et al. [120].

The conducted literature shows that little attention has been devoted to investigating the

feasibility of using FRCM composites in strengthening corrosion-damaged RC members.

Corroded structures are characterized by the deterioration of concrete and the loss of structural

integrity as a result of expansive corrosion products. Using cement-based mortars in repair might

be a challenge from the practical and technical points of view. To the authors’ knowledge, only

two studies [86,121] have documented the potential of using FRCM composites to restore the

ultimate capacities and the serviceability of corrosion-damaged beams. However, many

parameters that might affect the performance of FRCM-strengthening have not been fully

documented and yet need to be thoroughly investigated.

The aim of the current study was twofold, namely:

a) expand the understanding of flexural behavior of corrosion-damaged RC beams strengthened

with FRCM composites by i) including parameters that were not included in previous studies

[86,121] and ii) comparing their performance with that of FRP-strengthened beams.

b) validate newly-developed FE models that utilize the bond stress-slip model proposed by

D’Ambrisi et al. [120] to describe the bond behavior between PBO-FRCM and concrete.

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Details about the test specimens, the accelerated corrosion process, the test setup, and the test

results of some of the beams have been reported in [121] and are reiterated here for convenience.

The FE models were then extended in parametric studies to examine the effect of varying the

compressive strengths of the concrete substrates and the thickness of concrete cover on the flexural

enhancement of the strengthened beams.

8.3 Experimental Investigation

8.3.1 Test Matrix

The test matrix of the experimental program is given in Table 8.1. The beams were subdivided

into two groups A and B and were subjected to accelerated corrosion process to obtain theoretical

mass losses of 10 and 20%, respectively, in the middle third of their tensile steel reinforcement.

Details about the accelerated corrosion process can be found in [121]. At the end of the corrosion

process, one beam in each group was not strengthened (beams CU-A and CU-B) and were used as

benchmarks while other beams were strengthened with the designated externally-bonded

composite system. In addition, two virgin beams (i.e., not corroded nor strengthened: beams UUa

and UUb) were used as controls. The test parameters included the level of corrosion damage (10

and 20%), the type of the externally-bonded composite system (FRCM and FRP), and the volume

fraction of the fabric used in the FRCM composite (1, 2, or 4 layers).

8.3.2 Test Specimen

The geometry and details of reinforcement of the test specimen are shown in Figure 8.1. The

beams were 2.8 meters long with a cross section of 150×250 mm and a clear span of 2.2 meters.

The bottom and top reinforcement consisted of two 15M (diameter 15 mm) and 8M (diameter 8

mm) deformed bars, respectively, placed at a clear cover of 25 mm. Steel stirrups of 10 mm

diameter were provided along the shear spans to avoid shear failure. After testing the beams, the

corroded steel bars were extracted from the beams and the actual mass loss of the corroded bars

were determined according to ASTM G1-03 provisions [26]. Table 8.1 lists the actual average

steel mass loss for each beam.

The cylinder compressive strength of concrete used was 41.8 MPa with standard deviation of 4.8

MPa. The splitting tensile strength of concrete was 3 MPa with standard deviation of 0.3 MPa.

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The yield strength of the longitudinal reinforcing steel bars of diameters 15 and 8 mm were 466

MPa (standard deviation of 4.2 MPa) and 573 MPa (standard deviation of 17.7 MPa) as determined

by the authors.


Beam ID* Avg. mass loss

(%)

As

(mm2)

Strengthening regime

Reference Type

No. of

layers

Af

(mm2)

𝐾𝑓

(KN)

Control beams: Uncorroded unstrengthened beams

UUa, UUb - 400 - - - - [121]

Group (A): Theoretical steel mass loss of 10%

CU-A 12.9 348.4 - - - - [121]

CS-A-1C 11.6 353.6 C-FRP 1 19.5 4485 -

CS-A-1P 13.7 345.2 P-FRCM 1 6.9 1421 -

CS-A-2P 13 348 P-FRCM 2 13.8 2842 [121]

CS-A-4P 12.6 349.6 P-FRCM 4 27.6 5684 [121]

Group (B): Theoretical steel mass loss of 20%

CU-B 18 328 - - - - [121]

CS-B-2P 19.6 321.6 P-FRCM 2 13.8 2842 [121]

CS-B-4P 19.4 322.4 P-FRCM 4 27.6 5684 [121]

* UU, CU, and CS refer to Uncorroded-Unstrengthened, Corroded-Unstrengthened, and Corroded-Strengthened beams,

respectively. A and B refer to beam groups A and B, respectively. 1, 2, and 4 in the beam’s label refer to the number of

composite layers. P and C refer to P-FRCM and C-FRP, respectively.

Figure 8.1: Geometry and details of steel, P-FRCM, and C-FRP reinforcement of the tested

beams (all dimensions in mm)

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8.3.3 Externally-bonded Composite Systems

Two types of externally-bonded composites were used to strengthen the beams namely, the P-

FRCM composite and the C-FRP composite. The P-FRCM composite consisted of a fabric made

of polyparaphenylene benzobisoxazole (PBO) fibers impregnated and bonded to the concrete

surface by a polymer-modified cementitious matrix having compressive and flexural strengths of

43.9 and 3 MPa (standard deviations of 0.4 and 0.3 MPa), respectively, as determined by the

authors. The PBO fabric consisted of an unbalanced net of spaced fiber bundles organized along

two orthogonal directions as shown in Figure 8.2a. The fabric openings were 5 and 15 mm wide

in the primary and the secondary directions, respectively. The bundles were 5 and 2.5 mm wide

with thicknesses of 0.046 and 0.011 mm in the main and secondary directions, respectively. The

tensile tests conducted by D’Antino et al. [78] on the same fabric indicated a tensile modulus of

206 GPa, a tensile strength of 3014 MPa, and an ultimate elongation of 14.5‰. The P-FRCM

composite had a cracked tensile modulus of 121 GPa, a tensile strength of 1550 MPa, and an

ultimate elongation of 14‰ as determined by Ebead et al. [97].

Figure 8.2: a) PBO fabric used in the FRCM composite system and b) Carbon sheets used in the

FRP composite system

On the other hand, the C-FRP composite consisted of flexible unidirectional carbon fiber sheet

(Figure 8.2b) impregnated and bonded to concrete with an epoxy resin having a tensile modulus

of 3.8 GPa and a tensile strength of 30 MPa according to the manufacturer’s data sheet. The dry

carbon fiber sheet had a tensile modulus of 230 GPa, a tensile strength of 3.22 GPa, an ultimate

elongation of 14.5‰, and a nominal thickness of 0.13 mm. The cured laminate had a tensile

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modulus of 65.4 GPa, a tensile strength of 0.894 GPa, an ultimate elongation of 13.3‰, and a

nominal thickness of 0.38 mm, as reported in the manufacturer’s data sheet.

To compare between the two strengthening systems, the equivalent axial stiffness, 𝐾𝑓, for each

composite system was determined based on their tensile modulus, 𝐸𝑓, and the cross-sectional area

of the fibers embedded within the composite, 𝐴𝑓, as given in Equation (8.1). Table 8.1 lists the

values of 𝐴𝑓 and 𝐾𝑓 for all of the strengthened beams.

𝐾𝑓= 𝐴𝑓𝐸𝑓 Eq. (8.1)

8.3.4 Strengthening Procedure and Configuration

The strengthening procedure is shown in Figure 8.1. At the end of the corrosion process, the

deteriorated concrete was first removed using a hydraulic hammer (Figure 8.3a). The corroded

steel bars were then brushed and a cementitious repair mortar (Sikacrete-08SCC) was used to fill

the damaged zone. The repair mortar had a compressive strength of 55.4 MPa (standard deviation

of 5 MPa) and a flexural strength of 3.4 MPa (standard deviation of 0.3 MPa). After curing, the

beam surface was sandblasted before applying the externally-bonded composite system (Figure

8.3b). The P-FRCM was installed using the hand lay-up method. The first layer of the cementitious

matrix was applied to the concrete substrate with a thickness of 4 to 5 mm (Figure 8.3c). Then, the

fabric was installed and coated with a second layer of matrix of similar thickness. The procedure

was then repeated until the specified number of layers was attained according to the test matrix in

Table 8.1. Similar procedure was followed while applying the C-FRP composite system (Figure

8.3d).

Each layer of the designated composite system was 150 mm wide (width equal to that of the

beam) and was applied to the soffit of the beam over a length of 2400 mm. The fibers were oriented

so that their primary directions were parallel to the longitudinal axis of the beam. The flexural plies

were anchored at each end using one U-shaped transverse strip of 300 mm width and 200 mm

height as shown in Figure 8.1.

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Figure 8.3: Strengthening procedure of corroded beams: a) removing the deteriorated concrete,

b) patch repairing and sandblasting, c) installation of P-FRCM composite, and d) installation of

C-FRP sheets


The beams were tested under four-point loading configuration as shown in Figure 8.1. The load

was applied in displacement control at a rate of 2 mm per minute using a MTS actuator. Deflections

were measured by means of three linear variable differential transducers (LVDTs) located at mid-

span and under the point loads. All beams were instrumented at mid-span with a 60 mm long

concrete strain gauge bonded to the top surface of the beams and 5 mm steel strain gauges bonded

to the tensile reinforcing bars. The strengthened specimens were instrumented with 5 mm strain

gauges installed directly on the outer fabric of the composite system at mid-span.

8.4 Numerical Simulation

Three-dimensional finite element (FE) models have been developed to simulate the nonlinear

flexural behavior of the tested beams using the software package ATENA 3D [122]. Due to

symmetry of loading and support conditions, only half of the beam was modeled and the

appropriate symmetry conditions were applied in order to reduce the computational time. The C-

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FRP composite system was modeled as a discrete reinforcement bonded directly to the beam soffit

while the P-FRCM composite was modeled using a more detailed approach that involved modeling

the fabric and the matrix layers. The mass losses in steel bars due to corrosion were represented

by reductions in the cross-section area of the bars, As, according to the actual mass losses given in

Table 8.1. The mechanical properties of the materials reported earlier were used as input data in

the 3D model. The materials constitutive laws, elements type, and boundary conditions adopted in

the models are presented in the following sections.

8.4.1 Constitutive Laws

8.4.1.1 Concrete and Cementitious Matrix

The 3D nonlinear cementitious material model of the FE package (CC3DNonLinCementitious2)

is used to simulate concrete and the cementitious matrix associated with the P-FRCM system. The

model consists of two parts representing the compressive (plastic) and the tensile (fracturing)

behaviors. The compressive model is based on the Menétrey-Willam failure surface [123] while

the tensile model is based on the classical orthotropic smeared crack formulation and crack band

model. The compressive model consists of an ascending branch that represents the hardening phase

as shown in Figure 8.4a and a descending branch that represents the softening phase shown in

Figure 8.4b. In the hardening phase, the material behaves linearly up to a compressive stress value

of 𝑓𝑐𝑜= 2𝑓𝑡, where fco is the compressive stress at the onset of the nonlinear compressive behavior

and ft is the tensile strength of the material. The nonlinear behavior is governed by Equation (8.2)

as follows:

𝜎𝑐 = 𝑓𝑐𝑜 + ( 𝑓𝑐′ - 𝑓𝑐𝑜 ) √1 − {

ɛ𝑐𝑝− ɛ𝑝

ɛ𝑐𝑝}

2

Eq. (8.2)

where 𝜎𝑐 = the compressive stress in the nonlinear hardening part, 𝑓𝑐′ = cylinder compressive

strength, ɛ𝑝 = plastic strain, and ɛ𝑐𝑝 = plastic strain corresponding to the compressive strength.

164

Figure 8.4: Constitutive laws of concrete and cementitious matrix: a) compressive hardening law,

b) compressive softening law, and c) tensile softening law

In the softening phase, the material is assumed to behave linearly as shown in Figure 8.4b. The

plastic strain, ɛ𝑝, is transformed into displacement, 𝑤𝑐, through the length scale parameter, 𝐿𝑐,

which corresponds to the projection of the element size in the direction of minimal principal stress

as shown in Figure 8.4b. The plastic displacement, 𝑤𝑑, shown in Figure 8.4b is used to define the

end of the softening phase in compression and is assumed equal to 0.5 mm [124].

The compressive model also includes a reduction in the compressive strength after cracking in

the direction parallel to the cracks in a way similar to that proposed by Vecchio and Collins [125].

The reduced strength after cracking is estimated using Equation (8.3) and Equation (8.4) as

follows:

𝑓𝑐𝑒𝑓𝑓

= 𝑟𝑐 𝑓𝑐′ Eq. (8.3)

𝑟𝑐 = 1

0.8 +170ɛ1 , 𝑟𝑐

𝑙𝑖𝑚 ≤ 𝑟𝑐 ≤ 1.0 Eq. (8.4)

where 𝑓𝑐𝑒𝑓𝑓

= the effective compressive strength in a direction parallel to the direction of cracks,

ɛ1= the strain in a direction normal to the direction of the crack, 𝑟𝑐 is a reduction factor due to

cracking, and 𝑟𝑐𝑙𝑖𝑚= maximal compressive strength reduction factor taken as 0.8.

The tension model is assumed linear elastic up to the material tensile strength, 𝑓𝑡, followed by an

exponential softening shown in Figure 8.4c. The slope of the linear ascending branch is taken equal

to the material elastic modulus. In the softening branch (Figure 8.4c), the fixed crack model

proposed in [126] is adopted. In this model, the crack direction is given by the principal stress

direction at the moment of crack initiation. During further loading, this direction is assumed fixed

165

and represented the material axis of orthotropy. The exponential function of the crack opening is

given in Equation (8.5) and Equation (8.6) as derived experimentally by Hordijk [127]:

𝜎𝑡

𝑓𝑡 = {1 + (𝑐1

𝑤𝑡

𝑤𝑡𝑐)

3

} 𝑒𝑥𝑝 (−𝑐2𝑤𝑡

𝑤𝑡𝑐) −

𝑤𝑡

𝑤𝑡𝑐(1 + 𝑐1

3) 𝑒𝑥𝑝(−𝑐2) Eq. (8.5)

𝑤𝑡𝑐 = 5.14 𝐺𝑓

𝑓𝑡 Eq. (8.6)

where 𝑤𝑡 = crack opening displacement, 𝑤𝑡𝑐= crack opening at the complete release of stress, 𝜎𝑡

= the normal stress in the crack (crack cohesion), 𝐺𝑓 = fracture energy of the material needed to

create a unit area of stress free crack, and the constants 𝑐1and 𝑐2 taken equal to 3 and 6.93,

respectively [127]. The crack opening, 𝑤𝑡, is derived from the material strain, ɛ𝑐𝑓, and the crack

band length, 𝐿𝑡, is assumed equal to the size of the element projected into the crack direction.

The shear strength of the cementitious material is calculated using Equation (8.6), which is based

on the modified compression field theory developed by Vecchio and Collins [125] as follows:

𝜏𝑒𝑓𝑓 = 0.18√𝑓𝑐

′

0.31+24𝑤

𝑎𝑔 + 16

Eq. (8.7)

where 𝜏𝑒𝑓𝑓 = effective shear strength of a cracked cementitious material, w = maximum crack

width at a given location, and 𝑎𝑔 = maximum aggregate size. The values of the parameters used in

the constitutive model of the cementitious materials are presented in Table 8.2.

Table 8.2: Characteristics of concrete and P-FRCM matrix used in FE models

Parameter Concrete P-FRCM

Cementitious

matrix C41.8 C30 C20

Cylinder compressive strength, 𝑓𝑐′, (MPa) 41.8 30 20 43.9

Cube compressive strength, 𝑓𝑐𝑢, (MPa) 49.2 35.3 23.5 51.6

Tensile strength, 𝑓𝑡, (MPa) 3.2 2.6 1.9 3.2

Elastic modulus, E, (GPa) 36.7 32.4 27.3 37.35

Poisson’s ratio, ʋ 0.2 0.2 0.2 0.2

Specific fracture energy, 𝐺𝑓, (MN/m) 8.05×10-5 6.45×10-5 4.93×10-5 8.32×10-5

Critical compressive displacement, 𝑤𝑑, (m) -5 ×10-4 -5×10-4 -5×10-4 -5×10-4

Minimum compressive strength reduction factor, 𝑟𝑐𝑙𝑖𝑚 0.80 0.80 0.80 0.80

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8.4.1.2 Steel and Fabrics

The stress-strain law of the steel reinforcing bars is assumed elastic perfectly plastic (i.e. bilinear)

Prefect bond is assumed between the steel bars and concrete and between the steel bars and the

repair material in all of the strengthened beams. The carbon and the PBO fabrics are assumed to

behave linearly elastic up to failure.

8.4.1.3 Interfacial Bond Stress-slip Models

Interfacial bond stress-slip models have been adopted at the interfaces of C-FRP/concrete and

PBO-fabric/matrix to simulate the experimental observations in which debonding of the

externally-bonded system occurred at the C-FRP/concrete interface in beam CS-A-1C while it took

place at the fabric/matrix interface in the beams strengthened with P-FRCM.

8.4.1.4 C-FRP/Concrete Interfacial Bond Stress-slip Model

The bilinear bond stress-slip model proposed by Lu et al. [117] has been adopted to model the

interfacial behavior between the C-FRP and concrete. The model is governed by Equation (8.8) to

Equation (8.12) and is shown in Figure 8.5.

𝜏𝑚𝑎𝑥 = 1.5𝛽𝑤𝑓𝑡 Eq. (8.8)

𝑆0 = 0.0195𝛽𝑤𝑓𝑡 Eq. (8.9)

𝑆𝑓 = 2𝐺𝑓𝑡 𝜏𝑚𝑎𝑥 ⁄ Eq. (8.10)

𝐺𝑓𝑡 = 0.308𝛽𝑤2 √𝑓𝑡 Eq. (8.11)

𝛽𝑤 = √2.25 – 𝑏𝑓 𝑏𝑐⁄

1.25 + 𝑏𝑓 𝑏𝑐⁄ Eq. (8.12)

where 𝜏𝑚𝑎𝑥 = maximum bond stress, 𝑆0 = slip at maximum bond stress, 𝑆𝑓 = slip at failure, 𝐺𝑓𝑡

= interfacial fracture energy of concrete, 𝛽𝑤 = width coefficient factor, 𝑏𝑓 = width of the C-FRP

sheet, 𝑏𝑐 = width of the concrete substrate, and 𝑓𝑡 = concrete tensile strength.

167

Figure 8.5: C-FRP/concrete interfacial bond stress-slip model according to Lu et al. [117]

adopted for various concrete mixes

8.4.1.5 PBO-fabric/matrix Interfacial Bond Stress-slip Model

The bond between the PBO-fabric and the associated cementitious matrix is described by

Equation (8.13) and is shown in Figure 8.6, as proposed by D’Ambrisi et al. [120]. Equation

(8.13) is a modification of the bond-slip relation proposed in [128] for FRP sheets, where 𝜏 (s) =

bond strength, s = the corresponding slip, 𝜏𝑖 = initial bond strength, 𝑆𝑓 = slip at failure. A, 𝛼, and

𝛽 are curve-fitting parameters and their values were evaluated as 0.92, 8.94, and 43.85,

respectively, for P-FRCM [120]. In their study, D’Ambrisi et al. [120] also proposed the values of

𝜏𝑖 = 0.15 MPa and 𝑆𝑓 = 1.18 mm for P-FRCM.

𝜏 (s) = [ 𝜏𝑖+ A ( 𝑒−𝛼𝑆 - 𝑒−𝛽𝑆 )] . (1 − 𝑆

𝑆𝑓) 0≤ 𝑆 ≤ 𝑆𝑓 Eq. (8.13)

0

0.5

1

1.5

2

2.5

3

3.5

0 0.05 0.1 0.15 0.2 0.25

𝜏(s)

(M

Pa)

S (mm)

C 41.8

C 30

C 20

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Figure 8.6: PBO-fabric /matrix interfacial bond stress-slip model according to D’Ambrisi et al.

[120]

8.4.1.6 Element Type

To model the concrete and the cementitious matrix, 3D solid brick elements with 8 nodes were

used. The size of the concrete and matrix elements were 30 and 10 mm, respectively, based on a

mesh sensitivity study conducted by the authors. Based on test observations, perfect bond was

assumed between the concrete and the cementitious matrix. The steel, carbon, and PBO were

modeled using 3D truss elements. The typical mesh discretization of the beams along with the

details of the steel and the externally bonded reinforcements are shown in Figure 8.7.

8.4.1.7 Loading, Boundary Conditions, and Monitoring Points

Loading procedure and support steel plates similar to the ones actually used in the tests were

adopted in the FE model. The beams in FE models were loaded by means of prescribed

displacements located at the midpoint of the top surface of the steel loading plate, at an increment

of 0.1 mm. The support plate was restrained from movement in the vertical and transverse direction

(z and y directions, respectively) by means of a line support placed at the middle line of the bottom

surface of the plate. Since only half of the beam was modeled, the vertical surface along the axis

of symmetry was restrained from the horizontal movement (i.e., in the x direction).

The load was monitored at the midpoint of the top surface of the load plate where the prescribed

displacements were applied. Another point at the beam mid-span on the bottom surface of the

concrete or the composite system was used to record the beam deflections. The strains in the

169

bottom steel bars and the outer fiber layer of the composite system were monitored by means of

monitoring points at the beam mid-span. Strain monitoring point was also used to measure the

concrete compressive strains at mid-span.

Figure 8.7: a) Meshing of P-FRCM strengthened beams, b) reinforcement layout for beams

strengthened with 4 layers of P-FRCM, and c) reinforcement layout for beams strengthened with

C-FRP sheet

(a) 1

2 Front view 3

4 Bottom View 5

(b) 6

Front view 7

Bottom View 8

(c) 9

Steel reinforcement

Loading plate

Support plate

Concrete mesh

Externally-bonded P-FRCM

P-FRCM-end anchor

170

8.5 Results and Discussion

In this section, the experimental and numerical results are presented and discussed. The finite

element models were verified by comparing the results of the numerical analysis with the

experimental results. More details about the experimental results can be found in [121].

8.5.1 Crack Pattern at Failure

The predicted cracking pattern at failure was compared to that obtained from the tests in Figures

8.8a, 8.8b, and 8.8c for beams UUa, CS-A-1C, and CS-A-4P, respectively. Both the experimental

and numerical investigations indicated that the control beams (UU) and the corroded

unstrengthened beams (CU-A and CU-B) failed due to yielding of the steel bars followed by

concrete crushing. This mode of failure can be depicted in Figure 8.8a for the control beam UUa.

This finding indicated that the corrosion damage of the bottom steel bars with mass loss up to 18%

(group B) did not change the ductile mode of failure of the RC beams [121]. Beam CS-A-1C failed

due to a longitudinal crack that developed within the concrete layer close to the C-FRP

laminate/concrete interface, which was followed by a sudden rupture of the C-FRP laminate. A

thin layer of concrete was attached to the laminate at failure as shown in Figure 8.8b. The same

mode of failure was captured also by the FE model. A good correlation between the predicted and

experimental crack patterns was achieved as shown in Figure 8.8b.

On the other hand, all beams strengthened with two or four layers of P-FRCM in groups A and

B failed due to FRCM delamination at the fabric/matrix interface adjacent to the concrete substrate

[121] as shown in Figure 8.8c. This was similar to the mode of failure obtained from the numerical

simulation of beam CS-A-4P, which occurred due to the slippage of fabric within the cementitious

matrix. It is important to note that this mode of failure was not similar to that observed in beam

CS-A-1P, which was strengthened with one layer of P-FRCM composite, in which no bond

degradation between the P-FRCM composite and the concrete substrate was observed during the

test. Beam CS-A-1P failed by steel yielding followed by concrete crushing prior to the

delamination of the P-FRCM composite.

For all of the strengthened with two and four layers of P-FRCM, good correlation between the

predicted and experimental crack patterns as shown in Figure 8.8c. These finding indicate that the

inclusion of the interfacial bond stress-slip models between the C-FRP and concrete and between

171

the PBO-fabric and the matrix enabled the FE models to detect the appropriate failure modes of

the strengthened beams.

a)

b)

172

c)

Figure 8.8: Numerical and experimental crack patterns at failure for a) beam UUa, b) beam CS-

A-1C, and c) beam CS-A-4P

8.5.2 Load-deflection Response

The experimental load-deflection responses are compared with those obtained numerically in

Figure 8.9 and Figure 8.10 for beams of groups A and B, respectively. Both experimental and

numerical load-deflection responses consisted of three stages with two turning points indicating

concrete cracking and yielding of steel bars. As can be seen in Figure 8.9 and Figure 8.10, there is

a good agreement between the numerical and experimental load-deflection plots, which verifies

the accuracy of the FE models in capturing the nonlinear responses of the strengthened beams. The

corrosion damage of the bottom steel reinforcement, which was represented in the numerical

analysis by a reduction in the cross-section area of the steel bars, was offset after strengthening

and most of the beams restored their initial capacities as will be detailed later. It is important to

note that the use of externally-bonded composites (i.e. C-FRP or P-FRCM composites) had no

notable impact on the stiffness of the strengthened beam prior to steel yielding. However, all of

the strengthened beams exhibited various degrees of enhancement in their post-yielding stiffness,

which was mainly dependent on the type and the amount of fiber used as has been confirmed

numerically.

The main experimental and numerical results of the load-deflection responses of the tested beams

are summarized in Table 8.3 and Table 8.4. For the unstrengthened beams, the results given in

Table 8.3 indicated that the yield and ultimate loads measured experimentally for beams CU-A

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and CU-B were 7 and 5% and 14 and 7% lower than the corresponding values of the control

uncorroded beams (UU). The numerical results of the same beams were in agreement with those

measured experimentally. The ratio of the yield loads predicted numerically to those measured

experimentally (𝑃𝑦𝐹𝐸 𝑃𝑦

𝐸𝑋𝑃⁄ ) was 0.99 and 1.02 for beams CU-A and CU-B, respectively. The

corresponding ratios of the ultimate loads (𝑃𝑢𝐹𝐸 𝑃𝑢

𝐸𝑋𝑃⁄ ) for the same beams were 0.94 and 0.93,

respectively.

a) Beam UU, CU-A, and CS-A-1C

a) Beam UU, CU-A, and CS-A-1C

Figure 8.9: Numerical and experimental load-deflection responses for beams of group A

174

Table 8.3: Predicted and experimental strength and fiber strain results

Beam Yield load (kN) Ultimate load (kN) Ultimate fiber strain (‰) Reference

𝑃𝑦𝐸𝑋𝑃 𝑃𝑦

𝐹𝐸 𝑃𝑦𝐹𝐸 𝑃𝑦

𝐸𝑋𝑃⁄ 𝑃𝑢𝐸𝑋𝑃 𝑃𝑢

𝐹𝐸 𝑃𝑢𝐹𝐸 𝑃𝑢

𝐸𝑋𝑃⁄ ɛ𝑓𝑢𝐸𝑋𝑃 ɛ𝑓𝑢

𝐹𝐸


UUa, UUb* 75.1 78.22 1.04 79.7 81.7 1.03 - - [121]


CU-A 69.5 68.7 0.99

76.1 71.8 0.94 - - [121]

CS-A-1C 77.9 77.6 1 96.5 103.4 1.07 13.8 14.48 -

CS-A-1P 71.1 74.3 1.04 82.9 79.5 0.96 14.92 10.96 -

CS-A-2P 79.5 75.3 0.95 86.4 89.6 1.04 8.49 12.47 [121]

CS-A-4P 83.3 82.3 0.99 99.6 106 1.06 9.44 13.11 [121]


CU-B 64.5 65.7 1.02

74.2 68.7 0.93 - - [121]

CS-B-2P 71.8 71 0.99 85.6 85.1 0.99 8.18 13.28 [121]

CS-B-4P 79.6 78.6 0.99 102.6 103.8 1.01 10.66 13.32 [121]

* Average values are reported

175

Table 8.4: Predicted and experimental deflections and ductility indices

Beam Deflection at yield load Deflection at ultimate load

Ductility index

Reference 𝛿𝑦

𝐸𝑋𝑃 (mm) 𝛿𝑦𝐹𝐸 (mm) 𝛿𝑦

𝐹𝐸 𝛿𝑦𝐸𝑋𝑃⁄ 𝛿𝑢

𝐸𝑋𝑃 (mm) 𝛿𝑢𝐹𝐸 (mm) 𝛿𝑢

𝐹𝐸 𝛿𝑢𝐸𝑋𝑃⁄ 𝜇 𝐸𝑋𝑃 𝜇 𝐹𝐸 𝜇 𝐹𝐸/𝜇 𝐸𝑋𝑃


UUa, UUb* 11.7 10.7 0.91 32.9 31.3 0.95 2.81 2.93 1.04 [121]


CU-A 10.9 10.5 0.96

35.4 29.5 0.83

3.25 2.81 0.86 [121]

CS-A-1C 12.3 11 0.89 30.4 38.7 1.27 2.47 3.52 1.42 -

CS-A-1P 11.7 10.7 0.91 32.9 26.1 0.79 2.81 2.44 0.87 -

CS-A-2P 11.8 10.5 0.89 33 33 1 2.8 3.14 1.12 [121]

CS-A-4P 12.8 11.2 0.88 31.5 33.7 1.07 2.46 3.01 1.22 [121]


CU-B 9 10.1 1.12

28.4 34.9 1.23

3.16 3.46 1.1 [121]

CS-B-2P 11.3 10.6 0.94 33.7 31.1 0.92 2.98 2.93 0.98 [121]

CS-B-4P 12.5 11.51. 0.92 37.4 35.6 0.95 2.99 3.1 1 [121]

* Average values are reported

176

Figure 8.10: Numerical and experimental load-deflection responses for beams of group B

For the strengthened beams, both the experimental and numerical results listed in Table 8.3

revealed that flexural strengthening of the corrosion-damaged RC beams with C-FRP or P-FRCM

enhanced their load capacities. As can be seen in Table 8.3, the ratio of the predicted to measured

yield loads of the tested beams (𝑃𝑦𝐹𝐸 𝑃𝑦

𝐸𝑋𝑃⁄ ) ranged between 0.95 and 1.04 with an average ratio =

1.0, a standard deviation of 2.8%, and a coefficient of variation of 2.8%. The ratio of the predicted

to measured ultimate loads (𝑃𝑢𝐹𝐸 𝑃𝑢

𝐸𝑋𝑃⁄ ) of all beams ranged between 0.93 and 1.07 with an average

ratio = 1, a standard deviation of 5%, and a coefficient of variation of 5%.

Table 8.4 compares the deflections of the tested beams at both yield and ultimate loads. The

deflections predicted numerically were in a good agreement with those measured experimentally.

The predicted deflections at ultimate loads for all beams strengthened with P-FRCM were within

7% error band except in the case of beam CS-A-1P for which the numerical model underestimated

the deflection at ultimate by 21%. This was attributed to the discrepancy in the experimental and

predicted modes of failure observed for the beam CS-A-1P. While test observations indicated

prefect bond between the P-FRCM composite and the concrete substrate up to failure, the FE

model predicted a premature failure due to the fabric slippage within the cementitious matrix. This

finding can be depicted in Figure 8.9b from the sudden drop in the predicted load-deflection curve.

On the contrary, the FE model overestimated the defection at ultimate load for the beam CS-A-1C

strengthened with C-FRP laminate by approximately 27% as can be observed in Figure 8.9a. This

177

was related to the actual premature rupture of the C-FRP laminate at ultimate, which was

confirmed by the measured fiber strain at ultimate as will be explained later.

8.5.3 Load-carrying Capacity

The decline in the ultimate load, 𝑃𝑢, in the corroded unstrengthened beams due to corrosion of

the steel bars and the capacity gain experienced by the strengthened beams are shown in Figure

8.11. As previously noted, the measured ultimate loads for the corroded unstrengthened beams

CU-A and CU-B were 5% and 7% lower than that of the control virgin beam (UU), respectively.

The corresponding values predicted numerically were 12.1 and 15.9%, respectively, which means

that the numerical model tended to overemphasize the effect of corrosion damage on the load-

carrying capacities of the corroded beams. This could be explained by the fact that the FE models

didn’t accurately represent the mass loss that physically occurred during the corrosion process.

While the loss of the bars’ lugs due to corrosion contributed significantly to the mass loss measured

in the lab, it didn’t affect significantly the cross-section of the steel bars and hence didn’t affect

the load-carrying capacities of the corroded beams. However, the mass loss adopted in the FE

models significantly reduced the cross-section area of the bars and therefore contributed directly

to the decline in the load-carrying capacities of the beams.

Figure 8.11: Gain and decline in % in the ultimate loads, 𝑃𝑢, with respect to that of the control

beam (UU)

178

For the strengthened beams, both the experimental and numerical results indicated the

effectiveness of the externally-bonded composite systems (C-FRP and P-FRCM) in restoring the

yield and ultimate loads of the corroded beams as given in Table 8.3. The gain in strengths was

highly dependent on the amount and type of fibers used rather than the level of corrosion damage

in the tensile steel bars [121]. Beam CS-A-1P that was strengthened with one layer of P-FRCM

exhibited load-carrying capacity of 82.9 kN, which was 104% of that of the control beam. The

failure load predicted by the numerical model was 79.5 kN, representing a decline of -4%

compared to the experimental value. Increasing the number of P-FRCM layers further increased

the load-carrying capacities of the strengthened beams. For instance, beams CS-A-2P and CS-A-

4P failed at 86.4 and 99.6 kN, respectively, due to the delamination of the P-FRCM composite.

The predicted ultimate loads for the same beams were 89.6 and 106 kN, respectively, representing

differences of +4 and +6% compared to the experimental values. Both the experimental and

numerical results revealed that the gain in the load-carrying capacities of the strengthened beams

was approximately proportional to the number of P-FRCM layers used (i.e., the cross-section area

of the composite, 𝐴𝑓). The predicted gain in the ultimate strengths for the beams strengthened with

one, two, and four layers of P-FRCM were 7.7 kN (for CS-A-1P), 17.8 kN (for CS-A-2P), and

34.2 (for CS-A-4P), respectively. This finding was also confirmed by the similar ultimate fiber

strains reported for these beams.

Beams CS-B-2P and CS-B-4P encountered load-carrying capacities of 85.6 and 102.6 kN,

respectively, compared to 85.1 and 103.8 kN predicted numerically, representing differences of -

0.6 and +1.1% compared to the experimental values. The measured and predicted ultimate loads

for the beams of group B were approximately similar to their counterparts in group A (CS-A-2P

and CS-A-4P) that experienced lower level of corrosion. This finding indicated that approximately

doubling the level of corrosion had an insignificant impact on the load-carrying capacities of the

strengthened beams. More details about the effect of corrosion level on the performance of the

FRCM-strengthened beams are reported in Elghazy et al. [121].

The beam strengthened with C-FRP laminate failed at 96.5 kN, which was 21.1% higher than

that of the control beam. The finite element model predicted a load-carrying capacity of 103.4 kN,

representing a difference of +7% compared to the experimental value. This slight discrepancy

between the numerical and experimental capacities is generally acceptable. It is important to note

179

that the equivalent axial stiffness, 𝐾𝑓, of one layer of C-FRP, which strengthened beam CS-A-1C,

was about 80% of that of four layers of P-FRCM that’s strengthened beam CS-A-4P. It was

therefore expected to experience a higher load-carrying capacity in the latter beam than its former

counterpart. However, experimental and numerical data confirmed that beam CS-A-1C (with the

lower axial stiffness) showed an ultimate capacity of 97.5% of that of beam CS-A-4P (with the

higher axial stiffness). This finding indicated that the strengthening effectiveness of the P-FRCM

system was slightly lower than that of the C-FRP system.


The experimental and predicted relationships between the applied loads and the strains measured

in the outer fiber and concrete are shown in Figure 8.12 for all of the tested beams. Tensile steel

strain responses for beams CS-A-1C and CS-A-4P are also shown in Figure 8.13. Representative

numerical strain profiles of the internal and external reinforcement at ultimate are shown in Figure

14 for beams CS-A-1C and CS-A-4P.

Similar to the load-deflection responses, the numerical and experimental load-strain curves

consisted of three segments with two turning points that indicated the concrete cracking and the

yielding of the tensile steel bars. It can be observed that the numerical models reasonably predicted

the strain responses in the fiber, the concrete, and the tensile steel bars, which further validated the

accuracy of the FE models.

Table 3 lists the strain values in the outer fiber of the composite systems at ultimate as determined

experimentally, ɛ𝑓𝑢𝐸𝑋𝑃, and predicted numerically, ɛ𝑓𝑢

𝐹𝐸. Both the experimental and numerical

results indicated that beam CS-A-1C failed due to fiber rupture, with strain values ɛ𝑓𝑢𝐸𝑋𝑃 and ɛ𝑓𝑢

𝐹𝐸

reaching 13.8 and 14.48‰, respectively. This finding suggested that the measured strains in the

C-FRP laminate at failure was lower than that reported in the manufacturer’s data sheet and utilized

in the FE model. This finding explained why the FE model slightly overestimated the ultimate load

of beam CS-A-1C by 7%. The predicted ultimate strain profile of the internal and external

reinforcement shown in Figure 14a for beam CS-A-1C indicated that the maximum C-FRP strain

took place within the maximum moment zone and decreased gradually outside this zone, which

confirmed the rupture of fibers approximately at mid-span as observed in the tests.

180

Figure 8.12: Fiber and concrete strain response

Experimental results

Experimental results

Numerical results

Numerical results

a) Group A

a) Group B

181

Figure 8.13: Tensile steel strain response for beams CS-A-1C and CS-A-4P

a) Beam CS-A-1C

b) Beam CS-A-4P

Figure 8.14: Strain profile of the internal and external reinforcement at ultimate

182

Beam CS-A-1P experimentally recorded the maximum tensile strains in the outer fabric of the

P-FRCM layer (14.92‰) of all the beams. This was attributed to the mode of failure of beam CS-

A-1P, in which no fabric delamination was observed until failure occurred. On the other hand, the

predicted fiber strain at failure for beam CS-A-1P was 10.96‰ before failure occurred numerically

due to fiber slippage within the matrix. This discrepancy between the experimental and numerical

results can be attributed to the accuracy of the adopted interfacial bond-slip model between the

PBO-strands and the matrix in the FE model.

The measured fiber strains, ɛ𝑓𝑢𝐸𝑋𝑃, for beams strengthened with two and four layers of P-FRCM

ranged between 8.18 and 10.66‰ while the predicted fiber strains, ɛ𝑓𝑢𝐹𝐸, ranged between 12.47 and

13.32‰. Both the experimental and numerical results of such beams showed that the strains in the

PBO-fiber at peak loads were lower than the PBO-fiber rupture strain, which confirmed the modes

of failure of these beams due to the fiber slippage within the matrix. This finding can also be

depicted from the strain profile of the internal and external reinforcements predicted at ultimate

load as shown in Figure 8.14b for beam CS-A-4P. The slippage of the PBO-strands occurred within

the maximum moment zone (Figure 8.14b), which was consistent with the test observations.

8.5.5 Ductility

The deflection-based ductility index, μ, has been employed in the present study to evaluate the

effect of the externally-bonded composites on the ductility of the strengthened beams. The ductility

index is the ratio of the mid-span deflection at ultimate, 𝛿𝑢, to the mid-span deflection at yielding,

𝛿𝑦. A higher ductility index indicates a higher ability for the beam to provide sufficient

deformation, and hence ample warning, prior to failure. The experimental and numerical ductility

indices of the tested beams are given in Table 8.4. The predicted ductility indices were in

reasonable agreement with the experimentally obtained ones with a tendency of the FE model to

overestimate the ductility indices of the strengthened beams.

The ductility indices obtained experimentally ranged between 0.87 and 1.07 times that of the

control beam, which was adequate to guarantee satisfactory ductility. All beams strengthened with

P-FRCM showed ductility indices approximately similar to that of the control beam except beam

CS-A-4P that showed a ductility index 13% less than that of the control beam. Beam CS-A-1C

183

exhibited a ductility index equal to that of the beam CS-A-4P. These results indicated that the

externally-bonded composites slightly reduced the ductility of the strengthened beams.

Numerically, the ductility indices ranged between 0.83 and 1.2 times that of the predicted index

of the control beam. As previously noted, the FE model tended to overestimate the indices of the

strengthened beams except in case of beam CS-A-1P, in which the predicted index was 13% less

than the experimental index. This was explained by the different mode of failure predicted by the

FE model.

8.6 Parametric Studies

The verified FE models were used to investigate the effect of varying the compressive strength

of concrete substrate and the thickness of concrete cover on the flexural behavior of FRCM- and

FRP-strengthened beams. In each case, three FE models corresponding to three beams of the

experimentally-tested beams namely, beams CS-A-1C, CS-A-2P, and CS-A-4P, were selected as

baseline models for comparison. Then, the parameter of interest was varied to study its influence

on the flexural behavior of each beam. The results of the parametric study are summarised in Table

8.5 and Table 8.6 and discussed in the following sections.

8.6.1 Effect of Concrete Compressive Strength (𝒇𝒄′ )

This investigation aims to assess the potential of using the externally-bonded strengthening

technique to restore the service and ultimate capacities of RC beams having low compressive

concrete strengths. The tested beams had concrete compressive strength of 41.8 MPa (C41.8). Two

other concrete models C30 and C20 with compressive strengths 30 and 20 MPa, respectively, were

used in the numerical analysis to predict the flexural performance of beams CS-A-1C, CS-A-2P,

and CS-A-4P. The parameters used in the concrete models C41.8, C30, and C20 are given in Table

8.2.

The results of varying the concrete strength are summarized in Table 8.5 and the predicted load-

deflection responses are shown in Figure 8.15. The load-deflection response was similar in all

models with a noticeable lower stiffness response in beams with low concrete strengths. The use

of concrete with low compressive strengths decreased the load-carrying capacities of the

strengthened beams and decreased their mid-span deflections at ultimate. For example, the use of

184

C30 and C20 to model beam CS-A-4P reduced its ultimate load by 5.9 and 12.1% of that of the

baseline model (i.e. beam with concrete C41.8). Similar trend was observed in beams CS-A-1C

and CS-A-2P as can be depicted from Table 8.5.

Table 8.5: Summary of the parametric study results (effect of fc’)

FE Model Concrete 𝑃𝑢 (kN) Change in 𝑃𝑢

(%)** 𝛿𝑢 (mm) ɛ𝑓𝑢 (%) Mode of failure

CS-A-2P

C41.8* 89.6 0 32.97 12.47 Fiber slippage

C30 83.1 -7.2 28.63 9.535 Fiber slippage

C20 79.4 -11.3 26.41 11.06 Fiber slippage

CS-A-4P

C41.8* 106 0 33.69 13.11 Fiber slippage

C30 99.8 -5.9 31.37 12.02 Fiber slippage

C20 93.2 -12.1 30.68 11.57 Fiber slippage

CS-A-1C

C41.8* 103.4 0 38.69 14.48 laminate rupture

C30 97.7 -5.5 38.46 13.21 laminate debonding

C20 90 -12.9 35.03 11.59 laminate debonding

* Baseline FE model

** Relative to that of the baseline FE model

Slippage of the PBO fabric within the matrix governed the mode of failure in all of the FRCM-

strengthened beams regardless of their concrete compressive strengths. This suggested that the

interfacial bond stresses encountered at the fabric/matrix interface were higher than those

encountered at the concrete/matrix interface, which initiated slippage at the fabric level rather than

the separation between the matrix and the concrete substrate.

For beam CS-A-1C strengthened with CFRP laminates, it was remarkable to observe that the use

of low concrete compressive strength changed its mode of failure from laminate rupture (as

observed in the experimental test with concrete C41.8) to separation at the laminate/concrete

interface. This was attributed to the fact that the bond strength at the C-FRP/concrete interface was

significantly affected by the mechanical properties of the concrete substrate as shown in Figure

8.5. At failure, the recorded strains in the outer fabric in almost all of the models with lower

compressive strengths were lower than those recorded in the baseline models (Table 8.5).

185

Figure 8.15: Effect of the concrete compressive strength on the load-deflection response for

beams a) CS-A-1C, b) CS-A-2P, and c) CS-A-4P

186

8.6.2 Effect of Concrete Cover

Different thicknesses of the clear concrete cover were modeled to investigate their influence on

the strengthening effectiveness of FRCM and FRP systems. The clear concrete cover of the tested

beams was 25 mm. Two other concrete covers of 10 and 50 mm were modeled to predict the

flexural performance of beams CS-A-1C, CS-A-2P, and CS-A-4P.

The results of this study are summarized in Table 8.6 and the predicted load-deflection responses

are shown in Figure 8.16. All modeled beams showed similar load-deflection responses. However,

increasing the concrete cover from 25 to 50 mm significantly decreased the flexural stiffness of

the strengthened beams compared to that of the baseline model with concrete cover 25 mm (Figure

8.16). In addition, both yield and ultimate capacities of the strengthened beams were affected by

varying the concrete cover. This was attributed to the change in the effective depth of the steel

reinforcing bars as can be depicted from Figure 8.16. For example, the use of 50 mm concrete

cover in beams CS-A-4P and CS-A-1C decreased their ultimate capacities by 9.6 and 17.4 % of

those of the baseline models, respectively, whereas the use of a concrete cover of 10 mm increased

the ultimate capacities by 5.4 and 4.7% of those of the baseline models, respectively.

Table 8.6: Summary of the parametric study results (effect of concrete cover)

FE Model Concrete cover

(mm) 𝑃𝑢 (kN)

Change in 𝑃𝑢

(%)** 𝛿𝑢 (mm) ɛ𝑓𝑢 (‰) Mode of failure

CS-A-2P

25 89.6 0 32.97 12.47 Fiber slippage

10 90.2 +0.7 27.63 9.084 Fiber slippage

50 79.4 -11.3 34.51 12.85 Fiber slippage

CS-A-4P

25 106 0 33.69 13.11 Fiber slippage

10 111.8 +5.4 34.95 12.31 Fiber slippage

50 95.8 -9.6 33.55 12.09 Fiber slippage

CS-A-1C

25 103.4 0 38.69 14.43 laminate rupture

10 108.2 +4.7 37.23 14.47 laminate rupture

50 85.4 -17.4 27.43 10.08 laminate debonding

* Baseline FE model

** Relative to that of the baseline FE model

For the beams strengthened with P-FRCM system, varying the concrete cover did not affect their

mode of failure, which was characterised by the slippage of the PBO fabric within the matrix. This

finding was consistent with the measured fabric strains at failure shown in Table 8.6. On the other

hand, increasing the concrete cover in beam CS-A-1C (beam strengthened with C-FRP laminates)

187

to 50 mm changed its mode of failure from laminate rupture (in case of using 25 and 10 mm

concrete covers) to premature debonding at the laminate/concrete interface. This premature failure

significantly reduced the strengthening effectiveness of the FRP system. The results of this

parametric study indicated that the strengthening effectiveness of P-FRCM system was higher than

that of C-FRP system in the presence of thick concrete cover. This finding can be attributed to the

fact that failure of P-FRCM system is mainly dependent on the bond characteristics at the

fabric/matrix interface rather than the FRP/concrete bond strength in the case of FRP system.

b)

a)

188

Figure 8.16: Effect of the concrete cover on the load-deflection response for beams a) CS-A-1C,

b) CS-A-2P, and c) CS-A-4P

8.7 Conclusions

This paper discussed the numerical simulation of corrosion-damaged RC beams strengthened

with C-FRP and P-FRCM composites under flexural loading using ATENA. The load carrying

capacities, load-deflection responses, and load-strains responses were evaluated and compared

with the experimental results to validate the accuracy of the model. The validated models were

used in parametric studies to investigate the effect of varying the concrete compressive strength

and the thickness of concrete cover on the flexural behavior of the strengthened beams. The

following conclusions can be drawn from this study:

• The experimental and numerical studies indicated that the virgin and the corroded

unstrengthened beams failed due to yielding of steel bars. Corrosion up to 18% mass loss in

the steel bars showed no effect on the ductile mode of failure of the beams. However, the

numerical models tended to overestimate the effect of corrosion damage on the load-carrying

capacities of the corroded beams.

• Corrosion had an insignificant effect on the mode of failure of the C-FRP and P-FRCM-

strengthened beams. All beams strengthened with two and four layers of P-FRCM failed due

189

to FRCM delamination at the fabric/matrix interface adjacent to the concrete substrate

regardless of the corrosion level.

• Beam CS-A-1P strengthened with one layer of P-FRCM composite failed after steel yielding

prior to the delamination of the P-FRCM composite while beam CS-A-1C strengthened with

one layer of C-FRP laminate failed due to the sudden rupture of the laminate.

• Both the experimental and numerical results approved the ability of C-FRP and P-FRCM

composites to restore/increase the load-carrying capacities of the corrosion-damaged beams.

However, C-FRP composites were slightly more efficient than P-FRCM with similar axial

stiffness in restoring the beams strength.

• FE models accurately predicted the mode of failures of all of the strengthened beams except

that of beam CS-A-1P. While test observations indicated prefect bond between the P-FRCM

composite and the concrete substrate up to failure, the numerical model predicted a premature

mode of failure due to fabric slippage within the cementitious matrix.

• The ratios of the predicted to experimental yield loads of the tested beams (𝑃𝑦𝐹𝐸 𝑃𝑦

𝐸𝑋𝑃⁄ ) ranged

between 0.95 and 1.04 whereas the ratios of the predicted to measured ultimate loads

(𝑃𝑢𝐹𝐸 𝑃𝑢

𝐸𝑋𝑃⁄ ) ranged between 0.93 and 1.07. These results indicated the accuracy of the FE

models to predict the flexural performance of the tested beams.

• The predicted deflections at ultimate of all beams strengthened with P-FRCM were within 7%

error band except in the case of beam CS-A-1P in which the FE model underestimated its

deflection at ultimate by 21%. On the contrary, the FE model overestimated the defection at

ultimate for the C-FRP strengthened beam by 27%.

• The numerical results indicated the feasibility of using the bond slip models of D’Ambrisi et

al. [120] and Lu et al. [117] to simulate the behavior of P-FRCM and C-FRP-strengthened

beams at the composite/concrete interface.

• The results of the parametric study indicated that lowering the concrete compressive strength

or increasing the concrete cover decreased the load-carrying capacities of the strengthened

beams regardless of the strengthening system used (C-FRP or P-FRCM).

190

• The parametric study suggested that failure of FRCM-strengthened beams was independent

of the compressive strength of the concrete substrate or the thickness of the concrete cover.

Failure of such beams was governed by fabric slippage within the matrix. For beams

strengthened with C-FRP, lowering the compressive strength of the concrete substrate or

increasing the concrete cover changed their mode of failure from laminate rupture to fiber

slippage within the laminate.

Finally, the results of this study showed a promising potential of using FRCM composites in

restoring the load-carrying capacities and the serviceability of corrosion-damaged beams.

However, the experimental and numerical work presented in this paper is limited to the composites

used in the research reported in this paper and should not be extended to other types. The long-

term performance and durability of FRCM-strengthened beams are currently under investigation.

More research is needed to develop refined bond-slip models for FRCM composites, which

incorporate various parameters such as the number of fabric layers, the fabric type, and the fabric

orientation. Accurate correlation between the actual mass loss in the steel bars due to corrosion

and the reduction of reinforcement cross-section in FE models is also needed to better represent

the effect of corrosion level on the flexural performance of the strengthened members.

191

Chapter 9

Conclusions and Recommendations

8.8 Summary

Corrosion of steel reinforcement is the main cause of RC structure deterioration. Hence, durable

and costless repair technologies are vastly needed. FRCM composites have been introduced

recently as a new strengthening and repair technology for RC structures. The current study aimed

at investigating the monotonic and fatigue flexural behaviors of corrosion-damaged RC beams

strengthened with FRCM composites. Moreover, the long-term performance of the FRCM-

strengthened beams were investigated by exposing the strengthened beams to further corrosive

environment. The study includes both experimental and numerical investigations. In addition, the

analytical predictions provided by the current designs have been verified against the experimental

results.

The conducted work herein expands on the current knowledge of FRCM applications and

confirms the potential of FRCM composites to optimize the structural performance of corrosion-

damaged RC structures. The outcomes of this work are presented in five journal articles. A

summary of the findings and conclusions are presented in the following section.

8.9 Conclusions

8.9.1 Effect of Corrosion on RC Beams

• Corrosion of steel reinforcement initiated longitudinal cracks parallel to the corroded steel

bars. Wider corrosion cracks indicated higher steel mass loss due to corrosion. The

maximum widths of the observed corrosion cracks were 1.5, 2.8, and 3.5 mm for average

steel mass losses of 12.5, 19, and 22%, respectively. All of the corroded beams failed to

meet the provisions of ACI-318 design code for crack width criteria.

• Corrosion of steel reinforcing bars in the moment zone with an average mass loss up to

22.7% had a marginal impact on the flexural response of the tested beams. The maximum

decrease in the yield and ultimate strengths of the corrosion-damaged beams were 15 and

9%, respectively, of those of the virgin beams.

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• Corrosion of steel reinforcement reduced the yield and ultimate strengths of the corrosion-

damaged beams as compared to those of the virgin (non-corroded) beams. The yield and

ultimate strengths of the damaged beams decreased at average rates of 0.66 and 0.40%,

respectively, per 1% of steel mass loss. This finding was attributed to the fact that the

corroded bars lost their lugs, which increased the measured mass loss without affecting

their effective cross-sectional areas.

• Corrosion of steel reinforcement severely affected the fatigue life of RC beams. Corrosion

damage of 19.8% mass loss reduced the fatigue life of the corroded beam by 80% of that

of the virgin (non-corroded) beam. Moreover, the corroded beam failed abruptly due to

the sudden rupture of the corroded steel bars.

8.9.2 Short-term Performance of Corroded Beams Strengthened with

FRCM

• The use of FRCM composites improved the flexural response of the corrosion-damaged

beams. FRCM significantly enhanced the post-yield stiffness of the corroded beams

compared to that of the corroded unstrengthened beams. The enhancement in the flexural

response was highly dependent on the FRCM strengthening scheme, its type, and the

number of FRCM layers used rather than the level of corrosion damage.

• The use of PBO-FRCM and C-FRCM increased the yield and ultimate strengths of the

corroded beams damaged at various levels of corrosion (up to 22.7% steel mass loss).

After 22.5% steel mass loss, the ultimate strengths of beams strengthened with PBO- and

C-FRCM were 54 and 51% higher than those of the corroded unstrengthened beam and

about 39 and 37% higher than those of the virgin (uncorroded unstrengthened) beam,

respectively. While, the yield strength of the beams strengthened with PBO-and C-FRCM

were 26 and 17 % higher than those of the corroded unstrengthened beam and about 0.8

and 0.1% higher than those of the virgin beam, respectively

• All of the corroded-strengthened beams failed due to the loss of the strengthening action

of the FRCM layers regardless of their corrosion damage level. The only exception was

for the beam strengthened with one layer of PBO-FRCM in which no distress of the

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FRCM layer was observed until failure occurred due to concrete crushing. Four distinct

failure mechanisms were observed in the strengthened beams:

a) FRCM delamination: Delamination occurred between the fabric and the first layer of

the matrix adjacent to the concrete substrate. This mode of failure was observed in all

beams strengthened with PBO-FRCM in Scheme I (end-anchored layers).

b) Fabric slippage: This mode of failure was observed in all beams strengthened with

PBO-FRCM of Scheme II (U-shaped wrapping). Slippage occurred between the PBO

fabric and the cementitious matrix accompanied by slight delamination at the

fabric/matrix interface.

c) Matrix cracking and fabric separation from the matrix: This mode of failure was

reported in all beams strengthened with C-FRCM. It can be described by a progressive

cracking in the cementitious matrix associated with the sudden debonding between the

carbon fabric and the matrix. This mode of failure was more brittle than that observed

in the PBO-strengthened beams, which can be attributed to the superior characteristics

of the cementitious matrix of the PBO-FRCM compared to those of the C-FRCM

counterparts.

d) C-FRP laminate rupture: This mode of failure was reported for the beam strengthened

with C-FRP sheets. A longitudinal crack initiated at the laminate/concrete interface

followed by the sudden rupture of the laminate.

• Increasing the number of FRCM layers increased the yield and ultimate loads of the

strengthened beams. The increase in strength was approximately proportional to the

added number of layers. For instance, the use of one, two, and four layers of PBO-FRCM

increased the ultimate strengths of the strengthened beams by 9, 14, and 30% of that of

the corroded beam with12.5 % steel mass loss due to corrosion, respectively.

• Anchoring the FRCM flexural layers with a continuous U-shaped layer (Scheme II) was

more effective than the use of end anchors (Scheme I) in increasing the yield and ultimate

loads of the strengthened beams. This observation was attributed to the effect of the

continuous U-shaped layer in mitigating and delaying the delamination of the FRCM and

consequently increasing their strengthening potential. The use of two layers of PBO-

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FRCM in Scheme I enhanced its yield and ultimate loads by 6 and 8%, respectively, while

the use of the same number of layers but in Scheme II improved its yield and ultimate

strengths by 14 and 28%, respectively.

• PBO-FRCM composites were more efficient than C-FRCM in improving the flexural

performance of the corroded beams. The beam strengthened with four layers of PBO-

FRCM in Scheme II showed an ultimate strength 15% higher than that of beam

strengthened with two layers of C-FRCM with the same Scheme and corrosion-damage

despite the fact that two layers of C-FRCM had similar axial stiffness of the four layers

of PBO-FRCM. Moreover, the PBO-FRCM strengthened beams showed more ductile

failure than their counterparts strengthened with C-FRCM. These observations were

related to the superior bond characteristics of the matrix of the PBO-FRCM system as

compared to those of C-FRCM.

• C-FRP and PBO-FRCM composites restored/increased the load-carrying capacities of the

corrosion-damaged beams. However, C-FRP composites were slightly more efficient

than the PBO-FRCM with similar axial stiffness in restoring the beam strengths.

• Although the equivalent stiffness of the FRCM composites, Kf, and the stiffness factor,

𝛽𝑓, are believed to be good indicators of the strengthening effectiveness of the FRCM

systems, they should not be solely used to compare the strength gain in beams without

considering the bond characteristics at the matrix/fabric interface and the anchoring

Scheme used.

• Beams strengthened with FRCM composites showed ductility indices and energy

absorption indices that ranged between 86 and 118% and between 111 and 153%,

respectively, of those of the virgin beams.

8.9.3 Long-term Performance of Corroded Beams Strengthened with

FRCM

• The use of FRCM systems reduced the corrosion rate in the steel bars with no evidence

on the effect of the FRCM strengthening scheme on such rate. Exposing the FRCM-

strengthened beams to post-strengthening corrosion resulted in 23% reduction in the steel

195

mass loss. This was attributed to the reduction in the amount of water and air diffused to

the tensile reinforcing bars due to the presence of the FRCM layers, which also explained

the reduction in the corrosion rate in the long-term beams.

• For beams exposed to the post-strengthening corrosive environment, longitudinal cracks

parallel to the corroded reinforcement on one or both lateral surfaces of the beams were

observed. However, the use of continuous U-shaped layers (Scheme II) reduced the width

and the length of the corrosion cracks compared to the use of end anchors (Scheme I).

• Exposing the beams strengthened in Scheme II to post-strengthening corrosive

environment had no impact on their mode of failure regardless of the type of fabric used.

However, corrosion cracks resulted in premature delamination of FRCM for beams

strengthened in Scheme I, which significantly affected the strengthening action of the

FRCM system. This can be attributed to the wrapping effect of the FRCM layer in Scheme

II that offset the effect of corrosion cracks in weakening the concrete cover and prevented

the premature delamination of FRCM flexural layers.

• Exposing the beams to corrosion after FRCM strengthening had a marginal impact on

their load-deflection response and strength. The long-term beams strengthened in Scheme

II demonstrated higher ductility and energy absorption indices than those of their short-

term counterparts. Short-term beam strengthened in Scheme I showed higher ductility

and energy absorption indices than those of their long-term counterparts due to the

premature FRCM delamination caused by corrosion cracks.

8.9.4 Validation of ACI-549.4R-13 Design Equations

• Strain values recorded on the FRCM layers indicated that the assumption of prefect bond

suggested by the ACI 549.4R-13 while limiting the maximum design strain in the FRCM

system to 0.012 mm/mm is a simplification that appears justifiable and easy to implement

by engineers.

• The theoretical formulations of ACI 549.4R-13 reasonably predicted the ultimate

capacities of the corrosion-damaged RC beams strengthened with FRCM composites in

Scheme I (end-anchored bottom layers) but underestimated the capacities of those

196

strengthened in Scheme II (with a continuously-wrapped layer). The ratios of the

experimental to the theoretical ultimate capacities (𝑃𝑢

𝑃𝑢𝑡ℎ⁄ ) was in the range of 1.03 to1.18

and 1.1 to 1.23 for the beams strengthened in Scheme I and Scheme II, respectively. This

finding was attributed to the fact that ACI 549.4R-13 provisions don’t consider the effect

of the U-shaped FRCM layer present in Scheme II in confining the bottom fabric layers

and in delaying their delamination from the matrix.

• The obtained results suggested increasing the ultimate capacities predicted by ACI

549.4R-13 for beams strengthened in Scheme II by 10% to consider the effect of the

continuous anchoring.

• ACI 549.4R-13 provisions conservatively predict the ultimate capacities of the FRCM-

strengthened beams exposed to post-strengthening corrosive environment. The ratios of

the experimental to the theoretical ultimate capacities (𝑃𝑢

𝑃𝑢𝑡ℎ⁄ ) in the range of 1.21 to 1.3

for the beams strengthened in Scheme II, while it was 1.02 for the specimen strengthened

in Scheme I.

8.9.5 Fatigue Performance of Corroded Beams Strengthened with

FRCM

• Strengthening corrosion-damaged beams with FRCM composites increased their fatigue

life by 38 to 377% of that of the corroded-unstrengthened beams without restoring the

fatigue life of the virgin beams. The enhancement in fatigue life was dependent on the

FRCM type, amount, and strengthening Scheme.

• Rupture of steel bars at the locations of corrosion pits was the governing mode of failure

in all of the unstrengthened and strengthened beams tested under fatigue. However, the

presence of FRCM composites mitigated this brittle failure in the strengthened beams.

• PBO-FRCM composite was more effective than the C-FRCM counterpart in restoring the

fatigue life of the corrosion-damaged beams. Beam (FCS-3C-II) strengthened with three

layers of C-FRCM (with 𝛽𝑓 = 8) survived 80% less cycles than those survived by the

beam (FCS-4P-II) strengthened with four layers of PBO-FRCM (with 𝛽𝑓 =5.48).

197

• The effect of FRCM configuration was more pronounced in fatigue than in monotonic

tests. With the same number of PBO fabric layers, the beams strengthened in Scheme II

exhibited fatigue life 150% more of that of its counterpart beam strengthened in Scheme

I.

• The FRCM material had a notable effect on the rate of stiffness degradation of the

strengthened beams tested in fatigue rather than the number of fabric layers and the

strengthening Scheme applied. The beam strengthened with C-FRCM demonstrated a

higher rate of stiffness degradation with cycling than that strengthened with PBO-FRCM.

8.9.6 Numerical simulation

• The comparison between the experimental and numerical results revealed that the

developed finite element (FE) models were able to predict the nonlinear flexural response

of the corroded beams strengthened with C-FRP and PBO-FRCM from initial loading to

ultimate with good accuracy. In addition, the numerical model was able to detect the

failure modes of the strengthened beams.

• The ratios of the predicted to experimental yield loads of the tested beams (𝑃𝑦𝐹𝐸 𝑃𝑦

𝐸𝑋𝑃⁄ )

ranged between 0.95 and 1.04 whereas the ratios of the predicted to experimental ultimate

loads (𝑃𝑢𝐹𝐸 𝑃𝑢

𝐸𝑋𝑃⁄ ) ranged between 0.93 and 1.07. These results indicated the accuracy of

the FE models to predict the flexural performance of the tested beams.

• Ultimate deflections predicted for beams strengthened with PBO-FRCM were within 7%

error band except in the case of beam strengthened with one layer of PBO-FRCM in

which the FE model underestimated its ultimate deflection by 21%. On the contrary, the

FE model overestimated the ultimate deflection for the C-FRP strengthened beam by

27%.

• The numerical results indicated the feasibility of using the bond slip models of D’Ambrisi

et al. [120] and Lu et al. [117] to simulate the behavior of PBO-FRCM and C-FRP-

strengthened beams at the composite/concrete interface.

198

• The results of the parametric study indicated that the use of concrete with low

compressive strengths decreased the load-carrying capacities of the strengthened beams

regardless of the strengthening system used (C-FRP or PBO-FRCM).

• The parametric study suggested that failure of FRCM-strengthened beams was

independent of the compressive strength of the concrete substrate. Failure of such beams

was governed by fabric slippage within the matrix. For beams strengthened with C-FRP,

lowering the compressive strength of the concrete substrate changed their mode of failure

from FRP rupture to debonding at concrete/laminate interface.

8.10 Recommendation for Future Work

The findings and conclusions of the current study approved the feasibility of FRCM composites

to strengthen corrosion-damaged RC beams. However, some topics still require further

investigations. Recommendations for future studies are:

• Studying the structural performance of RC beams damaged at higher corrosion levels

(more than 23% mass loss) and strengthened with FRCM.

• Examining the post-strengthening performance of FRCM-strengthened beams under

harsh environmental conditions including freeze/thaw and high temperatures.

• Investigating the ability of various FRCM systems to reduce the corrosion rate in the steel

reinforcement when applied at various levels of corrosion damage and quantifying the

change in corrosion rates.

• Investigating the possibility of using FRCM composites in strengthening continuous RC

beams and quantifying their ability to redistribute moments between the sections.

8.11 Impact of Current Research

The results and outcomes of this research work can be used as a fundamental step to include

design provisions for the flexural strengthening using FRCM composites in the Canadian codes

(S6 and S806). It can also be used to update the current design guidelines ACI 549.4R-13.

199

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10. Bibliography

During this research work at Laval University, the candidate has authored the following

publications:

Journal Articles:

1. Elghazy, M., El Refai, A., Ebead, U., and Nanni, A. “Corrosion-Damaged Reinforced

Concrete Beams Repaired with Fabric-Reinforced Cementitious Matrix (FRCM).” Journal of

Composites for Construction, ASCE. Under review. In review. Submitted in the revised form:

February 2018.

2. Elghazy, M., El Refai, A., Ebead, U., and Nanni, A. (2017). “Effect of Corrosion Damage on

the Flexural Performance of RC Beams Strengthened with FRCM Composites.” Journal of

Composite Structures. Date of acceptance: https://doi.org/10.1016/j.compstruct.2017.08.069

3. Elghazy, M., El Refai, A., Ebead, U., and Nanni, A. “Post-repair Flexural Performance of

Corroded Beams Rehabilitated with Fabric-Reinforced Cementitious Matrix (FRCM) under

Corrosive Environment.” Journal of Construction and Building Materials. Date of acceptance:

23 January 2018. https://doi.org/10.1016/j.conbuildmat.2018.01.128

4. Elghazy, M., El Refai, A., Ebead, U., and Nanni, A. “Fatigue and Monotonic Behavior of

Corrosion-damaged Reinforced Concrete Beams Strengthened with FRCM Composites.”

Journal of Composites for Construction, ASCE. In review. Submitted in the revised form:

February 2018.

5. Elghazy, M., El Refai, A., Ebead, U., and Nanni, A. “Finite Element Modeling and

Experimental Results of Corroded Concrete Beams Strengthened with Externally-Bonded

Composites.” Journal of Engineering Structures. In review. Date of submission: November

2017.


213

Refereed Conference Papers:

1. Elghazy, M., El Refai, A., Ebead, U. and Nanni, A. “Post-Repair Flexural Performance of

Corrosion-Damaged Beams Repaired with Fabric-Reinforced Cementitious Matrix.” The 9th

International Conference on Fibre-Reinforced Polymer (FRP) Composites in Civil

Engineering - CICE 2018, Paris, France, July 2018.

2. Elghazy, M., El Refai, A., Ebead, U. and Nanni, A. “Experimental Results and Modeling of

Corroded Concrete Beams Strengthened with FRCM.” The 10th International Conference on

Short and Medium Span bridges, Quebec City, Canada, August 2018.

3. Elghazy, M., El Refai, A., Ebead, U. and Nanni, A. “Corrosion-Damaged Beams Repaired

with Fabric-Reinforced Cementitious Matrix under Fatigue Load.” The 10th International

Conference on Short and Medium Span bridges, Quebec City, Canada, August 2018.

4. Elghazy, M., El Refai, A., Ebead, U. and Nanni, A. (2017). “Corrosion-Damaged Beams

Repaired with Carbon-Fabric-Reinforced Cementitious Matrix,” The 5th International

Conference on Durability of Fiber Reinforced Polymer (FRP) Composites for Construction

and Rehabilitation of Structures, CDCC 2017, Sherbrooke, Canada, July 2017.

5. Elghazy, M., El Refai, A., Ebead, U. and Nanni, A. (2016). “Performance of Corrosion-Aged

Reinforced Concrete Beams Rehabilitated with Fabric-Reinforced Cementitious Matrix

(FRCM),” The 4th International Conference on Sustainability Construction Materials and

Technologies, SCMT4, Las Vegas, USA, August 2016.

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