ABSTRACT WILLIAMS, MARGARET KATHERINE. Debonding

ABSTRACT

WILLIAMS, MARGARET KATHERINE. Debonding Resistance of NSM CFRP Strips Bonded to Masonry under Various Environmental Exposures. (Under the direction of Dr. Rudolf Seracino).

Much research has been done on the repair or strengthening of masonry walls with both externally bonded and near-surface mounted FRP systems. However, research is needed to determine the durability of an FRP-reinforced masonry wall subjected to environmental conditions similar to what it may witness over its lifespan. The pull-test is commonly used to determine the bond behavior of the FRP and the masonry prism as it simulates the critical intermediate crack debonding failure mode.

This thesis describes the results of pull-tests performed on hollow core clay brick prisms strengthened with NSM carbon FRP. Three series of tests were performed at approximately four-month intervals during which the prisms were exposed to varying environmental conditions. While a portion of the constructed prisms remained in a controlled humidity and temperature setting, others were weathered outside in a natural environmental condition. The remaining prisms were placed in an extreme environmental condition, in which the prisms were subjected to large temperature ranges through the use of an environmental chamber. The change in debonding behavior due to varying the masonry pattern bond and exposure to environmental conditions was experimentally determined and compared to existing models. Further, a numerical analysis using the Applied Element

Method was completed on six different pull-test specimens in an effort to determine the suitability of using such a program to accurately model the pull-test.

The resulting pull-test behavior exhibited a significant change in the debonding failure of a naturally conditioned specimen due to exposure to high temperatures and moisture. The failure mechanism changed to shear failure in the adhesive. Further, the debonding resistances of the control specimens and specimens placed in the environmental chamber were also compromised by premature splitting of the prisms through the cores. This is turn lessened the debonding resistance of FRP-reinforced prisms. However, the numerical analysis program utilized in this research was able to capture the splitting failure of the control prisms with accuracy. Debonding Resistance of NSM CFRP Strips Bonded to Masonry under Various Environmental Exposures

by Margaret Katherine Williams

A thesis submitted to the Graduate Faculty of North Carolina State University in partial fulfillment of the requirements for the degree of Master of Science

Civil Engineering

Raleigh, North Carolina

2010

APPROVED BY:

______Dr. James Nau Dr. Sami Rizkalla

______Dr. Rudolf Seracino Chair of Advisory Committee DEDICATION

To my parents, Todd and Kitty Williams, and to my grandparents, Byron and Gayle

Williams, for their love, guidance, and support.

ii BIOGRAPHY

Margaret Williams was born to Todd and Kitty Williams and grew up in a small community outside of Cincinnati, Ohio. Her initial pursuit of a degree in civil engineering led her to Elon University located outside of Burlington, North Carolina. Elon University had a dual-degree engineering program in which she majored in Engineering Physics at Elon and was also selected as an Honors Fellow. She then spent the last two of the five-year program at North Carolina State University. Upon completion of the coursework at Elon and

N.C. State in May 2008, she graduated Cum Laude with a Bachelor of Science degree in

Engineering Physics from Elon University and Magna Cum Laude with a Bachelor of

Science degree in Civil Engineering from North Carolina State University. In the fall of

2008, she decided to continue her education at North Carolina State University for a Master of Science degree in Civil Engineering with a concentration in Structural Engineering and

Mechanics. Margaret is a registered Engineer-In-Training and following graduation, she will begin her professional career with an engineering firm located in Raleigh, N.C.

iii ACKNOWLEDGMENTS

First, I would like to express thanks to my advisor, Dr. Rudolf Seracino, for his direction and support throughout this research. The past two years in which he has served as my advisor has proved to be a valuable and enjoyable learning experience due to his involvement in the project. In addition, I would also like to express gratitude to committee members, Dr. Sami Rizkalla and Dr. James Nau, for their guidance and contributions to this research and throughout my graduate studies.

Further, I would like to thank the entire staff of the Constructed Facilities Lab, and particularly Jerry Atkinson, Greg Lucier, and Johnathan McEntire for their assistance throughout this research. Their experience and technical expertise facilitated the completion of my experimental program. Genuine thanks is also due to fellow graduate students John

Wylie, Adolfo Obregon-Salinas, Dillon Lunn, and Hartley Grimes for their help and advice throughout this research.

Finally and most importantly, I would also like to thank my family for their love, support, and enthusiasm in all my life’s endeavors. The successes that I have enjoyed have been a result of the opportunities provided to me by their guidance and encouragement.

iv TABLE OF CONTENTS

LIST OF TABLES ...... ix

LIST OF FIGURES ...... xi

CHAPTER 1 – INTRODUCTION ...... 1

1.1 Background ...... 1

1.2 Objective ...... 3

1.3 Scope...... 4

CHAPTER 2 – LITERATURE REVIEW ...... 6

2.1 Introduction ...... 6

2.2 FRP and Historic Masonry Structures ...... 7

2.3 FRP Material Properties ...... 9

2.3.1 History and Production ...... 9

2.3.2 Fiber Properties ...... 11

2.3.3 Resin Properties ...... 15

2.4 Durability ...... 16

2.5 FRP Strengthening Techniques ...... 21

2.6 Failure Mechanisms of FRP-Reinforced Structures ...... 24

2.7 Conclusions ...... 36

CHAPTER 3 – EXPERIMENTAL PROGRAM ...... 38

3.1 Introduction ...... 38

3.2 Pull-Test ...... 38

3.2.1 Test Specimens ...... 38

v 3.2.2 Masonry Prism Construction ...... 40

3.2.3 Experimental Set-Up ...... 51

3.2.4 Instrumentation ...... 52

3.2.5 Testing Procedure ...... 53

3.3 Testing Matrix ...... 53

3.4 Material Properties ...... 56

3.4.1 FRP Coupon Testing ...... 56

3.4.2 Masonry ...... 60

3.4.2.1 Brick Prism Compression Test...... 60

3.4.2.2 Brick Lateral Modulus of Rupture Test ...... 68

3.4.2.3 Compressive Strength of Brick Unit ...... 72

3.4.2.4 Mortar Cube Test ...... 75

CHAPTER 4 – TEST RESULTS ...... 78

4.1 Introduction ...... 78

4.2 Pull-Test Results...... 79

4.2.1 First Series (Control)...... 79

4.2.2 Second Series (5 Months Conditioning) ...... 86

4.2.3 Third Series (9 Months Conditioning) ...... 92

4.3 Discussion of Results ...... 98

4.3.1 Effects from Environmental Conditioning and Construction Pattern ...... 98

4.3.2 Discussion ...... 102

4.4 Comparisons to Previous Models ...... 110

vi 4.4.1 Willis Masonry Model ...... 110

4.4.2 Comparison to Petersen et al. (2009) Tests ...... 116

4.4.3 Comparison to ACI 440.7R-10 ...... 119

4.5 Conclusions ...... 120

CHAPTER 5 – NUMERICAL ANALYSIS ...... 123

5.1 Introduction ...... 123

5.2 Applied Element Modeling...... 123

5.3 Modeling Parameters in Extreme Loading for Structures® ...... 128

5.4 Numerical Analysis ...... 130

5.4.1 Introduction ...... 130

5.4.2 Current Study’s Specimen ...... 131

5.4.3 Numerical Analysis of Willis et al. (2009) Specimens ...... 146

5.4.4 Numerical Analysis of Petersen et al. (2009) Specimens ...... 161

5.5 Sensitivity Analysis ...... 174

5.6 Conclusions ...... 178

CHAPTER 6 – CONCLUSIONS & RECOMMENDATIONS ...... 181

6.1 Conclusions ...... 181

6.1.1 Splitting of Control Specimens ...... 181

6.1.2 Naturally Conditioned and Shear Failure in Adhesive ...... 183

6.1.3 Numerical Simulation ...... 184

6.2 Recommendations and Future Work ...... 185

6.3 Summary ...... 186

vii REFERENCES ...... 189

viii LIST OF TABLES

Table 2-1. Typical Fiber Properties of Common Materials (Gilstrap & Dolan, 1998) ...... 12

Table 3-1. Composite Gross Laminate Properties (Provided by Fyfe Co., LLC) ...... 42

Table 3-2. Tyfo TC Epoxy Properties (Provided by Fyfe Co., LLC) ...... 46

Table 3-3. Test Matrix ...... 55

Table 3-4. Tested FRP Coupon Properties ...... 60

Table 3-5. Masonry Prism Compressive Strengths and Elastic Modulus ...... 65

Table 3-6. Lateral Modulus of Rupture Results...... 72

Table 3-7. Compressive Strength of Brick Units ...... 75

Table 3-8. Compressive Strength of Mortar Cubes ...... 76

Table 4-1. Summary of Test Results for Series 1 ...... 85

Table 4-2. Summary of Test Results for Series 2 ...... 90

Table 4-3. Summary of Test Results for Series 3 ...... 96

Table 4-4. Theoretical versus Experimental Test Results ...... 112

Table 4-5. Theoretical versus Experimental Test Results, Petersen et al. (2009) Specimens

...... 117

Table 5-1. Material Properties ...... 134

Table 5-2. Comparison of Numerical to Experimental Data – Stack Bonded Specimens .. 139

Table 5-3. Comparison of Numerical to Experimental Data – Running Bonded Specimens

...... 145

Table 5-4. Daniels et al. (2009) Material Properties ...... 149

ix Table 5-5. Comparison of Numerical to Experimental Data – Daniels et al. (2009) Stack-

Bonded Specimens ...... 155

Table 5-6. Willis et al. (2009) Material Properties ...... 156

Table 5-7. Comparison of Numerical to Experimental Data – Running Bonded Specimens

...... 161

Table 5-8. Petersen (2009) Material Properties ...... 164

Table 5-9. Comparison of Numerical to Experimental Data – Stack Bonded Specimens .. 169

Table 5-10. Comparison of Numerical to Experimental Data – Running Bonded Specimens

...... 173

Table 5-11. Comparison of Stack-Bonded Specimens through Sensitivity Analysis of Mortar

...... 176

Table 5-12. Comparison of Stack-Bonded Specimens through Sensitivity Analysis of Epoxy

...... 177

x LIST OF FIGURES

Figure 2-1. Out-of-Plane Failures in Historic Masonry Structures (Valluzzi, 2007) ...... 8

Figure 2-2. Applications of Fiber Reinforced Polymers Noted in Literature Review

(Uomoto, Mutsuyoshi, Katsuki, & Misra, 2002) ...... 10

Figure 2-3. Typical Stress-Strain Curves (Bakis C. E., 2009) ...... 13

Figure 2-4. Tensile Failure of FRP Rods (Uomoto et al., 2002) ...... 14

Figure 2-5. Pull-Test Set-Up (Fava et al., 2009)...... 17

Figure 2-6. Typical Strengthened Specimen (Bati & Rotunno, 2001) ...... 19

Figure 2-7. Schematic of FRP-Reinforced Cross-Sections ...... 22

Figure 2-8. IC Debonding Mechanism (Seracino, Jones, Ali, Page, & Oehlers, 2007) ...... 25

Figure 2-9. Typical Masonry Pull-Test Specimen ...... 26

Figure 2-10. Prism Specimens and FRP Locations (Willis et al., 2009) ...... 28

Figure 2-11. Shear Stress in an EB Pull-Test (Yuan et al., 2004) ...... 30

Figure 2-12. Bilinear Bond-Slip Model (Yuan et al., 2004) ...... 30

Figure 2-13. Debonding Failure Planes (Seracino et al., 2007b) ...... 31

Figure 2-14. Pull Test Specimens (Petersen et al., 2009) ...... 33

Figure 2-15. Crack Pattern on Face of Specimen (Petersen et al., 2009) ...... 34

Figure 2-16. Cracking Through Specimen (Petersen et al., 2009) ...... 35

Figure 3-1. Brick Dimensions ...... 39

Figure 3-2. FRP-to-Masonry Joint Prisms (A) Stacked Bond and (B) Running Bond ...... 39

Figure 3-3. Placing a Brick Course ...... 41

Figure 3-4. Leveling Brick Course ...... 41

xi Figure 3-5. Marking Groove Position ...... 43

Figure 3-6. Creating Groove in Brick Prism ...... 44

Figure 3-7. Final Preparation of Brick Prisms for FRP ...... 45

Figure 3-8. Filling Groove with Epoxy...... 47

Figure 3-9. (A) Application of Brick Dust (B) Brick Dust Set on Prism ...... 47

Figure 3-10. Epoxy Filled Groove ...... 48

Figure 3-11. Orientation of Fibers for Tabs ...... 50

Figure 3-12. Construction of Tab...... 51

Figure 3-13. Pull-Test Set-up ...... 52

Figure 3-14. Pull-Test Instrumentation ...... 53

Figure 3-15. Typical Set-Up for FRP Coupon Testing ...... 58

Figure 3-16. Typical Failure of FRP Coupon ...... 58

Figure 3-17. Stress - Strain Curves for FRP Coupon Tensile Testing ...... 59

Figure 3-18. Set-Up of Compression Tests...... 61

Figure 3-19. Back Face of Masonry Prism with Hydro-stone ...... 62

Figure 3-20. Vertical Cracks on Front and Back of Masonry Prism ...... 63

Figure 3-21. Masonry Prism at Failure ...... 63

Figure 3-22. Masonry Stress-Strain Curves at 5 Months ...... 66

Figure 3-23. Load vs. Stroke for Prism C-5...... 67

Figure 3-24. Strength Increase in Control Specimens with Time ...... 68

Figure 3-25. Compressive Strength of Conditioned Specimen vs. Control ...... 68

Figure 3-26. Lateral Modulus of Rupture Test Set-Up ...... 69

xii Figure 3-27. Typical Break in Center Brick Course ...... 70

Figure 3-28. Typical Break through Thinnest Section of Brick ...... 70

Figure 3-29. Half-Brick Test Specimen ...... 73

Figure 3-30. Brick Compression Test Set-Up...... 74

Figure 3-31. Failure of Specimen ...... 74

Figure 3-32. Mortar Cube Test Set-Up ...... 76

Figure 3-33. Typical Failure Mode of Mortar Cube ...... 77

Figure 4-1. Crack Propagation Across Top of Brick ...... 80

Figure 4-2. Crack Propagation on Front Face of Masonry Prism ...... 81

Figure 4-3. Continuation of Crack Propagation ...... 81

Figure 4-4. Complete Debonding of FRP ...... 82

Figure 4-5. Splitting of Masonry Prism ...... 82

Figure 4-6. Splitting Across Top Brick Course ...... 83

Figure 4-7. FRP Strip Removed from Specimen ...... 84

Figure 4-8. Load-Slip Curves for Specimens in First Series ...... 86

Figure 4-9. Crack Propagation in CR1-5 ...... 88

Figure 4-10. Cracking on Back Face of CR1-5 ...... 88

Figure 4-11. FRP Strip Removed from (A) Control (B) NR-5 ...... 89

Figure 4-12. Specimen NR-5 with FRP Removed (Front Face) ...... 89

Figure 4-13. Load-Slip Curves for Stack-Bonded Specimens in Series 2 ...... 91

Figure 4-14. Load-Slip Curves for Running Bonded Specimens in Series 2 ...... 92

Figure 4-15. Color of Epoxy Following Conditioning ...... 93

xiii Figure 4-16. Cracking on Masonry Prism (Natural Conditioning) ...... 94

Figure 4-17. FRP Removed from Specimen (Natural Conditioning) ...... 94

Figure 4-18. Groove in Specimen (Natural Conditioning) ...... 95

Figure 4-19. Load-Slip Curves for Stack-Bonded Specimens in Series 3 ...... 97

Figure 4-20. Load-Slip Curves for Running Bonded Specimens in Series 3 ...... 98

Figure 4-21. Peak Loads of Control Specimens with Time ...... 100

Figure 4-22. Debonding Resistance of Conditioned Specimens with Time ...... 102

Figure 4-23. FRP Removed from Natural Conditioning (A) At 0 Months (B) At 5 Months

Figure 4-24. Daily Temperature Data for Naturally Conditioned Specimens ...... 106

Figure 4-25. Daily Precipitation Data for Naturally Conditioned Specimens ...... 107

Figure 4-26. Extreme Conditioning – Cycle Type 1...... 109

Figure 4-27. Extreme Conditioning – Cycle Type 2...... 109

Figure 4-28. Comparison of Load-Slip Curves for Willis et al. (2009) Study versus

Specimen in Current Study ...... 115

Figure 4-29. Load-Slip Behavior, Petersen et al. (2009) Specimens vs. Current Study

Specimens ...... 118

Figure 5-1. AEM Modeling (A) Element Generation (B) Spring Distribution of each Pair of

Springs (Meguro & Tagel-Din, 2000) ...... 126

Figure 5-2. Current Study's Brick Dimensions (Actual) ...... 131

Figure 5-3. Pull-Test Specimens ...... 131

Figure 5-4. Crack Propagation in Stack-Bonded Specimen ...... 135

xiv Figure 5-5. Comparison of Numerical to Experimental Load Curves – Stack-Bonded

Specimens ...... 139

Figure 5-6. Splitting of Top Brick ...... 140

Figure 5-7. Crack Propagation in Running Bonded Specimens ...... 141

Figure 5-8. Comparison of Numerical Analysis to Experimental Data – Running Bonded

Specimens ...... 145

Figure 5-9. Nominal Brick Dimensions in Willis et al. (2009) and Daniels et al. (2009)

Specimens ...... 146

Figure 5-10. Daniels et al. (2009) and Willis et al. (2009) Pull-Test Specimens ...... 147

Figure 5-11. Crack Propagation in Daniels et al. (2009) Stack-Bonded Specimen ...... 150

Figure 5-12. Comparison of Numerical Analysis to Experimental Data – Daniels et al.

(2009) Stack-Bonded Specimens ...... 154

Figure 5-13. Crack Propagation in Willis et al. (2009) Specimen ...... 157

Figure 5-14. Comparison of Numerical Analysis to Experimental Data – Willis et al. (2009)

Running Bonded Specimens ...... 160

Figure 5-15. Brick Dimensions Used in Petersen et al. (2009) Study ...... 161

Figure 5-16. Petersen et al. (2009) Pull-Test Specimens ...... 162

Figure 5-17. Cracking Behavior of Petersen et al. (2009) Specimen ...... 165

Figure 5-18. Load-Slip Curves – Petersen (2009) Stack-Bonded Specimen ...... 168

Figure 5-19. Cracking Behavior of Petersen (2009) Running Bond Specimen ...... 170

Figure 5-20. Load-Slip Curves – Petersen (2009) Running Bonded Specimens ...... 173

Figure 5-21. Load-Slip Curves with Decrease in Mortar Strength and Stiffness ...... 176

xv Figure 5-22. Load-Slip Curves with Decrease in Epoxy Strength and Stiffness ...... 177

Figure 6-1. Splitting Through of Prisms ...... 182

Figure 6-2. Cracking Pattern of URM Wall (Wylie, 2009) ...... 183

xvi CHAPTER 1 – INTRODUCTION

1.1 Background

Today, masonry buildings make up a large portion of not only our entire building stock, but also the historical structures sector. These structures are usually in need of repair due to flaws in the construction process, settling of the foundation, or deterioration from weathering. They are also often strengthened following an increased demand of its load capacity due to changes in occupancy for example. Further, these buildings must be continually upgraded to the most current building codes (Triantafillou, 1998). Building codes for new masonry construction published in the 1970’s increased the amount of bracing required by half in an effort to reduce damage from seismic activity. This shows how necessary it was to also strengthen older structures (Gilstrap & Dolan, 1998). Failure of unreinforced masonry structures during seismic activity causes the most damage and results in more human death than other structural failures. It is imperative to strengthen these unreinforced masonry structures to guarantee the safety of the occupants as well as to save money by avoiding damages (Korany & Drysdale, 2006).

Previously, masonry structures were upgraded by grouting cracks, stitching large cracks with additional bricks, applying reinforced grouted perforations, using steel to post tension the structure, or using concrete with steel reinforcement to jacket the entire structure.

The last method was the most common, however, the added mass to the structure caused problems, as the lower floors could not resist the increase in dead load. Further, the increase in dead load impairs the dynamic behavior of the structure, which is particularly worrying in

1 high seismic regions. Jacketing of the structure also altered its aesthetics, which in most cases was extremely undesirable as many structures are of historic and architectural value.

Finally, the process was labor-intensive and areas near the repair work must to be taken out of service (Triantafillou, 1998).

One potential solution to the problems caused by the use steel jacketing is the use of fiber-reinforced polymers (FRP). These composites are typically made from glass, carbon, or aramid and are bonded together using a polymer matrix. FRP offers a high-strength and stiffness at a low weight. Further, FRP is not prone to corrosion, and is available in many different types such as plates or fabrics (Triantafillou, 1998). FRP is also relatively easy to install and subsequently, the time required to repair the structure will be decreased (Korany

& Drysdale, 2006).

The first use of FRP to strengthen a masonry system was by Sweiden in 1991 who analyzed the use of FRP in post-tensioning masonry structures. Triantafillou and Fardis then used FRP as a method to circumferentially prestress masonry structures. This was followed by Schwegler in 1994 who used carbon FRP (CFRP) that was externally bonded to the face of the masonry surface. This acted as a tensile reinforcement for the structure and is the basis of the research presented in this thesis (Triantafillou, 1998).

Further research was conducted that proved the efficiency of using FRP to strengthen masonry structures. FRP is applied to the surface of a masonry structure in one of two ways; it is either externally bonded (EB) to the surface or it is inserted into a precut groove on the face of the structure in what is called the near-surface mounted (NSM) technique. Studies have shown that the NSM technique is more efficient in increasing the load capacity of the

2 wall, especially when using thin FRP strips. It is also less visibly invasive as the FRP is hidden and protected in the face of the structure. With this, the FRP is less prone to damage from vandalism or weathering (Petersen, Masia, & Seracino, 2009).

The pull-test, which subjects the FRP in a NSM-FRP reinforced masonry prism to a tensile load, simulates the behavior of an FRP-reinforced wall subjected to out-of-plane bending. The pull-test replicates a section of the wall centered about a FRP strip and can help determine the increase in strength of a wall with the addition of FRP (Willis, Yang,

Seracino, & Griffith, 2009). The effects of weathering or masonry wall construction type can be studied through changes in the ultimate load observed in a pull-test.

1.2 Objective

Much research has been focused on the advantages of the NSM technique but little has been studied on the behavior of FRP-reinforced masonry structures after being exposed to various environmental conditioning. This research is necessary in order to determine the behavior of these structures in service. Utilizing the pull-test, the effects of environmental degradation on the bond strength of FRP-to-masonry joints can be assessed by observing the change in debonding resistance after exposing the masonry specimens to environmental conditioning. Further, the effects of the pattern bond of the masonry will be studied, as prisms will be constructed in the stack-bond and running bond patterns. The type of construction changes the distance between the FRP strip and the nearest vertical mortar joints as this will affect the debonding resistance determined by the pull-test.

3 The debonding resistances determined in this study will be compared to published models that predict the debonding resistance of a FRP-reinforced masonry prism subjected to a pull-test. Furthermore, the pull-test will be analytically studied through the program

Extreme Loading for Structures®, which utilizes the Applied Element Method. This numerical method will be validated by comparing experimental results found in this study as well as in other published studies as prism construction and material properties can be easily altered within the program.

1.3 Scope

The objectives described in the previous section will be met in this thesis through the research presented in the following chapters.

 Chapter 2 encompasses the literature review that examined past studies on FRP-

reinforced masonry. The chapter includes background on FRP material properties,

use, and durability in masonry structures, and models that predict the debonding

resistance. Furthermore, the technical background of the Applied Element Method is

presented.

 Chapter 3 describes the experimental program implemented in this research. It

includes specimen construction and pull-test set-up as well as the testing protocols

executed for testing the material properties of the brick, mortar, and FRP.

 Chapter 4 discusses the resulting debonding behavior of specimens subjected to the

pull-test. Effects of environmental conditioning and construction pattern are also

examined in this chapter. The debonding resistances determined in this chapter will

4 be compared to previously published studies and reasons for any deviations will be

discussed.

 Chapter 5 explores the use of the program Extreme Loading for Structures® and

Applied Element Modeling to study the behavior of FRP-reinforced masonry prisms

subjected to a pull-test.

 Chapter 6 concludes the findings presented in this thesis as well as presents future

research needs and recommendations associated with FRP-reinforced masonry.

5 CHAPTER 2 – LITERATURE REVIEW

2.1 Introduction

Reinforcing techniques that include the use of fiber-reinforced polymers (FRP) are both efficient and cost-effective when strengthening or repairing masonry structures.

Unreinforced masonry (URM) structures worldwide are prone to failure stemming from environmental degradation and exposure to extreme loading. Previously, steel was chiefly used as reinforcement for strengthening or repairing URM structures yet the associated increase in weight and durability issues was undesirable in many applications. FRP can be used in many strengthening and repair applications due to its high strength and stiffness to weight ratio. It can also be applied in numerous orientations to a structure. Certain installations of FRP may be more desirable than others in preserving the aesthetics of the structure by minimizing the visibility of the FRP itself. This advantage can be particularly attractive when strengthening or repairing historic structures.

Much research has focused on the advantages of utilizing FRP for strengthening or repairing URM structures but little has been studied on the behavior of FRP-reinforced masonry structures after being exposed to various environmental conditions. This research is necessary in order to determine the behavior of these structures in service. In order to simulate the fundamental intermediate crack debonding mechanism, small-scale pull-tests, can be performed. With the pull-test, the effects of environmental degradation on the bond strength of FRP-to-masonry joints can be assessed by observing the reduction in debonding resistance after exposing the specimens to environmental conditioning. This literature review

6 will focus on the following topics associated with the use of FRP in unreinforced masonry structures:

 FRP and Historic Masonry Structures

 FRP Material Properties

 Durability

 FRP Strengthening Techniques

 Pull Tests

2.2 FRP and Historic Masonry Structures

Historical structures must not only be repaired for the current occupant’s sake but they must also be preserved for their heritage. Many of these structures contain historical pieces of art such as murals, carved stone, and other artistic details. To properly preserve the historic structure, these pieces must be salvaged. Many of these structures do not meet current building code standards or have been overstressed due to extreme loadings. Although many historical structures have endured over a long period of time, the quality of the materials used in the construction of such buildings is below today’s standard. Historical masonry structures are particularly susceptible to failure from seismic activity and are vulnerable to weathering associated with freeze-thaw cycles and moisture. Many structures are also experiencing problems stemming from modern technologies such as vibrations from traffic or household appliances (Szabó & Kirizsán, 2008). Figure 2-1 depicts failures stemming from out-of-plane bending for different masonry structures (Valluzzi, 2007).

Figure 2-1. Out-of-Plane Failures in Historic Masonry Structures (Valluzzi, 2007)

FRP systems have the ability to prevent failures for historical structures by strengthening walls for shear and bending, columns for axial loadings, and supporting curved members or sharp corners in a structure. Allowable repair methods set in place by the

Charter of Venice, the Declaration of Amsterdam, and the Chart of Cracovia require that any strengthening system for existing masonry structures must be minimally invasive, reversible, durable, include compatible materials to the original materials used, and uphold the original function of the structure. FRP meets many of these requirements however removing the bonded system if necessary and the durability and compatibility of FRP with masonry structures have not yet been studied fully (Valluzzi, 2008).

Nevertheless, there have been multiple case studies where FRP has successfully been applied to historical structures. Large cracks in a vault at the Villa Bruni in Megliadino S.

Vitale and at the church of S.ta Corona of Vicenza have been repaired with FRP. The walls at the church of S.ta Maria in Organo of Verona have been effectively reinforced in the in-

8 plane direction. The top of columns were confined using FRP at the Palazzo della Ragione in Padova. Even statues have been reinforced with FRP such as the equestrian statue at the tomb of Cansignorio in Verona in which the base of the tomb was damaged but was remedied through the application of carbon FRP. Marble chipping is a technique that can then be used to hide the FRP system on the statue. Therefore, FRP’s ability to be integrated into the aesthetics of the structure can make it an advantageous strengthening system compared to steel, particularly in the case of historic structures (Valluzzi, 2008).

2.3 FRP Material Properties

2.3.1 History and Production

Development of the FRP materials that are used today to repair or strengthen historical structures began in the 1940’s as the petrochemical industry boomed as a result of

World War II. The increase of funding to programs that encouraged the exploration of space and to extending air travel during the 1960’s and 70’s led to new fiber materials such as boron, carbon, and aramid. These new materials possessed even higher strengths and stiffness at lower weights than their predecessors. To further expand the market for FRP, an effort was made to reduce the cost in producing and distributing FRP. Companies addressed sporting goods companies in an effort to attract them in using FRP in their products.

Particularly with the end of the Arms Race in the 1980’s, it was crucial for the FRP industry to adapt to other markets. Further lowering the cost of the product made FRP an attractive option to those interested in rehabilitating civil infrastructure. Additionally, the government encouraged the use of FRP in civil engineering applications by sponsoring research that

9 incorporated FRP products (Bakis, et al., 2002). FRP systems are now used in civil engineering applications such as strengthening structures in seismic areas and repairing structures failing in flexure (Valluzzi, 2008). Evidence that FRP is becoming increasingly more accepted in a vast array of construction uses is apparent in Figure 2-2.

Figure 2-2. Applications of Fiber Reinforced Polymers Noted in Literature Review (Uomoto, Mutsuyoshi, Katsuki, & Misra, 2002)

Fiber-reinforced polymers (FRP) are comprised of fibers that are joined together and protected by a polymeric matrix. FRP is a composite material that is low in weight yet high in tensile strength. FRP is also of high stiffness and is a desirable material to use in situations where corrosion and fatigue are problematic. Strengthening systems that use FRP

10 are also easy to install (Valluzzi, 2008). Further, FRP systems are less expensive than other typical reinforcing materials for the significant gain in strength (Bakis, et al., 2002).

FRP cross-sections are produced through a pultrusion process in which fibers are impregnated with resin and then are pulled through a heated shaping die. When pulling the fibers through the resin, a small tensile load is applied. This tensile load enables the fibers to straighten resulting in a uniform cross-section. The applied load is kept to a minimum so that it does not deform the FRP, resulting in a lower ultimate load capacity. Fibers can also be bonded together in various directions to create grids or fabrics (Uomoto et al., 2002).

Typically, FRP fabrics are described as being chopped mats, woven, or individual tapes of fibers where the first group involves assembling FRP in random orientations. These chopped mats have been shown to strengthen structures for in-plane shear forces and with proper design, out-of-plane forces. Conversely, fibers in woven mats are interwoven in only two directions whereas tapes are thinner slices of a larger fabric piece (Gilstrap & Dolan, 1998).

FRP sheets can be adhered together in varying directions or thicknesses on-site as the resin is applied in the field. Further, the thicknesses of FRP sheets versus plates are different in that sheets typically have thicknesses of 0.0079 inches (0.2 mm) and plates are 0.047 inches (1.2 mm) (Balendran, Rana, & Nadeem, 2001).

2.3.2 Fiber Properties

For typical FRP products, 50 to 65% of the product is comprised of fibers. The strength of the FRP is dependent on the material properties of these fibers (typically carbon, aramid, glass, or polyvinyl alcohol), the matrix material, and the amount of fiber in the product. Carbon fibers are comprised of fine graphite crystals whereas aramid fibers are

11 comprised of organic materials and are classified as Kevlar, Twaron, or Technora. The fiber’s nuclei in types Kevlar and Twaron’s are bound by aramid whereas the nucleus is bound by ether in the Technora type of fiber. Glass fibers are classified as E-glass or alkali- resistant fibers where alkali-resistant fibers have the distinction of containing zirconia, which deters corrosion. High-strength polyvinyl alcohol fibers are rolled during production to further increase the strength and ductility of the fibers. Recent research has focused on testing hybrid FRP materials that use different combinations of fiber materials. For example, aramid fibers can be used to surround and protect glass fibers from corrosive environments

(Uomoto et al., 2002). Table 2-1 briefly summarizes the differences in common fiber properties.

Table 2-1. Typical Fiber Properties of Common Materials (Gilstrap & Dolan, 1998)

Composite Fiber Tensile Fiber Composite Fiber Tensile Strain at Strength, GPa Modulus, Modulus GPa, Type Strength, GPa Failure (ksi) GPa (ksi) (ksi) (ksi) Aramid 3.66 (525) 125 (18,000) 1.54 (220) 84 (12,000) 0.024 Carbon 3.66 (525) 228 (33,000) 1.75 (250) 132 (19,000) 0.012 E-glass 2.10 (300) 75 (11,000) 0.83 (120) 49 (7,000) 0.03

Additionally, the difference between the ultimate tensile strength and ductility of these FRP materials compared to steel is represented in Figure 2-3. In this figure, it is apparent that FRP materials reach higher tensile strengths than steel however, steel exhibits a more ductile behavior. FRP strains are low before observing sudden failure. Up to failure, the stress-strain curve shows that the FRP material behaves linear-elastically. At ultimate, the fibers fail by rupture as seen in Figure 2-4, however failure can also happen by cracking

12 of the matrix, failure of the adhesive, debonding between the fiber and the polymer matrix, and cracking down the length of the FRP (Uomoto et al., 2002). The yield strain typically associated with Grade 60 (420 MPa) reinforcement steel is 0.002, which is only 10% of the ultimate strain reached by FRP. Therefore, structures reinforced with FRP will normally fail by failure through the adhesive, FRP rupture, or crushing of the masonry or concrete. With additional FRP reinforcement, the structure will behave more brittle in bending.

Furthermore, the shear capacity observed for FRP is only one-tenth of the tensile strength whereas in steel, it is normally 45% of the tensile strength (Gilstrap & Dolan, 1998).

600 Carbon 4,000 500 Aramid Glass 3,500 3,000 400 2,500

300 2,000 Stress, ksi Stress, 200 1,500 MPa Stress, 1,000 100 Steel 0.2 500 0 0 0 0.02 0.04 0.06 0.08 0.1 Strain

Figure 2-3. Typical Stress-Strain Curves (Bakis C. E., 2009)

Figure 2-4. Tensile Failure of FRP Rods (Uomoto et al., 2002)

It has also been observed that the fibers cannot uphold sustained loads as well as steel. When fibers are stressed to up to 90% of their maximum load, the fibers will eventually rupture due to a phenomenon called creep-rupture (Gilstrap & Dolan, 1998).

Studies have shown that glass, aramid, and carbon can hold up to 30, 50, and 90% of their maximum strength, respectively. Per the guidelines of the American Concrete Institute, the maximum load that should be applied to the fibers to avoid rupture and to include cyclic loading is 20, 30, and 55% of ultimate for glass, aramid, and carbon fibers, respectively (ACI

440.2R-08, 2008). This is primarily for prestressing applications and does not apply when

FRP is adhered to the exterior of a structure (Gilstrap & Dolan, 1998).

14 2.3.3 Resin Properties

The resins used for tying the fibers together have no structural purpose other than to protect and transfer the load among fibers. When each individual fiber is covered in resin, the overall durability of the FRP composite is increased. For each material type of fiber used, it is necessary to choose the corresponding resin that has compatible strain capacities.

For example, epoxy resins are compatible with carbon fibers due to the small strain deformations observed in each material. Epoxy is not as well matched with aramid or glass fibers because these types of FRP have larger strain deformations. It is also important to properly apply the resin to the FRP. Uneven application of the resin during FRP production may result in high stress concentrations on some fibers, which will eventually cause failure of the specimen at a lower load than anticipated (Uomoto et al., 2002).

Previously identified problems regarding the use of FRP are being solved with further research and the incorporation of its use in engineering codes and guidelines. These problems include the elevated initial cost, low strength in directions not parallel to the fibers, low ductility, and lack of use in the field. Furthermore, the long-term behavior of FRP has not been observed in real structures as this is a relatively new strengthening system for civil engineering structures. The material’s lack of resistance to high temperatures and the limited research on the durability and necessary upkeep for FRP-reinforced structures has also been cited as a potential issue for the use of FRP (Bakis C. E., 2009). Known durability issues such as damage from UV radiation and water absorption, particularly for aramid fibers, have also been cited to cause these concerns (Erki & Rizkalla, 1993).

15 2.4 Durability

Although the material properties of FRP have been well studied and shown to create an advantageous system in strengthening and repairing existing masonry structures, how the

FRP system endures environmental degradation and aging has not been well researched. Past studies have looked at the service life of the masonry and FRP as individual structural components however, little is known about the durability of the bond between the FRP and the masonry structure (Bati & Rotunno, 2001). There is potential for the bond to weaken if the reinforced wall is subjected to a wide range of temperatures or water, particularly when present in walls that undergo freeze-thaw cycles. Deterioration of the bond may not be visible from the surface of the wall but could cause premature failure of the strengthening system (Gilstrap & Dolan, 1998).

Individually, the fibers themselves have varying coefficients of thermal expansion depending on the material and direction of the fibers. For example, carbon fibers have a longitudinal coefficient of thermal expansion near zero but glass fibers have a similar coefficient of thermal expansion as concrete. In the transverse direction, coefficients of thermal expansion can be twice that of concrete, which could cause degradation of the adhesive layer between the FRP and substrate when the structure is subjected to high-low temperature cycles (Gilstrap & Dolan, 1998).

Some studies on the durability of the bond between FRP and concrete structures have exposed specimens to either extreme temperature or humidity ranges. In many cases, these studies have not exposed specimens to cyclic actions of either condition. Deicing salts and freeze-thaw conditions can also adversely affect the bond but are currently under-researched

16 topics such as in the study by Fava, Mazotti, Poggi, and Savoia (2009). A number of FRP- reinforced concrete specimens were exposed to thermal cycles as well as salt-spray. Each concrete specimen was reinforced with externally bonded carbon FRP plates or sheets. The first set of specimens was set in an environmental chamber while another set of specimens were exposed to the natural environment in the open air. Specimens in the environmental chamber underwent freeze-thaw cycles in which the temperature was held for five hours at -

0.4° F (-18° C) or 39° F (4° C) or they were subjected to salt spray fog conditioning. A final set of specimens were held as controls. After a predetermined number of cycles, specimens from each conditioning group were tested using the pull-test shown in Figure 2-5 in which the FRP was subjected to a purely tensile load.

Figure 2-5. Pull-Test Set-Up (Fava et al., 2009)

The experimental results showed that the freeze-thaw cycles decreased the peak load whereas the salt spray actually increased it for specimens reinforced with FRP plates. The load and elongation of the FRP was recorded for each specimen and it was observed that initially, the graph was linear until reaching a peak load at which point the line began to plateau. For the conditioned specimens, the plateau began at a lower load than the control

17 specimens. It was also observed that the debonding strength of the FRP-reinforced specimens increased with time, which was a product of the concrete gaining strength with age. Furthermore, it was observed that the sheets protected the concrete better than the plates against the freeze-thaw cycles and the salt spray. Therefore, these specimens were less affected by the conditioning treatments (Fava et al., 2009).

Bati and Rotunno (2001) from the University of Florence studied the effects of exposing masonry specimens strengthened with FRP to wet-dry and freeze-thaw cycles.

Seventy-eight specimens were divided into sets of six with an additional set of six specimens set aside as controls. Each specimen was assembled with three bricks, bonded with mortar, and then adhered to an FRP strip across both sides of the masonry prism as shown in Figure

2-6.

Three conditioning methods as part of an accelerated testing program were employed that would simulate the natural weathering that a masonry structure would be subjected to during its service-life in central Italy. A wet-dry cycle was achieved by submerging the specimens into a water bath for ten minutes and then drying for 48 hours. This was performed on 24 of the specimens, divided into sets of six that were then subjected to six, twelve, twenty-four, or forty-eight cycles. Similarly, another 24 specimens were submerged in water for 20 minutes and then allowed to dry for 24 hours. An additional four sets of specimens were subjected to freeze-thaw cycles controlled by a thermostatic treatment chamber that consisted of exposing the specimens to 122° F (50° C) for six hours and then

18° F (-8° C) for six hours. Each set of six specimens were either exposed to twelve, twenty- four, forty-eight, or ninety-six cycles.

Figure 2-6. Typical Strengthened Specimen (Bati & Rotunno, 2001)

Following conditioning, each specimen was tested in shear by applying load to the center brick unit. Although the application of load is different from a typical pull-test in which load is applied to the FRP, comparisons on the effects of environmental conditioning can be drawn. Bati and Rotunno (2001) found that when comparing the peak load of the control specimens to an unconditioned set of reinforced specimens, the maximum load was twice that of the unreinforced specimens and were stiffer, showing the added advantage of strengthening with the FRP. When comparing the specimens that were subjected to wet-dry cycles, it was observed that for the specimens immersed in water for twenty minutes, the strength was initially lower than for those specimens immersed for only ten minutes.

However, with an increase in the number of cycles, both sets reached a region of little variation in which no further loss in strength was observed. The final, total loss in both

19 conditioning cases was only 18%. Therefore, the researchers concluded that FRP-reinforced specimens performed well in wet-dry cycles.

The specimens exposed to freeze-thaw cycles on the other hand, exhibited a 50% reduction in strength by the 96th cycle due to a complete deterioration of the bond between the FRP and masonry. Up to the twenty-fourth cycle, there was only a slight reduction in strength compared to the control specimens however following this cycle, visible cracks were present on the bond between the FRP and brick. These cracks were more evident as the number of cycles increased and eventually ended with the complete loss of the FRP reinforcement. The peak load recorded at the 96th cycle is comparable to the peak load observed for the unreinforced specimens. Therefore, significant freeze-thaw cycles could potentially cause the loss of the entire FRP strengthening system in a masonry structure (Bati

& Rotunno, 2001).

High temperatures are more critical for the FRP systems as the adhesive’s strength and elastic modulus are adversely affected when temperatures reach the glass transition temperature (Tg). Typically, the Tg for epoxies used in FRP systems is 131° F (55° C). At this temperature, the epoxy resin changes texture; what was once glass-like becomes of a plastic consistency. Further, as temperatures reach up to just 68° F (20° C) less than the Tg, creep effects increase and the strength decreases significantly. Currently, the recommendation is to use 50% of the short-term load of the adhesive for designing structures due to the uncertainty of the long-term in-situ behavior of these systems. Research has concluded that the long-term durability of FRP-reinforced structures can be traced to the

20 adhesive used in the system and thus, must be studied further as a part of the FRP strengthening system (Borchert & Zilch, 2005).

Further, both the FRP and adhesive are sensitive to ultra-violet (UV) rays, which can impair the long-term behavior of an FRP-reinforced structure. Typically, only the top surface of the system is affected by UV rays, however this could still cause a premature failure. Aldajah, Al-omari, and Biddah (2009) subjected both FRP sheets and resin to UV rays in 8-hour increments. Results showed that the elastic modulus decreased by 11% on average for three different types of commercially available resins after being subjected to a three-point bending test. Further, the elastic modulus of the FRP system also decreased for all three brands. The resin and FRP were studied further through the use of a scanning electron microscope and it was discovered that the UV rays caused significant microcracking in the resin. The FRP also displayed signs that the matrix fibers had debonded and cracked

(Aldajah et al., 2009). Therefore, it is important to understand the long-term behavior of the adhesive to ensure the durability of the system. Current design guidelines for strengthening structures when exposed to harsh weather conditions must be extremely conservative due to the lack of research on the durability of the FRP system (Fava et al., 2009)

2.5 FRP Strengthening Techniques

FRP can be applied to a structural member in different installation methods, some of which may better protect the FRP from environmental degradation. Externally bonded (EB) systems are those in which FRP sheets or plates are adhered to the surface of the structure.

When applied in the near-surface mounted (NSM) scheme, the FRP is placed in an epoxy-

21 filled groove cut into the face of a structure as seen in Figure 2-7. In an effort to preserve the aesthetics of the structure, FRP bars or strips can be installed in masonry bed joints in what is referred to as structural repointing (Valluzzi, 2008).

EB NSM

Figure 2-7. Schematic of FRP-Reinforced Cross-Sections

Research has shown that there are many advantages in using the near-surface mounted technique over externally bonding the FRP to the structure. Although application to curved or overhead surfaces is easier with the use of EB sheets or plates, respectively, these systems are not as efficient (Erki & Rizkalla, 1993; Petersen et al., 2009). Near-surface mounting the FRP into grooves cut on the face of the masonry wall significantly increases the strength developed by a FRP-reinforced structure. When the FRP debonds from the structure, the strengthening system is lost as the FRP is no longer adhered to the structure

(Petersen et al., 2009). Furthermore, NSM FRP can be anchored into other structural members more easily than in EB FRP systems. This can limit the likelihood that debonding occurs. This application is beneficial for situations where the FRP is anchored at an area of high moment as debonding will be less likely to occur (De Lorenzis & Teng, 2007).

Additional advantages of using a NSM FRP strengthening system include the concealment of the FRP in the face of the masonry structure, thus safeguarding the aesthetics of the structure. Being protected within the face of the wall also decreases the chance that

22 the FRP will be damaged from vandalism, fire, or weathering (Petersen et al., 2009).

Externally bonding FRP to a structure requires the surface to be prepped by creating a clean and smooth surface. NSM FRP only requires cutting of the grooves and therefore, the amount of time required to install the NSM FRP strengthening system is significantly less than it is for EB FRP systems. Not only does this save on labor costs but it also helps the project stay on schedule (De Lorenzis & Teng, 2007).

The cross-sectional shape of the FRP also varies with its intended use. Round bars are typically used in prestressing applications (De Lorenzis & Teng, 2007). However, bars are not as well confined in a masonry wall and it is more likely that debonding will occur.

Therefore, the most efficient geometric shape for NSM FRP is a thin, rectangular strip as it is the most confined by a cut groove (Petersen et al., 2009). Additionally, the bonded surface area is increased by using narrow strips. The surface texture of the FRP can also vary from being smooth, rough, ribbed, or spirally wound (De Lorenzis & Teng, 2007).

Generally, a two-part epoxy is used for filling the grooves in the NSM technique as well as in externally bonded applications. The epoxy serves the purpose of transferring the load from the FRP to the structure. The epoxy can be either high-viscosity, which is helpful when working overhead as it does not drip, or low-viscosity, which can be easily poured. In some cases, sand may also be added to the epoxy to alter the epoxy’s material properties such as the volume, viscosity, or thermal coefficient of expansion. However, this may weaken the bond strength. Some experimentation has been completed on using cement paste or mortar instead of epoxy however, the strength of each of these materials is significantly lower than epoxy (De Lorenzis & Teng, 2007).

23 2.6 Failure Mechanisms of FRP-Reinforced Structures

Intermediate crack (IC) debonding is considered the most important failure mechanism in an FRP strengthened structure subjected to flexure. IC debonding is the governing mechanism in determining the ultimate moment capacity of the strengthened section as well as the ductility. This failure mode typically occurs as a crack in the substrate crosses the FRP plate. Then additional cracks form between the FRP and substrate. When these cracks merge and propagate towards the end of the plate within the substrate as seen in

Figure 2-8, the strain in the FRP reduces. At this point, in which the FRP detaches from the structure, IC debonding has occurred (Seracino, Saifulnaz, & Oehlers, 2007). The load at which IC debonding occurs is typically denoted as PIC.

The Australian Masonry Code assumes that a horizontal crack will form at mid-height of an unreinforced masonry (URM) wall when subjected to a low load. As a result, the vertical bending moment capacity needs to be improved, as it is not adding to the ultimate strength of the wall after it cracks. FRP strips can then efficiently be used to strengthen the wall by increasing the vertical bending capacity, which in turn would result in a greater overall moment capacity of the wall (Willis et al., 2009).

24 NSM FRP

Intermediate Crack Out-of-Plane Interface Crack Loading Propagation

Masonry Wall

Figure 2-8. IC Debonding Mechanism (Seracino, Jones, Ali, Page, & Oehlers, 2007)

The increase in vertical bending capacity with the addition of a FRP strengthening system can be quantified by a pull-test. A pull-test simulates a load that is applied to a wall section centered about an FRP strip as seen in Figure 2-8. The pull-test can quantify the bond-slip behavior of the strengthening system, which is necessary in numerical simulations and the development of analytical models. The pull-test also provides the strain at which debonding occurs, the critical bond length, and IC debonding load. A typical pull-test specimen is shown in Figure 2-9. The debonding behavior observed in pull-tests performed on concrete specimens strengthened with FRP has been well researched, however analyzing the debonding behavior of masonry specimens is more complex in that masonry specimens are non-homogenous. The specimen is comprised of both masonry and mortar, which affects the bond behavior differently than in concrete (Willis et al., 2009).

Load P

FRP Strip

Restraint

12 3/4" (324 mm)

7 5/8" (194 mm)

3.5" (89 mm)

Figure 2-9. Typical Masonry Pull-Test Specimen

Failure mechanisms typically observed in pull-tests performed on NSM FRP-reinforced sections are debonding, FRP rupture, and pull out which is sometimes also referred to as sliding. Cracking in the substrate identifies the debonding failure mechanism as the concrete or masonry layer is the weakest part of the system. Herringbone shear cracks, which are at an angle to the FRP, propagate near or along the NSM FRP strip. As these cracks continue to propagate, the concrete or masonry can no longer transfer the stress from the FRP to the substrate thus the strip detaches from the structure. The increase in slip, in which the FRP is pulling out of the specimen, indicates that debonding is about to occur. Following debonding, a thin layer of substrate remains on the FRP strip. FRP rupture is the desirable failure mode as the FRP is used to its full capacity however the debonding strain is normally reached before the rupture strain of the FRP and therefore FRP rupture is not commonly

26 observed. The final failure mechanism is pull out failure in which the bond in the adhesive between the substrate and FRP fails. This failure should be avoided, as the FRP and concrete are not being used to their full capacities. However, this type of failure can be avoided by using an adhesive with ample strength, thinner layers of the epoxy in the groove, proper cleaning of the FRP and groove, and using an adequate bonded length (Seracino et al.,

2007a).

Willis et al. (2009) tested twenty-nine specimens that were constructed with either 10- hole clay brick units or solid clay brick pavers. The experimental study altered the following parameters for each pull-test:

 The strengthening technique of either NSM or EB

 The preparation of the surface with either a needle gun, sander, or grinder

 The geometric properties such as the width, thickness, and length of the FRP

 The location of the FRP in relation to the mortar joints and brick cores

 The bed joint material which was either mortar or quick drying paste

 The bonding method for the glass fiber sheets which was plate bonding or dry lay-up

 The FRP material of either carbon or glass

Each prism stood five bricks tall however they altered in width. The first set of prisms was stacked vertically similar to the prism St1.0 seen in Figure 2-10. The next set was constructed in a half overlap stretcher bond and was only one brick wide such as prism

HO1.0. The final set, similar to prism HO1.5 in Figure 2-10, was again in a half overlap stretcher bond however they were one and a half bricks wide (Willis et al., 2009).

Figure 2-10. Prism Specimens and FRP Locations (Willis et al., 2009)

The five failure modes, as previously described in this thesis, observed during testing of the masonry specimens were full debonding of the FRP, FRP rupture, FRP pull out

(sliding), FRP rupture and partial debonding, and FRP pull out and rupture. The most common failure mechanism was by debonding of the FRP. In this study, FRP rupture occurred as a result of the grips not being properly aligned. Therefore, higher stress concentrations were located on a percentage of the fibers. Rupture was not due to a satisfactory bond between the FRP and masonry specimen, which would indicate that the

FRP was being used to its fully capacity (Willis et al., 2009).

By and large, it was observed that the material in the bed joints played a small role in the failure mechanisms of the pull-tests as cracking occurred in the brick. Having a larger width of FRP strip, improved the ultimate load capacity of the test specimen although the

FRP can split if it extends into the cores as a result of the variation of stresses. The ultimate strength was decreased by 8.5% when the FRP strip was located through a vertical, perpendicular joint as a result of having less confinement from the surround brick. Locating the FRP strip between the cores of the brick also increased the strength slightly compared to

28 when the strip was located at the core, which was also a result of the increased confinement provided by surrounding bricks. For prisms that utilized the NSM technique, the strength and ductility were higher than when using the EB technique. It was concluded that this was a result of the FRP being better confined in a NSM application (Willis et al., 2009). This was also observed in FRP-to-concrete pull-tests in which the debonding crack was more confined with an increase in concrete area surrounding the FRP (Seracino et al., 2007a).

The pull-tests performed on these masonry specimens follow a similar load-slip relationship to pull-tests for concrete specimens (Willis et al., 2009). A bilinear bond-slip model best describes the shear transfer mechanism of the debonding crack interface. The axial force developed in the externally bonded FRP is transferred to the concrete through the adhesive layer in shear, as illustrated schematically in Figure 2-11. Both the slip and shear stress increase linear-elastically until a peak stress, τf, is reached, as seen in Figure 2-12. The slip, which is the relative displacement of the plate at the loaded end, is denoted as δ1 at this peak stress. In what is referred to as interfacial softening or microcracking, the shear stress begins to decrease linearly while the slip continues to increase. This continues until macrocracking or debonding of the plate is observed as the shear stress reaches zero at a slip of δf. The slip will continue to increase as it is being pulled out of the specimen (Yuan,

Teng, Seracino, Wu, & Yao, 2004).

Figure 2-11. Shear Stress in an EB Pull-Test (Yuan et al., 2004)

Figure 2-12. Bilinear Bond-Slip Model (Yuan et al., 2004)

30 Willis et al. (2009) developed Equation 2-1 to predict the IC debonding resistance load for a masonry pull test specimen based on a model developed by Seracino et al. (2007a) for both EB and NSM reinforcing techniques for FRP-reinforced concrete. The generic model presented by Seracino et al. (2007b) is as follows:

Equation 2-1

Where τf is the maximum interface shear stress, δf describes the maximum slip in the bond- slip model, Lper is the perimeter of the debonding failure plane as denoted by the dashed line in Figure 2-13, and Ep and Ap are the Elastic modulus and cross-sectional area of the FRP plate, respectively. The perimeter of the debonding failure plane is typically offset by 0.039 inches (1 mm) on all sides of the plate.

Masonry Surface bp Ap of EB Plate Ap of NSM Plate tp

df Lper Lper bf Masonry bp Failure plane EB plate df Element

tp Failure plane NSM plate bf

Figure 2-13. Debonding Failure Planes (Seracino et al., 2007b)

Substituting the bond-slip model for FRP-to-concrete, Equation 2-1 can be converted to:

Equation 2-2

Where:

for EB systems

for NSM systems

bp = width of the FRP strip, as depicted in Figure 2-13 for EB and NSM tp = thickness of the FRP strip, as depicted in Figure 2-13 for EB and NSM fc = cylinder compressive strength of the concrete

To alter the concrete model for use with masonry prisms, a relationship between the tensile strength of concrete, fct, and the compressive strength of the concrete was utilized:

Equation 2-3

Substituting the lateral modulus of rupture, fut, of the brick for the tensile strength of concrete, fct, Equation 2-2 then becomes:

Equation 2-4

When comparing the predicted IC debonding resistance from Equation 2-1, also known as the bond-slip model, to the “concrete” model shown in Equation 2-4, it is apparent that the bond-slip model is more accurate as it directly uses the bond-slip curve for that particular pull test. The “concrete” model is based on the material properties of the brick and a general relationship between the peak shear stress and slip at zero shear stress. Equation

2-1 is correct nevertheless, τfδf is based on a relationship that would need to be

32 experimentally determined by the pull-tests in the study. Therefore, it is recommended to use this expression when predicting the IC debonding load (Willis et al., 2009).

Petersen et al. (2009) continued the work initiated by Willis et al. (2009) through an experimental study that involved eighteen pull-tests on masonry prisms comprised of solid brick units without cores as seen in Figure 2-14 (A). In this study, the effect of the cores would be removed and only the influence of the FRP strip’s position to the head joints would be studied. The first series was constructed in the same orientation that Willis et al. (2009) used shown in Figure 2-10. This study also performed pull-tests on specimens in which the

FRP strip was orientated parallel to the bed joints similar to what would be found in a wall that was reinforced with horizontal FRP. This configuration can be seen in Figure 2-14 (B).

The masonry prisms in from Figure 2-14 (B) were also subjected to a compressive force on each side of the prism that was perpendicular to the bed joints (Petersen et al., 2009).

(A) (B)

Figure 2-14. Pull Test Specimens (Petersen et al., 2009)

Similar to the failure mechanism observed by Willis et al. (2009), the masonry prisms failed by debonding through the masonry units with cracks initiating at the loaded end of the prism and propagating downwards at a 45 degree angle to the FRP as seen in Figure 2-15 in

33 which selected cracks are highlighted. Failure occurred when the FRP strip had completely debonded from the prism. Upon further examination of the FRP strip, it was observed that the strip had a thin layer of brick still attached to it. For prisms in which the FRP strip was centered in the mortar joint, debonding occurred in the mortar first with very little cracking in the brick units. Occasionally, sliding was the dominant failure mechanism however, this was most likely due to the strain gauges on the FRP strip reducing the bonded surface area. It was also observed that the cracking that initiated at the loaded end of the specimen had also propagated across the top of the prism as seen in Figure 2-16 (A) and on the backside of the specimen as shown in Figure 2-16 (B) (Petersen et al., 2009).

Figure 2-15. Crack Pattern on Face of Specimen (Petersen et al., 2009)

(A) (B)

Figure 2-16. Cracking Through Specimen (Petersen et al., 2009)

The series in which the FRP was aligned parallel to the bed joints failed as a result of the strengthened middle row of bricks that was attached to the FRP pulling through the entire prism with no normal compression was applied. Cracking propagated only through the bed joints in this series. Once the middle section had separated from the prism, only the middle section remained loaded until the FRP detached from the prism. These prisms demonstrated a 31% reduction in strength compared to the stack bonded prisms in the first series. Diagonal cracking was observed in the middle row of bricks when the prism was subjected to normal compressive loading. However, this indicates that again the mortar was cracking through the mortar because only cracks were visible in the middle row of bricks. The maximum applied compressive force of 145 psi (1 MPa) that was applied to the prisms did produce a similar ultimate load to the stack bonded specimens from the first series (Petersen et al., 2009).

Petersen et al. (2009) concluded that when the FRP strip was located parallel to a mortar head joint, the ultimate load capacity of the specimen was decreased by 8% compared to the stack-bonded prism. That strength was further reduced by 11% when the FRP strip ran

35 directly through a mortar joint. When comparing the relative reduction in debonding resistances between test specimens, Petersen et al. (2009) found that the 1.5 brick wide specimen and the stacked prism in the study completed by Willis et al (2009) showed similar debonding resistances. This was unlike the results observed by Petersen et al. (2009) and the difference was concluded to be due to the cores in the bricks used by Willis et al (2009). It was determined that the cores had more of an influence on the bond strength than the position of the FRP to the mortar head joints. However, both Willis et al. (2009) and Petersen et al.

(2009) observed that the bond strength was reduced when the FRP passed through the mortar head joints. Therefore, it is unadvisable to position an NSM FRP strip through a mortar joint

(Petersen et al., 2009).

2.7 Conclusions

Much research has looked into strengthening URM walls with FRP. It has been determined that both the EB and NSM techniques contribute significant strength gains to the structure however, past studies have shown that using vertical NSM FRP strips is more efficient at strengthening structures for out-of-plane loading by increasing the vertical moment capacity. Further, NSM FRP has many other advantages such as less effort involved in preparing the structure for application and the ability to conserve the aesthetics of the structure by minimizing the visibility of the strengthening system. This is a particularly attractive quality in strengthening and repairing historical structures. For this research, it was decided to use NSM FRP in pull-tests in order to determine the change in debonding behaviour due to varying masonry pattern bond and environmental conditions. The

36 advantages of NSM FRP would likely cause its selection in most design applications, therefore, the durability of this system must also be determined as discussed previously in this chapter.

It is imperative to know the long-term characteristics of a strengthening system before putting it into service. Few studies have looked at the response of FRP-reinforced masonry structures to weathering and aging. Even fewer have studied the use of pull-tests to quantify the durability of the system. Thus, this research exposes specimens reinforced with NSM

FRP to various environments for differing time periods as a preliminary attempt to determine how the system would perform in in-service applications. The conditioning treatments included natural weathering and extreme conditioning in which the specimens will be subjected to extreme high and low temperatures. The specimens were tested at various intervals during a nine-month period to determine the difference in strength and behavior.

The subsequent chapter will discuss in further detail the experimental program created to explore the debonding behavior of FRP-reinforced masonry specimens when subjected to environmental conditioning. Specimen construction details as well as the set-up for the pull- tests will be described in detail.

37 CHAPTER 3 – EXPERIMENTAL PROGRAM

3.1 Introduction

To determine the interface load-slip behavior of the fiber-reinforced polymer (FRP) and masonry system due to varying exposures to environmental conditioning, FRP-to- masonry prisms were constructed and tested in the Constructed Facilities Lab on Centennial

Campus at North Carolina State University in Raleigh, North Carolina. Pull tests simulate the failures associated with the IC debonding mechanism that can occur in unreinforced masonry walls under out-of-plane bending. The following chapter will discuss the construction process and the testing procedures completed on all components of the prisms.

3.2 Pull-Test

3.2.1 Test Specimens

Each masonry prism was constructed using five courses of modular hollow core brick with nominal dimensions of 8” x 4” x 2 2/3” (203.2 x 101.6 x 67.7 mm) which corresponds to the length, width and height of the prism. The actual dimensions, including the spacing of the 1 ½” diameter cores are shown in Figure 3-1. In each series of prisms that were tested, the bond configuration varied between specimens in that they were in a stacked bond configuration or in a running bond configuration, which were one and a half bricks wide (see

Figure 3-2).

38 2" (51 mm) 2 1/4" (57 mm)

1.5" (38 mm) 3 1/2" (89 mm)

1 3/4" (44 mm) 2" (51 mm) 7 5/8" (194 mm)

Figure 3-1. Brick Dimensions

FRP

(A) (B)

Figure 3-2. FRP-to-Masonry Joint Prisms (A) Stacked Bond and (B) Running Bond

As seen in Figure 3-2, the FRP was positioned in the center of the front face of the masonry prism for the stacked bonded prisms. For the prisms in a running bond configuration, the FRP was placed slightly to the left or right of a vertical mortar joint depending on which orientation was closest to the center of the face for that particular prism.

The FRP was ½” (12.7 mm) in width, which prevented it from cutting into the hollow cores of the brick. On average, the actual height of each specimen was 12 ¾” (323.9 mm).

39 3.2.2 Masonry Prism Construction

On February 11, 2009, masons from Aztec Masonry of Garner, North Carolina, constructed the masonry prisms used for the pull-tests. The specimens were set on large pieces of plywood supported by wooden boards. This enabled the masons to have a flat working surface while constructing the prisms, which helped ensure that they would be level.

Mortar was placed by use of a trowel on the plywood before the first brick course was set to help initially level the brick prism. For the stacked bond configuration, each brick was set on the mortar joint that was approximately 3/8” (9.5 mm) thick as seen in Figure 3-3. Following the placement of each brick, the end of the trowel was used to tap the top of the brick to confirm that each brick course was level before placing the next brick course (see Figure

3-4). After all five bricks were placed, the prisms were confirmed to be level through use of a level. The one and half-wide specimens differed in that they were constructed by splitting one brick in half and placing the bricks in a running bond pattern. Mortar then filled in the vertical joints between the whole and the half brick. Upon completion of construction, a waterproof tarp was placed over the prisms to protect them from the elements while they were curing. After two days, during which the mortar initially set, the prisms were brought inside to protect them against any further exposure to wind or rain.

Figure 3-3. Placing a Brick Course

Figure 3-4. Leveling Brick Course

41 After a week, the mortar was cured and the prisms to be used for the pull-tests were ready to be cut. The high modulus, high tensile strength epoxy-carbon FRP was provided by

Fyfe Co., LLC and the manufacturer’s material properties are given in Table 3-1. The FRP came in 2-inch (50.8 mm) wide rolls and therefore needed to be cut to the desired length and width using a wet tile saw. The FRP was cut to a length of 24” (609.6 mm), which would provide enough length for the instrumentation that would be used in the pull-tests. From

Figure 3-2, it can be seen that the FRP does not penetrate through the hollow cores due to the

½” (12.7 mm) width of the FRP strip used. Cutting into the cores was avoided, as this would allow epoxy to seep into the core, thereby escaping the groove and potentially weakening the strengthening system. A high-quality bond might have been harder to achieve if the epoxy seeped out of the groove. The wet tile saw was then used to cut the 2-inch (50.8 mm) wide strips into ½-inch (12.7 mm) strips. Immediately after cutting of the FRP, the strips were cleaned with acetone to remove any excess carbon particles or debris on the FRP.

Table 3-1. Composite Gross Laminate Properties (Provided by Fyfe Co., LLC)

ASTM Typical Test Design Property Method Value Value Ultimate Tensile Strength in Primary Fiber 364,500 D-3039 405,000 (2.79) Direction, psi (GPa) (2.51) Elongation at Break D-3039 1.8% 1.67% 22.5 x 106 20.2 x 106 Tensile Modulus, psi (GPa) D-3039 (155) (139)

Because the FRP was first cut to a width of ½ inch (12.7 mm), a groove approximately 5/8 inch (15.9 mm) deep would be cut into the brick prism to allow the epoxy to encase the FRP on all sides. Two passes with a skill saw would be needed to make this

42 groove wide enough for a 0.079 inch (2.0 mm) thick strip of FRP and 0.039 inch (1 mm) of epoxy on each side of the FRP. A vice was made to hold the brick prism in place as the skill saw cut through the specimen. The vice made of 2” x 4” (50.8 x 101.6 mm) pieces of lumber was placed on concrete slabs that were in the yard of the CFL. For the stacked bond specimens, the groove would be placed in the center of the brick prism. For each prism, the centerline was marked based on that particular brick prism’s dimensions as seen in Figure

3-5. For the running bond prisms, the groove was also cut approximately down the center of the prism but occasionally this line had to be adjusted to avoid cutting through the vertical mortar joint that ran parallel to the groove. The brick dust seen in Figure 3-6 was created from cutting into the brick prisms was collected and saved for later use. Upon completion of cutting into the bricks, the prisms were brought inside and an air hose was used to push air through the grooves to remove any remaining brick dust.

Figure 3-5. Marking Groove Position

Figure 3-6. Creating Groove in Brick Prism

With the masonry prisms and the FRP cut, the final stage of construction could commence. Approximately half of the masonry prisms were laid down on a clean, flat surface that was covered with plastic sheets. Air was blown through the grooves to ensure the groove was thoroughly cleaned of any remaining brick dust or debris. Duct tape was placed along the sides of the groove to protect the surrounding brick from getting stray epoxy on it (see Figure 3-7). The FRP was once again cleaned with acetone to ensure it was completely free of dirt and debris.

Figure 3-7. Final Preparation of Brick Prisms for FRP

At this point, the two-part Tyfo TC (Tack Coat) Epoxy produced by Fyfe Co., LLC was mixed. The material properties given by the manufacturer are shown in Table 3-2. The epoxy was mixed according to weight in the ratio of 100 A: 25 B, according to manufacturer’s instructions. A long drill bit attached to a hand drill was used to mix the two parts of the epoxy for approximately 5 minutes. The epoxy was then scooped into empty caulking tubes using a spatula. The caulking tubes facilitated the filling of the grooves by allowing them to be filled more neatly and efficiently.

45 Table 3-2. Tyfo TC Epoxy Properties (Provided by Fyfe Co., LLC)

Property Value Elastic Modulus, ksi (GPa) 224.6 (1.55) Compressive Strength, psi (MPa) 4,088 (28) Tensile Strength, psi (MPa) 2,423 (17) Tensile Elongation, % 2.7

For each prism, the entire depth of the groove was filled with epoxy and a strip of

FRP was laid into each groove as seen in Figure 3-8. The FRP was set flush with the top face of the brick prism. Because the groove was filled with epoxy, the insertion of the FRP displaced the epoxy onto the face of the prism. Subsequently, a small plastic scraper was used to push the epoxy back into the grooves and to clean the epoxy off the face of the prism.

On a small number of the prisms, the brick dust, which was saved from when the prisms were cut, was sprinkled on top of the epoxy as depicted in Figure 3-9. A light weight was placed on the groove to hold the FRP in place in case the FRP was accidently shifted while curing. The brick dust was used to camouflage the FRP and epoxy system, which is apparent in Figure 3-10. This makes the NSM strengthening scheme an advantageous one when the preservation of the aesthetics of the structure is essential.

Figure 3-8. Filling Groove with Epoxy

(A) (B)

Figure 3-9. (A) Application of Brick Dust (B) Brick Dust Set on Prism

Figure 3-10. Epoxy Filled Groove

Before the first series of prisms were tested, tabs were cut from 15-gauge steel sheets.

The tabs were cut using a band saw to 4 inches (101.6 mm) in length and 1 ½ inches (38.1 mm) in width. The brick prisms were laid on a surface covered with plastic sheets to protect the surface from getting epoxy on it. Tabs were glued to opposing sides of the FRP strip using Fyfe Type S epoxy thickened with the product by Cabot, Treated Amorphous Fumed

Silica TS-720. This makes a thicker paste to make the application of the epoxy less difficult.

Small C-clamps were used to hold the tabs in place while the epoxy cured. The tabs were

48 adhered to the FRP so that the grips on the MTS Q-Test would bite into the steel tabs instead of into the FRP, which could possibly cause local damage.

An alternative method for constructing the tabs was used for the following series of test specimens. Glass FRP sheets were found to be an effective material for constructing tabs. Steel plates were stacked beneath the FRP strip so that the tabs would have a level surface to sit on. Glass FRP sheets of dimensions 4” x 6” (101.6 x 152.4 mm) were cut, in which half of the sheets had the FRP fibers going in the longitudinal direction and the other half, in the transverse direction. To prepare the specimens, the masonry prism was set on its smallest dimension so that the FRP strip was parallel with the level surface.

Fyfe Co., LLC, supplied the Type S epoxy that would be used to adhere the sheets to the FRP strip. After the epoxy was mixed to manufacturer’s instructions, three sheets of alternating FRP orientations were set on a sheet of plastic as seen in Figure 3-11. The FRP was then impregnated with epoxy. After one side of the sheet was completed, it was turned to its other side to make sure both sides held sufficient amounts of epoxy. The second sheet of opposite FRP orientation was set on top of the first and was smoothed down to decrease the likelihood of any air bubbles remaining between the sheets. This was then repeated for the third sheet. The FRP tab was then set on the steel plates and the brick prism was oriented so that the FRP strip sat in the center of the tab. The FRP strip was pushed down on to the sheets to encourage a complete bond to form there. Because the FRP strip was of such a large thickness, smaller strips that were 6” (152.4 mm) long but only 1 ¾ inch (44.45 mm) wide were cut also seen to the right of the larger FRP strips in Figure 3-11. These smaller strips were impregnated with the epoxy and then layered along the edges of the strip to avoid

49 creating a void when the top tab would be placed on the strip. Three strips were used on each side of the FRP strip. Finally, the process was repeated for the top tab, which was also made of three glass FRP sheets. A plastic sheet was then laid across the FRP tab and then two to three more steel plates were set on the strip to hold the tab in place as shown in Figure 3-12.

After the epoxy initially set after two days, the steel plates were removed. Excess epoxy that had seeped to the sides of the tab, causing the tab to no longer be of rectangular dimensions, was trimmed using tin snips.

Figure 3-11. Orientation of Fibers for Tabs

Figure 3-12. Construction of Tab

3.2.3 Experimental Set-Up

An MTS Q-Test was utilized to apply the tensile force to the FRP. The Grade 50 steel plate of dimensions 24" x 28" x 1" (609 x 711 x 25.4 mm) was bolted to the MTS Q-

Test to create a level surface for the brick prism to sit. Occasionally, hydro-stone was necessary to level the prism. At this point, it was necessary to align the FRP with the grips in order to minimize any eccentricities that could cause a bending force to be applied to the

FRP. This would result in an undesirable failure mechanism such as premature rupture of the

FRP.

Once the FRP was aligned in the direction that the force would be applied, a top restraint, also being a steel plate of the same grade and dimensions, was lowered onto the brick prism by being slid along four 1" (25.4 mm) diameter Grade 50 steel rods that were 36"

(914.4 mm) in length. Hydro-stone again was applied to the top of the brick prism to ensure that the prism would remain level while being compressed against the top steel plate. The

51 FRP was allowed to go through the top steel plate by a 5" x 3" (127 x 76.2 mm) hole that was cut into the plate. Once the masonry prism was level, the bolts were tightened so that it would remain restrained while the FRP was subjected to a tensile load (see Figure 3-13).

(4) Steel Rods E-Glass Fiber P Tab NSM CFRP Extensometer

Linear String Potentiometer

Steel Reaction Plate Hydro-stone Specimen

Figure 3-13. Pull-Test Set-up

3.2.4 Instrumentation

An extensometer and linear string potentiometer were used to measure the strain in the FRP and the slip between the FRP and the masonry prism near the loaded end, as shown in Figure 3-14. The extensometer was calibrated for use with the Q-Test and recorded the change in displacement in the FRP over a 2" (50.8 mm) gauge length. The linear string potentiometer was zeroed as it was set against a steel angle that was in line with the top of the masonry prism. The steel angle was hot-glued to an extension arm that was connected to the bottom of the masonry prism. This ensured that cracking in the top course would not reposition the linear string potentiometer. The slip would then be recorded relative to the top

52 of the brick prism. All data, including the FRP strain, applied load, and slip from the linear string potentiometer, was electronically recorded through the MTS program, Test Works 4.

Extensometer Tab

Linear String Potentiometer

Figure 3-14. Pull-Test Instrumentation

3.2.5 Testing Procedure

The MTS Q-Test loaded the FRP in tension under static loading conditions until failure was reached. The loading rate was .012 in/min (0.3 mm/min). Once the MTS Test

Works 4 software recognized failure of the specimen following the decrease of applied load, it automatically stopped the test. Test Works 4 was then able to export the displacement of the crosshead, extensometer displacement, the load, stress, and strain for each pull-test.

3.3 Testing Matrix

The masonry prisms were tested under varying environmental conditions and at different time intervals. Approximately every four months, a series of masonry prisms was

53 tested. Each series included seven prisms: three of the prisms being in a stacked configuration and four in a running bond configuration. Of the prisms in a stacked configuration, one prism was held in a controlled environment, one was exposed to natural environment conditions, and the third was placed in an environmental chamber. Two of the four prisms in the running bond configuration were controls, one was exposed to the natural environment, and the final prism was exposed to extreme environmental conditions. Prisms that were made for compression tests were subjected to the same environmental conditions.

The prisms that were considered controls were placed on shelves in the basement of the CFL. The basement’s temperature and humidity remained steady at an average of 72.7°

F (22.6° C) and 53.6%. It was assumed that there would be little to no damage from the environmental conditions. The prisms exposed to the natural environment were placed on the roof of a small shed in the yard of the CFL. The openings in a large grate on top of the roof was utilized in holding the prisms upright while maintaining at least an inch of separation between the prisms. Because the prisms were upright, rain was not able to pool on the prisms, which is similar to real configurations of masonry walls. The separation also allowed each prism to be exposed to the environmental conditions. The environmental chamber in the CFL provided the extreme environmental conditions for the prisms. The environmental chamber in the laboratory is able to reach temperatures ranging from -90° F to

185° F (-68° C to 85° C) and a range of relative humidity from 20% to 95%. Table 3-3 presents the test matrix that identifies the prisms found in each series as well as the nomenclature used in this experimental program.

54 Table 3-3. Test Matrix

Series Test Specimen Environment Specimen Compression Control C1:3-0 Initial CS1:3-0 Pull Control CR1:4-0 Control C-5 Compression Natural N-5 Extreme E-5 Control CS-5 After 5 Months Natural NS-5 Extreme ES-5 Pull Control CR1:3-5 Natural NR-5 Extreme ER-5 Control C-9 Compression Natural N-9 Extreme E-9 Control CS-9 After 9 Months Natural NS-9 Extreme ES-9 Pull Control CR1:2-9 Natural NR-9 Extreme ER-9

55 The nomenclature is described as:

C R 1 - 9

Specimen Environment Bond Pattern Age in Months Number

Where:

. The environment is denoted as:

o C – Controlled Environment (Prisms located in basement)

o N – Natural Environment (Prisms located outside)

o E – Extreme Environment (Prisms located in environmental chamber)

. The bond pattern is identified as:

o S – Stacked Pattern

o R – Running Bond Pattern

. If more than one prism with the same identifying features is tested, then a final

number is used to distinguish the prisms from each other.

. The age in months for each series corresponds to 0 months for prisms tested in April,

5 months for September, or 9 months for January tests.

3.4 Material Properties

3.4.1 FRP Coupon Testing

To test the ultimate strength of the FRP used in the pull-tests, coupons were cut from the same roll of FRP used in the pull-tests. Each coupon was 1/2" (12.7 mm) in width, which

56 is of the same width used in the pull-test, and approximately 14 inches (355.6 mm) in length, which was of sufficient length to minimize the likelihood for bending stresses caused by any eccentricities of the grips. The ASTM "Standard Test Method for Tensile Properties of

Polymer Matrix Composite Materials" with designation D 3039/D 3039M - 08 was followed in testing the FRP coupons. The standard suggests using emery cloth as tabs to protect the unidirectional FRP from crushing due to the pressure from the grips, therefore, medium course sandpaper of 100 grit was used as tabs. The sandpaper was cut to 1/2" (12.7 mm) in width and was folded over each end of the specimen with the course side facing the specimen.

Each specimen was centered and aligned vertically in the grips of the MTS Q-Test as seen in Figure 3-15. Pressure was increased to 1500 psi (10.3 MPa) for both the top and bottom grip to hold the FRP specimen. A 2-inch (50.8 mm) extensometer was employed in measuring the strain due to the tensile loading (see Figure 3-15). The load was increased at a rate of 0.05 in/ min (2 mm/min) until failure was reached. Rupture of the FRP is seen in

Figure 3-16. The load and corresponding extensometer displacement data was recorded at a rate of 2 Hz.

57 Sandpaper Tab

Extensometer

Figure 3-15. Typical Set-Up for FRP Coupon Testing

Figure 3-16. Typical Failure of FRP Coupon

58 The stress was found by dividing the load at a particular data point by the cross- sectional area of the FRP specimen and the ultimate stress was found by dividing the peak load by that area. The tensile strain is found by dividing the extensometer displacement by the gage length of the extensometer, which is 2 inches (50.8 mm). The stress-strain curves for the four test specimens are shown in Figure 3-17.

400,000 2,500 350,000

300,000 2,000 250,000 1,500

200,000 Stress, psi Stress, 150,000 1,000 MPa Stress,

100,000 500 50,000

0 0 0 0.005 0.01 0.015 0.02 Strain

Specimen 1 Specimen 2 Specimen 3 Specimen 4

Figure 3-17. Stress - Strain Curves for FRP Coupon Tensile Testing

It can be seen that all four test specimens have the same initial, linear slope. To calculate the elastic modulus, a trendline was fit to the linear portion of each curve. Table 3-

4 displays the ultimate tensile stress, elongation at break, and the tensile modulus calculated

59 for each FRP specimen. The manufacturer’s data is also supplied for reference. The mean, standard deviation, and coefficient of variation of the data are also provided.

Table 3-4. Tested FRP Coupon Properties

Ultimate Tensile Stress, Elongation at Tensile Modulus, psi Specimen psi (GPa) Break (GPa) Manufacturer’s 405,000 (2.79) 1.80% 22,500,000 (155) Data 1 304,013 (2.10) 1.49% 26,614,922 (184) 2 379,416 (2.62) 1.72% 27,615,803 (190) 3 261,223 (1.80) 1.05% 26,545,801 (183) 4 319,445 (2.20) 1.67% 26,998,327 (186) Mean 316,024 (2.18) 1.48% 26,943,713 (186) Standard 48,914 (0.34) 0.31% 490,281 (3.38) Deviation Coefficient of 0.155 0.207 0.018 Variation

3.4.2 Masonry

3.4.2.1 Brick Prism Compression Test

At the time that the masonry prisms used for the pull-tests were constructed, prisms necessary for compression tests were also made. The prisms utilized in the compression test were each seven brick courses high, which resulted in an approximate total height of 18 inches (457 mm). The masonry prisms were tested in accordance with ASTM C 1314 – 07 which is the “Standard Test Method for Compressive Strength of Masonry Prisms.” Four linear string potentiometers were adhered to the masonry prism in the locations shown in

Figure 3-18. These linear string potentiometers were the same as those used in the pull-test to measure slip. The string pot was zeroed against a wooden block or angle so that as it was

60 compressed during testing, the values recorded would represent the deformation in that area.

Wooden blocks were used to create a gage length for the string pots in two locations: across a mortar joint and across a brick course. Two FRP arms were glued to an angle to create a gage length across the middle five courses of brick. One arm was placed on each side of the masonry prism to verify that bending was minimized during testing.

P Upper Platen

Hydro-stone String Pot Across Prism String Pot String Pot Across Brick Across Mortar

Test Specimen

Hydro-stone

Lower Platen

Figure 3-18. Set-Up of Compression Tests

Using the MTS testing apparatus, the masonry prism was set on the lower platen. A bag of hydro-stone was set on the top of the masonry prism to create a level surface for the load to be applied. Figure 3-19 shows the back face of the masonry prism with the fourth string potentiometer and the bag of hydro-stone on top of the prism.

Figure 3-19. Back Face of Masonry Prism with Hydro-stone

The prism was initially loaded to half of the expected total load. Then it was loaded at a rate of 20 kips/minute (89 kN/minute) with data being recorded at a rate of 2 Hz.

Cracking typically initiated with a single crack down the center of the prism. The same crack was seen on both the front and back face of the masonry prism as highlighted for clarity purposes in Figure 3-20. At ultimate failure, the brick course has broken along the cores, exposing the mortar that had filled the cores (see Figure 3-21).

Figure 3-20. Vertical Cracks on Front and Back of Masonry Prism

Figure 3-21. Masonry Prism at Failure

63 The peak stress of each masonry prism was determined by dividing the peak load by the net area of the brick, which did not include the cores. Table 3-5 presents the peak loads as well as the elastic modulus recorded from the linear string potentiometers that spanned across five bricks and four mortar joints. The elastic modulus was determined from the slope of the linear portion of the curve between 5% and 33% of the peak stress as exemplified in the stress-strain curves for specimens at 5 months in Figure 3-22. However, when looking at the load versus the stroke plot, it was found that occasionally 5% of the peak stress was not on the linear portion of the load versus the stroke curve as seen in Figure 3-23. Therefore, the elastic modulus was adjusted to be determined from the beginning of the linear portion on the load versus stroke curve to 33%. The mean elastic modulus, found from averaging the elastic moduli on the front and back of the prism, was 1,050 ksi (7.24 GPa) with a standard deviation of 400 ksi (2.76 GPa) and coefficient of variation of 0.38. This average does not include prism N-9 as the coefficient of variation increased to 1.44 with its inclusion.

64 Table 3-5. Masonry Prism Compressive Strengths and Elastic Modulus

Peak Stress, psi E – Front, ksi E – Back, ksi Mean E , ksi Prism m m m (MPa) (GPa) (GPa) (GPa) C.1-0 2010 (14) 600 (4.13) N/A 600 (4.13) C.2-0 1433 (10) 690 (4.76) 1,046 (7.21) 868 (5.98) C.3-0 3145 (22) 484 (3.33) 1,284 (8.85) 884 (6.09) C-5 2858 (20) 1,201 (8.28) 1,425 (9.82) 1,313 (9.05) N-5 3690 (25) 1,279 (8.82) 1,165 (8.03) 1,222 (8.43) E-5 2067 (14) 1,096 (7.56) 825 (5.69) 960 (6.62) C.1-9 3881 (27) N/A N/A N/A C.2-9 2086 (14) 1,844(12.7) N/A 1,844(12.7) N-9 2905 (20) 9,650 (66.5) N/A 9,650 (66.5) E.1-9 2903 (20) 706,582 N/A 706,582 E.2-9 2624 (18) N/A N/A N/A

65 4,000 N-5 - Back 25 3,500 N-5 - Front

3,000 20 E-5 - 2,500 C-5 - Back Front E-5 - Back 15 2,000 C-5 - Stress, psi Stress, Front 1,500 10 MPa Stress,

1,000 5 500

0 0 0 0.001 0.002 0.003 0.004 0.005 0.006 0.007 Strain

Figure 3-22. Masonry Stress-Strain Curves at 5 Months

66 Stroke, mm 0 5 10 15 20 25 3,000 20

2,500 15 2,000

1,500 10 Stress, psi Stress, 1,000 MPa Stress, 5 500

0 0 0 0.2 0.4 0.6 0.8 1 Stroke, in

Figure 3-23. Load vs. Stroke for Prism C-5

With time, it is obvious from Figure 3-24 that the average compressive strength of the control prisms increases significantly prior to reaching a plateau. Figure 3-25 presents the peak stresses observed for the conditioned specimens compared to the average control specimen’s peak stress of the same series. For example, at 5 months, the peak stress recorded for the naturally conditioned specimen was 3690 psi (25 MPa) which is 1.3 times the peak stress of the control prism at that time interval. The peak stresses of the naturally conditioned specimens remain at least as strong as the control specimens. However, the extreme environmental condition had adverse effects on the prisms as these dropped in strength, particularly at the 5-month interval.

67 3500

3000 20 2500 15 2000

1500 10 Stress (psi) Stress

1000 (MPa) Stress 5 500

0 0 0 Months 5 Months 9 Months

Figure 3-24. Strength Increase in Control Specimens with Time

1.4

1.2

0.8 Natural 0.6 Extreme

0.4 Condtioned/Control 0.2

0 0 Months 5 Months 9 Months

Figure 3-25. Compressive Strength of Conditioned Specimen vs. Control

3.4.2.2 Brick Lateral Modulus of Rupture Test

Three courses of the clay brick used to construct the masonry prisms were adhered end-to-end using the Tyfo Type TC epoxy that was used to adhere the FRP to the brick

68 prism. The test specimen was set on a flat surface and held in place so that the bottom of the test specimen would be level. The specimens were tested in accordance to the

Australian/New Zealand Standard AS/NZS 4456.15:2003, "Masonry Units, Segmental

Pavers and Flags - Methods of Test - Method 15: Determining Lateral Modulus of Rupture."

Each test specimen was set on two support bars at a distance of 15 inches (381 mm) apart.

Two loading bars were centered on the middle brick course at a distance of 5 inches (127 mm) apart. A spreader beam was placed along the loading bars and the load was applied through the platens of the Forney testing apparatus as shown in Figure 3-26. The peak load and location of the break was recorded for each test. For those specimens that fractured in the middle course of brick, the crack typically occurred at the location of one of the cores due to the reduced cross-sectional area as shown in Figure 3-27 and Figure 3-28.

Spreader Beam

Test Specimen Glued Joint

L1 = 5" (12.7 cm)

Support Bar L = 15" (38.1 cm)

Figure 3-26. Lateral Modulus of Rupture Test Set-Up

Figure 3-27. Typical Break in Center Brick Course

Figure 3-28. Typical Break through Thinnest Section of Brick

70 Only the tests in which the break occurred in the middle brick was used in calculating the lateral modulus of rupture. The following equation was used to determine the lateral modulus of rupture shown in Table 3-6:

Equation 3-1

Where Tn = Lateral modulus of rupture (MPa)

M = Bending Moment at Failure (N-mm)

= 1000 ( )

W = Load when specimen fails (kN) l = Distance between support bars (in mm), which is 15” (381 mm) l1 = Distance between loading bars (in mm), which is 5” (127 mm)

Z = Section modulus of test specimen (mm3)

= for solid and cored units

B = work size width of each specimen (in mm), which is 2.25” (57 mm)

D = work size depth of each specimen (in mm) which is 3.625” (92 mm)

71 Table 3-6. Lateral Modulus of Rupture Results

Test Location Load, lbs (kN) Tn, psi (MPa) 1 Center 989 (4.4) 520 (3.582) 2 Edge Brick 1950 (8.7) - 3 Center 848 (3.8) 440 (3.014) 4 Joint 1413 (6.3) - 5 Joint 1555 (6.9) - 6 Center 1300 (4.547) 7 Center 1555 (5.439) 8 Center 933 (3.264) 9 Center 848 (2.966)

The mean lateral modulus of rupture for prisms that ruptured in the center course was

551 psi (3.802 MPa). The standard deviation was 143 psi (0.989 MPa) and the coefficient of variation was 0.260.

3.4.2.3 Compressive Strength of Brick Unit

In accordance with ASTM C 67-09, “Standard Test Methods for Sampling and

Testing Brick and Structural Clay Tile,” seven half brick specimens were subjected to loading in order to measure the compressive strength of the brick unit. Each brick was placed in an oven for twenty-four hours at 230° F (110° C). For testing, each brick was cut in half to produce half brick units. Prior to the test, hydro-stone was placed on the top and bottom faces of the brick to ensure an even bearing surface as seen in Figure 3-29. Load was applied through the Forney testing apparatus (Figure 3-30) at a steady rate up to one-half of the total load. Following reaching one-half of the peak load, the remainder of the load was applied at a rate so that failure of the specimen would occur in one to two minutes. Failure of

72 the specimen is shown in Figure 3-31. Stress was then calculated by dividing the peak load by the net surface area of the brick after subtracting the area of 1.5 cores.

Table 3-7 presents the dimensions and peak load for each half-brick unit. The mean compressive stress for the half brick units was 10,849 psi (75 MPa). The standard deviation was 1,823 psi (12.6 MPa) and the coefficient of variation was 0.168.

Figure 3-29. Half-Brick Test Specimen

73 P Upper Platen 2.25" (5.72 cm)

Test Specimen

3.5" (8.89 cm) Lower Platen

Figure 3-30. Brick Compression Test Set-Up

Figure 3-31. Failure of Specimen

74 Table 3-7. Compressive Strength of Brick Units

Length, in Width, in Depth, in Load, kips Stress, psi Test (mm) (mm) (mm) (kN) (MPa) 1 3.875 (98) 3.625 (92) 2.25 (57) 122.3 (548) 10,735 (74) 2 3.75 (95) 3.4375 (87) 2.25 (57) 126.8 (568) 12,384 (85) 3 3.625 (92) 3.5 (89) 2.25 (57) 110.9 (497) 11,054 (76) 4 3.75 (95) 3.5 (89) 2.25 (57) 107.0 (479) 10,214 (70) 5 4.5 (114) 3.5 (89) 2.25 (57) 102.9 (461) 7,855 (54) 6 3.5 (89) 3.4375 (87) 2.25 (57) 127.6 (572) 13,603 (94) 7 3.875 (98) 3.5 (89) 2.25 (57) 110.2 (494) 10,095 (70)

3.4.2.4 Mortar Cube Test

Each mortar cube was cast from the same mortar found in the masonry prisms used for the pull-tests. On average, the mortar cubes were 2" x 2" x 2" (50.8 x 50.8 x 50.8 mm).

The mortar cubes were tested according to ASTM C 109/C 109M - 08, "Standard Test

Method for Compressive Strength of Hydraulic Cement Mortars (Using 2 in or [50-mm]

Cube Specimens)." The cubes were loaded at an approximate rate of 200 - 400 lb/s (900 -

1800 N/s) using the Forney testing apparatus as depicted in Figure 3-32. The mode of failure and peak compressive load was recorded for each specimen. Table 3-8 identifies the peak load for each compressive test completed on September 4, 2009 at an age of 6 months and 24 days (205 days).

Mortar cube 5 was unable to produce a compressive strength due to an error in the data acquisition from the Forney testing apparatus. The average compressive stress calculated for the remaining five cubes was 2,450 psi (16.9 MPa). The standard deviation

75 was 363 psi (2.5 MPa) and the coefficient of variation was 0.148. The typical conical failure shape for each of the mortar cubes is seen in Figure 3-33.

Table 3-8. Compressive Strength of Mortar Cubes

Cube Length Width Height Max Load, lb Comp. Strength, psi No. (in) (in) (in) (kN) (MPa) 1 2 2 2 8,680 (39) 2,170 (15) 2 2 2 1.9375 10,461 (47) 2,615 (18) 3 2 1.9375 2 11,677 (52) 3,013 (21) 4 2 1.9375 2 8,651 (38) 2,233 (15) 5 2 1.9375 2 N/A N/A 6 2 2 2 8,849 (39) 2,212 (15)

P Upper Platen 2" (5.08 cm)

Test Specimen

2" (5.08 cm) Lower Platen

Figure 3-32. Mortar Cube Test Set-Up

Figure 3-33. Typical Failure Mode of Mortar Cube

The following chapter will present the test results from pull-test completed experimentally over the course of a nine-month period. General behavior of the pull-tests will be discussed. The material properties determined experimentally in this chapter will be necessary for performing a numerical analysis on the pull-test specimens in Chapter 5.

77 CHAPTER 4 – TEST RESULTS

4.1 Introduction

The following chapter presents the experimental results of 24 pull-tests tested in three series over a nine-month period. The first series of tests were performed in April of 2009 and were followed by the second series of tests in September. The third and final series of testing was completed in January of 2010. The general behavior observed during testing of the FRP- reinforced masonry specimens will be discussed as well as the differing results between conditioning groups and construction patterns. Construction patterns of the specimens varied between stacked or running bond specimens as described in the previous chapter. The specimens were subjected to one of three conditioning groups that varied between a natural, extreme, or a control environment. Results recorded from these pull-tests will be compared to a published model that predicts the bond resistance of a NSM FRP-reinforced masonry prism in a pull-test.

Each prism in the following chapter is identified by the nomenclature first described in Chapter 3, in which the first letter describes the environment and is denoted as either C for the control, N for natural, or E for extreme environmental conditioning. The bond pattern is identified next as being S for stacked or R for the running bond pattern. If more than one prism with the same identifying features is tested in a series, then a final number is used to distinguish the prisms from each other. The age in months for each series corresponds to 0 months for prisms tested in April, 5 months for September, or 9 months for tests completed

78 in January. For example, prism CR1-9 describes the first control running bonded prism tested during the month of January.

4.2 Pull-Test Results

4.2.1 First Series (Control)

The first series describes the set of masonry prisms tested during April of 2009.

Construction of the prisms was complete on March 12, 2009 therefore the epoxy used in the

FRP strengthening system was allowed to cure for a full month prior to testing. All prisms for each series were tested at room temperature. Further, all specimens in each series were allowed to air dry at room temperature for a minimum of seven days prior to testing. In the initial series, seven masonry prisms were tested including three prisms constructed in a stack- bonded configuration and an additional four prisms constructed in a running bond pattern.

Aluminum tabs were used in this series to prevent the grips on the MTS Q-Test from prematurely rupturing the FRP.

Debonding of the FRP from the masonry specimen followed similar patterns for each specimen, regardless of whether the prism was constructed in a stacked or running bond.

Cracking initiated in the top course of brick and propagated along the groove cut for the FRP strip as indicated in Figure 4-1. Cracking then continued down the face of the top brick course while also propagating across the top course of brick as seen in Figure 4-2. The crack would propagate towards the nearest core and along its edge towards the back face of the prism. For prisms constructed in a running bond, the vertical mortar joints typically cracked along one edge of the joint in a clean, straight line as highlighted in Figure 4-3. Cracking

79 then continued to propagate along the face of the masonry prism until reaching the bottom brick courses at which point the FRP completely debonded from the specimen as seen in

Figure 4-4. Upon inspection of the back of the prism, it was also observed that cracking propagated from the top course of the prism downward in a similar manner to the front face of the prism. However, cracking in the top course on the front face of the prism did not coincide in time with cracking in the top course of the prism on the back face. Cracking on the back face of the prism typically trailed cracking on the front face.

When the FRP strip was removed from the prism, it was clear that the masonry prism had split through both sides. This became more obvious as the top steel plate was lifted off the specimen, and light could be seen through the prism as seen in Figure 4-5. The split was also evident in Figure 4-6 as the crack had propagated across the top of the brick.

Figure 4-1. Crack Propagation Across Top of Brick

Figure 4-2. Crack Propagation on Front Face of Masonry Prism

Cracking along Mortar Joint

Figure 4-3. Continuation of Crack Propagation

Figure 4-4. Complete Debonding of FRP

Figure 4-5. Splitting of Masonry Prism

Figure 4-6. Splitting Across Top Brick Course

FRP strips removed from the masonry prisms were similar in that a thin layer of brick was adhered to the strip as seen in Figure 4-7. Typically, a triangular section of brick was attached to the top of the FRP strip. It was also common to see the clean break of the vertical mortar joint remaining on the FRP strip for prisms constructed in a running bond.

Occasionally, small areas of the FRP strip were not covered in brick, which was evidence of a poor bond between the adhesive and the brick.

Figure 4-7. FRP Strip Removed from Specimen

Table 4-1 presents the results from the first series of pull-tests. The maximum slip identifies the final reading recorded during testing. Generally, specimens failed suddenly causing the final slip measurement to be taken when the load had decreased to almost zero.

Therefore, load-slip curves for most specimens have straight unloading curves post-peak due to the existence of the one final slip recording near zero load. The average debonding resistance of the stacked and running bonded specimens was 5,887 lbs (26.18 kN) and 5,723

(25.47 kN) respectively, which is only a 2% difference. The standard deviation for the stack- bonded specimens was 1,455 lbs (6.47 kN) with a coefficient of variation of 0.25 and the standard deviation for the running bond specimens was 1,156 lbs (5.14 kN) with a coefficient of variation of 0.20 which is within the acceptable limits of 0.3 for masonry (Daniels, Duffy,

Johnsson, & Sedmak, 2009).

84 Table 4-1. Summary of Test Results for Series 1

Specimen Peak Load, lbs (kN) Slip at Peak Load, in (mm) Maximum Slip, in (mm) CS1-0 7559 (33.62) 0.06 (1.46) 0.51 (13.0) CS2-0* 5193 (23.10) N/A N/A CS3-0* 4908 (21.83) N/A N/A CR1-0 6004 (26.71) 0.02 (0.60) 0.15 (3.89) CR2-0 6746 (30.01) 0.00 (0.06) 0.00 (0.11) CR3-0* 4066 (18.09) N/A N/A CR4-0 6091 (27.09) 0.00 (0.04) 0.12 (3.11) * Slip was not recorded due to failure of the E-glass fiber tab prior to debonding of the FRP.

Figure 4-8 plots the load-slip curves for specimens tested in the first series in which the slip data was recorded. Therefore, specimens CS2-0, CS3-0, and CR3-0 are not included in the figure. The aluminum tab of CS1-0 detached during testing thus a new tab was attached to the FRP and the test was restarted. The value of slip for the second test was set to match the slip recorded prior to loss of the tab. The new tab also detached after reaching the peak load, therefore the descending branch of the load-slip curve was not recorded as the test was ceased. Upon comparison of the stack-bonded and running bond specimens, it appears that the neither group of specimens exhibited large amounts of slip before reaching the peak load.

85 Slip, mm 0 1 2 3 4 8,000 35 7,000 CS1-0 30 6,000 25 5,000 20 4,000

15 Load, kN Load, Load, lbs Load, 3,000 CS2-0 CR1-0 10 2,000

1,000 CS4-0 5

0 0 0 0.05 0.1 0.15 Slip, in

Figure 4-8. Load-Slip Curves for Specimens in First Series

4.2.2 Second Series (5 Months Conditioning)

The second series of specimens was tested during September 2009. Upon visual inspection of the prisms prior to testing, it was clear that the exposed epoxy on specimens subjected to natural conditioning was more yellow in color than the other specimens. A small crack along the groove for the FRP was observed in the top of brick and another crack was located in the third brick course of specimen, NS-5. The top course of brick on specimen ER-5 was not sufficiently adhered to the rest of the prism, as it would slightly shift when moved. No other major cracks were observed for prisms from the different conditioning groups.

86 Generally, failure of the specimens was by FRP debonding. Cracking initiated in a similar manner to prisms tested in the first (control) series in that cracking propagated across the top of the first brick course and then down the face of the prism. Cracks were always along or at an angle to the FRP strip as seen in Figure 4-9. Cracking was also observed along the vertical mortar joints and on the back face of the prisms as depicted in Figure 4-10. Upon further inspection of the FRP strips removed from the specimens, it was observed that almost all had a thin layer of brick and mortar still attached to it, as was previously observed in the first series. However, for specimens subjected to natural conditioning such as NS-5 and NR-

5, only small sections of the strip still had a layer of brick attached as seen in Figure 4-11.

The rest of the FRP was covered in a layer of epoxy that was more of a rubber consistency rather than the typical glass-like texture. Visual inspection of specimen NR-5, which behaved similarly to NS-5, showed that very little cracking in comparison to the control specimens was present on the face of the prism as seen in Figure 4-12. This was unlike the cracking observed for specimen CR1-5 (Figure 4-9). Minimal cracking was also present on the inside of the groove as it was smooth in most areas. Cracking along the height of the prism was present on the back of specimens NS-5 and NR-5.

Figure 4-9. Crack Propagation in CR1-5

Figure 4-10. Cracking on Back Face of CR1-5

88 (A) (B)

Figure 4-11. FRP Strip Removed from (A) Control (B) NR-5

Figure 4-12. Specimen NR-5 with FRP Removed (Front Face)

89 Table 4-2 shows the results from the set of prisms tested during the second series in

September of 2009. The stack-bonded specimen, CS-5, had a 20% higher debonding resistance than the average control running bonded specimen, which had an average value of

5,388 lbs (24.0 kN). The standard deviation and coefficient of variation was calculated as

1,034 lbs (4.60 kN) and 0.19 for the running bond specimens. However, both running bond specimens from the natural and extreme conditioning groups had greater debonding resistances than their stack-bonded counterparts. Upon isolating the stack-bonded specimens, it is clear that the specimen exposed to the natural conditioning was the weakest of the group whereas the control was the strongest. However, running bonded specimens from both the natural and extreme environmental groups were stronger than the control specimens. Compared to the first series, the control stack-bonded specimens on average, increased in debonding resistance by 15% although the running bond specimens decreased by

6%.

Table 4-2. Summary of Test Results for Series 2

Specimen Peak Load, lbs (kN) Slip at Peak Load, in (mm) Maximum Slip, in (mm) CS1-5* 6778 (30.15) N/A N/A CR1-5* 6123 (27.23) N/A N/A CR2-5 4206 (18.71) 0.03 (0.80) 0.44 (11.2) CR3-5 5836 (25.96) 0.02 (0.53) 0.23 (5.86) NS-5 5011 (22.29) 0.05 (1.16) 0.17 (4.30) NR-5 6357 (28.28) 0.04 (0.95) 0.43 (10.9) ES-5 5580 (24.82) 0.04 (1.02) 0.27 (6.93) ER-5 6092 (27.10) 0.00 (0.00) 0.18 (4.64) * Slip not recorded due to malfunction of linear string potentiometer.

90 The load-slip curves for stack-bonded specimens are shown in Figure 4-13 whereas

Figure 4-14 shows the load-slip curves for running-bonded specimens. Due to a malfunctioning linear string potentiometer, the slip was unable to be recorded for the control stack-bonded specimen. For comparison, the load capacity of CS-5 is presented in Figure

4-13. However, upon comparison of the natural and environmentally conditioned specimens, both curves are similar in that there is a primary peak followed by a slight dip in load. The slip then continued to increase before reaching the ultimate load. Similar to the control specimens, the unloading portion of the curve is straight due to the sudden failure of the specimen. The final reading of the slip was recorded immediately after the specimen failed and the load returned to almost zero.

Slip, mm 0 2 4 6 7,000 30

6,000 CS-5 25 5,000 20 4,000 15

3,000 ES-5 Load, lbs Load, 10 kN Load, 2,000

1,000 NS-5 5

0 0 0 0.1 0.2 0.3 Slip, in

Figure 4-13. Load-Slip Curves for Stack-Bonded Specimens in Series 2

91 Slip, mm 0 2 4 6 8 10 7,000 30 6,000 25 5,000 CR3-5 20 4,000 15

3,000 Load, lbs Load, NR-5 kN Load, 10 2,000

1,000 CR2-5 5 ER-5 0 0 0 0.1 0.2 0.3 0.4 Slip, in

Figure 4-14. Load-Slip Curves for Running Bonded Specimens in Series 2

4.2.3 Third Series (9 Months Conditioning)

The final series of specimens was tested in January 2010. Figure 4-15 depicts three specimens from the three environmental conditioning groups prior to testing. The specimen on the far left is from the control environment, the center is from the environmental chamber, and the farthest right specimen was exposed to natural conditioning. Prior to testing, the epoxy visible on the surface of specimens exposed to the natural conditioning was more yellow in color than specimens from the control or extreme environments. Other than the color of the epoxy, there were very few visible differences such as cracking between specimens of the various conditioning groups.

92 Control

Extreme Natural

Figure 4-15. Color of Epoxy Following Conditioning

Specimens exhibited the same failure modes as observed in previous series. Failure of the specimen was by debonding of the FRP. In most instances, the masonry prism was split during testing. Specimens from the naturally conditioned group exhibited a different failure mode and much higher debonding resistances. Diagonal cracking towards the FRP strip was not observed in the naturally conditioned specimens as it was in other conditioning treatments. Cracking propagated down the edge of the FRP groove as seen in Figure 4-16.

Upon removal of the FRP strip, it was apparent that shear failure occurred within the epoxy.

As depicted in Figure 4-17, little to no brick was observed on the FRP strip indicating that the epoxy was not bonded to the brick. Further, upon inspection of the groove, no epoxy was observed to be attached to the brick as shown in Figure 4-18. The brick layer was also not cracked further supporting that the epoxy had failed in shear.

Figure 4-16. Cracking on Masonry Prism (Natural Conditioning)

Figure 4-17. FRP Removed from Specimen (Natural Conditioning)

Figure 4-18. Groove in Specimen (Natural Conditioning)

Table 4-3 presents the test results recorded during the third series of tests. Both stack-bonded and running bonded specimens are consistent with the exception of CR2-9 having a slightly higher debonding resistance. The average debonding resistance of the stack-bonded and running bonded control specimens was 7,383 lbs (32.8 kN) and 7,949 lbs

(35.4 kN), respectively. The standard deviation for the stack-bonded specimens was 81.3 lbs

(0.36 kN) with a coefficient of variation of 0.01. The running bonded specimens’ standard deviation was 531 lbs (2.36 kN) with a coefficient of variation of 0.07. When comparing specimens from the natural conditioned group, it is clear that these specimens have significantly higher debonding resistances than the controls for both construction patterns.

Furthermore, the stack-bonded specimen is only 7% stronger than the running bonded pattern. The weakest specimens were those that were subjected to the extreme temperatures in the environmental chamber. The running bonded specimen was significantly stronger than

95 the stack-bonded specimen. The average debonding resistance of the stack-bonded specimen subjected to extreme conditioning was 4,377 lbs (19.5 kN). When comparing these specimens to the controls, on average, the stack-bonded specimens were 40% weaker than the control stack-bonded specimens. The running bonded specimens were only 12% weaker than the control running bonded specimens.

Table 4-3. Summary of Test Results for Series 3

Specimen Peak Load, lbs (kN) Slip at Peak Load, in (mm) Maximum Slip, in (mm) CS1-9 7440 (33.09) 0.02 (0.60) 0.51 (12.9) CS2-9 7325 (32.58) 0.05 (1.15) 0.33 (8.33) CR1-9 7573 (33.68) 0.01 (0.37) 0.48 (12.2) CR2-9 8324 (37.02) 0.02 (0.51) 0.39 (10.0) NS-9 9240 (41.10) 0.07 (1.86) 0.47 (12.1) NR-9 8594 (38.22) 0.07 (1.88) 0.23 (5.96) ES1-9 4144 (18.43) 0.00 (0.11) 0.26 (6.72) ES2-9 4610 (20.51) 0.01 (0.20) 0.10 (2.52) ER-9 6939 (30.87) 0.05 (1.23) 0.36 (9.21)

Figure 4-19 shows the load-slip curves for specimens constructed in a stack-bonded pattern while Figure 4-20 depicts the load-slip curves for running bonded specimens.

Isolating the curves for the stack-bonded specimens, it is apparent that the slip for the naturally conditioned specimens is greater at the peak load than other specimens. This is also observed in the running bonded specimens. Furthermore, specimens tested from the environmental chamber showed very small amounts of slip before reaching the peak load.

The E-glass fiber tab detached during testing for specimen ER-9 thus the load dropped as the testing apparatus was no longer resisting load as it pulled the tab off the specimen. A new

96 tab was attached to the FRP and the test was restarted. The final slip recorded in the first test was set as the initial slip for the second test as the FRP had displaced from the prism. Figure

4-20 presents the combined curve that includes both trials of the test.

Slip, mm 0 2 4 6 8 10 12 10,000 9,000 40 NS-9 8,000 35 7,000 30 6,000 25 5,000 CS2-9

20 Load, lbs Load, 4,000 ES1-9 kN Load, 15 3,000 CS1-9 2,000 10 1,000 5 ES2-9 0 0 0 0.1 0.2 0.3 0.4 0.5 Slip, in

Figure 4-19. Load-Slip Curves for Stack-Bonded Specimens in Series 3

97 Slip, mm 0 2 4 6 8 10 12 10,000 9,000 40 NR-9 8,000 35 7,000 30 6,000 CR1-9 25 5,000

20 Load, lbs Load, 4,000 kN Load, 15 3,000 ER-9 CR2-9 2,000 10 1,000 5 0 0 0 0.1 0.2 0.3 0.4 0.5 Slip, in

Figure 4-20. Load-Slip Curves for Running Bonded Specimens in Series 3

4.3 Discussion of Results

4.3.1 Effects from Environmental Conditioning and Construction Pattern

From Figure 4-21, it is clear that with the exception of the running bond specimens from the second series, all control specimens increased in bond resistance. From the first to second, and then the second to third series, the stack-bonded control specimens increased in strength by 15% and 10% respectively. By the final series, the stack-bonded specimens had increased in strength by 25% from the initial series. Alternatively, the running bond

98 specimens tested in the second series had decreased in strength by 6% on average. Between the second and the third series, the running bond specimens then increased by 48%, which resulted in a 39% increase in strength from the initial series. When comparing just the stack- bonded construction to the running bond, the stack-bonded specimens for the first two series are 3% and 21% stronger than the running bond counterparts. If specimen CR2-5, which had a debonding resistance of 4,206 lbs (18.7 kN), is not included, then the stack-bonded specimen has only a 12% increase in debonding resistance. With masonry structures, variation should be expected due to the nature of the materials and the prism construction.

Acceptable coefficients of variation for masonry tests are as high as 30% (Daniels et al.,

2009). In this particular case, a large piece of mortar failed in the corner of the prism due to poor construction of the prism. By the final series, the stack-bonded specimen was 8% weaker than the running bonded specimen.

99 9,000 40

8,000 Running 35 Stacked 7,000 Stacked 30 Stacked 6,000 25 Running 5,000 Running 20

4,000

Load, lbs Load, Load, kN Load, 15 3,000 10 2,000

1,000 5

0 0 0 Months 5 Months 9 Months

Figure 4-21. Peak Loads of Control Specimens with Time

Figure 4-22 depicts the change in debonding resistance over time for specimens exposed to environmental conditioning. This figure displays the debonding resistance, P, of the conditioned specimens to the debonding resistance of the control specimens, identified as

Po, in the same series. Clearly, the bond resistance for specimens subjected to natural conditioning increased between the second and third series however the specimens tested in the third series are also associated with a complete change in the failure mode. Specifically, by the final series the stack-bonded specimen increased in strength by 84% while the running bonded specimen increased by 35%. When comparing the construction patterns of each specimen, the stack-bonded specimen in the second series is 27% weaker than the running

100 bonded specimen in the same series. However, the stack-bonded specimen in the third series is 7% stronger than the running bond specimen.

The change in peak loads for specimens that were subjected to extreme conditioning is also presented in Figure 4-22. The running bonded specimens increased by 14% in strength over all three series, however, the stack-bonded specimen actually decreased in strength by 22%. In the second and third series, the running bonded specimens were 9% and

59% stronger, respectively, than their stack-bonded counterparts.

Isolating the stack-bonded specimens in the second series, it is apparent that both the natural and extreme conditioned specimens decreased in strength by 26 and 18%, respectively, compared to the control in the same series. However, in the third series, only the extreme conditioned specimen dropped in strength by 41% whereas the naturally conditioned specimen significantly increased in strength by 25%.

The running bond specimens behaved slightly different. Both naturally and extreme conditioned specimens in the second series exhibited higher strengths than their control counterparts. The naturally conditioned specimen was 18% stronger and the extreme conditioned specimen was 13% stronger than the control from the same series. The running bond specimens from the third series did behave similarly to the stack-bonded specimens in that the naturally conditioned specimen was stronger than the control by 8% yet the extreme conditioned specimen was 13% weaker than the control. Therefore, the percent increase or decrease for each of the specimens was not as extreme as the stack-bonded specimens.

101 1.8

1.6 Natural Running

1.4 Natural Stacked 1.2 Extreme Running Natural Running 1 Extreme Running P/Po 0.8 Extreme Stacked Natural Stacked 0.6 Extreme Stacked

0.4

0.2

0 0 Months 5 Months 9 Months

Figure 4-22. Debonding Resistance of Conditioned Specimens with Time

4.3.2 Discussion

It is apparent from the results presented in the previous section that the naturally conditioned specimens increased in strength by the third series following a slight dip in strength in the second. Upon observation of FRP strips removed from the naturally conditioned specimens in the second and third series, it is clear that the bond between the epoxy and the masonry deteriorated with time. A thin layer of brick is still visible on the

FRP strip in the second series as depicted in Figure 4-23 (B) however, almost no brick was observed on the FRP removed from the final series as seen in Figure 4-23 (C). The difference is clearly evident in comparing the FRP removed from a control specimen shown

102 in Figure 4-23 (A) to the FRP in the naturally conditioned specimen in the final series. The naturally conditioned specimens were exposed to the natural elements that not only included daily variations in temperature but also in precipitation. Specimens from the extreme conditioning group were not exposed to moisture or extremely high temperature for extended periods of time, which would potentially alter the material properties of the epoxy. Further, these specimens were exposed to less variation in temperature as not many cycles were run in the environmental chamber compared to the exposure from the daily variations for the naturally conditioned specimens. However, it was suspected that high temperature fluctuations would cause thermal cracking due to the difference in thermal coefficients of expansion for each material. However, upon inspection of the specimens removed from the environmental chamber, there were no observable cracks due to thermal expansion.

Generally, cracking observed on the specimens was from construction.

103 (A) (B) (C)

Figure 4-23. FRP Removed from Natural Conditioning (A) At 0 Months (B) At 5 Months (C) At 9 Months

One reason for the deterioration of the bond could be attributed to the naturally conditioned specimens’ exposure to precipitation and temperature. It is clear from Figure

4-23 that this deterioration continued with time. Figure 4-24, which presents the daily maximum and minimum temperatures experienced by the naturally conditioned specimens, shows a significant drop in temperature between the second and third series (History, 2010).

The average maximum and minimum temperature measured from March 12 to

August 21, 2009, which corresponds to the time that the specimens were exposed to natural conditioning before the second series of testing, was 83°F (28°C) and 62°F (17°C) respectively. The maximum and minimum average temperatures dropped to 65°F (18°C) and

46°F (8°C) between the second and the third series. As seen in the study by Bati and

Rotunno (2001), exposing FRP-reinforced prisms to numerous freeze-thaw cycles resulted in

104 a complete loss of the FRP strengthening system. Furthermore, the amount of precipitation that the specimens saw also increased between the second and third series as seen in Figure

4-25. Between the first and the second series, a total of 13.8 inches (351 mm) of precipitation fell however, the amount of precipitation between the second and the third series increased to a total of 18.3 inches (465 mm). As FRP systems are exposed to moisture, the resins expand or distort the reinforcing system. The strength and stiffness of the system is then compromised (ACI 440.2R-08, 2008). The decrease and increase in temperatures were not so severe as to label them as freeze-thaw cycles therefore, the deterioration of the bond for the naturally conditioned specimens can be associated with the increase of moisture.

Additionally, the naturally conditioned specimens were located on the roof of a one- story structure where temperatures would tremendously exceed the air temperatures recorded for that day. Typically, for dark colored roofs, the roof temperature is 90 °F (50 °C) higher than the air temperature. For lighter colored roofs, this difference is reduced to 20° F (11 °C)

(Guyette, 2002). Therefore, the maximum air temperature observed during conditioning of

99 °F (37 °C) translates into a roof temperature of 119 °F (48 °C) to 189 °F (87 °C) depending on the absorbency of the roof. The manufacturer provided glass transition temperature (Tg) for the epoxy used in this study is 113 °F (44.9 °C) (S. Witt, personal communication, May 2, 2010). As per the literature review, on average, Tg for most adhesives used in FRP-strengthening applications is 131 °F (55 °C). Therefore, it is most likely that the Tg of the epoxy was reached during this study because the roof temperatures varied between 119 °F (48 °C) to 189 °F (87 °C). Additionally, the texture of the epoxy became rubbery, which is another sign that temperatures near the Tg were reached as

105 discussed in the literature review. Further, these specimens did not have UV protection applied. Therefore, with exposure to UV rays, the elastic modulus and strength of the FRP and epoxy resin were more than likely negatively impacted as discussed in the literature review. The change in the failure mechanism of the naturally conditioned specimens occurred after only a nine-month period of exposure to moisture, UV rays, and temperatures near the glass transition temperature with no protection.

100

90 Series 3 32.22 80

70 22.22

C ° 60 ° 12.22 50

2.22 Temperature, Temperature, 30 Temperature,

20 -7.78 Series 2 10

0 -17.78

Maximum Temperature Minimum Temperature

Figure 4-24. Daily Temperature Data for Naturally Conditioned Specimens

106 3 70 Series 2 Series 3 2.5 60

2 50

1.5 40 30

Precipitation, in Precipitation, 1 20 mm Precipitation, 0.5 10

0 0

Figure 4-25. Daily Precipitation Data for Naturally Conditioned Specimens

The environmental chamber provided the extreme environmental conditions for the prisms. Several cycles were ran in which the temperature ranged from ambient to -40°F (-

40°C) before being held at -40°F (-40°C) for 45 hours. In the following two hours, the temperature was allowed to return to ambient as shown in Figure 4-26. This cycle type was run four times in May of 2009, which was 3 months prior to tests run on the second series.

Three additional extreme cycles were run in June of 2009 in which the temperature was lowered to -40°F (-40°C) and held for 52 hours before returning to ambient. Then, the temperature was lowered to -50°F (-46°C) and held for 74 hours. In the following two hours, the temperature was increased to 180°F (82°C), which is also near the glass transition temperature of the epoxy, and held for 16 hours before allowing the temperature to return to

107 ambient as seen in Figure 4-27. An additional eight runs were ran in the environmental chamber in May and August of 2009, which occurred before the second series of testing, in which the temperature was lowered to -40°F (-40°C) and held for an hour before returning to ambient. Shear failure through the adhesive observed for the naturally conditioned specimens was not seen for those placed in the environmental chamber. Specimens from the environmental chamber were not exposed to the moisture that the naturally conditioned specimens saw. Although temperatures seen by the extreme conditioned groups were of a greater range than the naturally conditioned group, they were cyclic as the temperature was allowed to return to ambient. Naturally conditioned specimens sustained these temperatures while also being exposed to precipitation events. Further, the extreme conditioning treatment did not expose the specimens to high temperatures for sustained periods of time as in the case of the naturally conditioned specimens.

108 80 18.9 60 8.9

C ° -1.1 ° 20 -11.1 0 -21.1

0 10 20 30 40 50 60 Temperature, Temperature, -20 Temperature, -31.1

-40 -41.1

-60 -51.1 Time, hours

Figure 4-26. Extreme Conditioning – Cycle Type 1

200 86.9

150 66.9

46.9 F

100 C

° ° 26.9

50 6.9

-13.1 Temperature, Temperature, 0 Temperature, 0 50 100 150 200-33.1 -50 -53.1

-100 -73.1 Time, hrs

Figure 4-27. Extreme Conditioning – Cycle Type 2

109 As previously discussed, the deterioration of the bond between test dates encouraged the specimen to fail through the adhesive in shear. Therefore, it is fitting that more cracking was visible in the second series than in the last. The change in the failure mechanism, although different from what was expected, actually increased the bond resistance. As the failure mechanism shifted from debonding to shear failure in the adhesive between the second and third series, the bond resistance of all naturally conditioned specimens, regardless of construction type, increased.

However, this increase in debonding resistance for the naturally conditioned specimens is misleading. Splitting of the brick is typically associated with a premature failure.

Therefore, the debonding resistance is much lower than anticipated. The rough surface of the debonding crack that develops in the masonry substrate must dilate in order to slip. Without cracking in the masonry substrate, as observed in the naturally conditioned specimens, there is no dilation of the debonding failure plane and hence, splitting does not occur. Therefore, the premature failure caused by splitting of the brick was avoided as failure was in shear through the adhesive. Typically, shear failure through the adhesive would produce a smaller debonding resistance. However, because the control prisms are failing prematurely, it appears as though the naturally conditioned specimens fail at a higher debonding resistance.

4.4 Comparisons to Previous Models

4.4.1 Willis Masonry Model

The premature failure observed in the control prisms can be proven to drastically reduce the debonding resistance upon comparison with a published model that predicts the IC

110 debonding resistance of an FRP-reinforced masonry prism. In a recent paper by Willis et al.

(2009), 29 pull-tests on FRP-reinforced masonry prisms were performed. Willis et al. (2009) suggested a generic model to be used with FRP-reinforced masonry prisms that predicted the debonding resistance of the strengthening system. The model accounts for the axial rigidity of the FRP as well as the brick lateral modulus of rupture. The equation, as previously described in Chapter 2 of this thesis, is presented in Equation 4-1 in which units of Newtons and millimeters must be used.

Equation 4-1

The cross-sectional area of the FRP used in this study was of dimensions 0.5 x 0.079 inches (12.7 x 2 mm). The elastic modulus of the FRP, tested as described in Chapter 3, was determined to be 26,944 ksi (185,770 MPa) and the lateral modulus of brick was found to be

478 psi (3.3 MPa). Incorporating these values into the generic masonry model, Equation 4-1, the IC debonding resistance, PIC, was found to be 12,190 lbs (54.22 kN). This predicted debonding resistance applies to all control tests however it is not expected to capture differences from conditioning treatments.

Table 4-4 presents the experimental debonding resistance, Pexp, compared to the theoretical debonding resistance, PIC, found from the generic masonry model developed by

Willis et al. (2009). Because the Willis Masonry Model does not take into account environmental conditioning, the most relevant comparison can be made from the control specimens. The average difference between the predicted and the experimental debonding resistance is 2.03, which signifies that the Willis Masonry Model is overpredicting the

111 debonding resistance by more than double. Comparing the predicted to the experimental values for the extreme conditioned specimens, the Willis Masonry Model more than doubles the debonding resistances determined experimentally. Further, the naturally conditioned specimens, which reached the highest loads during testing, were also overpredicted by 177%.

The coefficients of variation for all specimen groups were less than 0.3, which is the acceptable limit for masonry tests (Daniels et al., 2009).

Table 4-4. Theoretical versus Experimental Test Results

* Specimen Experimental Load (Pexp), lbs (kN) PIC /Pexp CS1-0 7,559 (33.62) 1.61 CS2-0 5,193 (23.10) 2.35 CS3-0 4,908 (21.83) 2.48 CR1-0 6,004 (26.71) 2.03 CR2-0 6,746 (30.01) 1.81 CR3-0 4,066 (18.09) 3.00 CR4-0 6,091 (27.09) 2.00 CS1-5 6,778 (30.15) 1.80 CR1-5 6,123 (27.23) 1.99 CR2-5 4,206 (18.71) 2.90 CR3-5 5,836 (25.96) 2.09 CS1-9 7,440 (33.09) 1.64 CS2-9 7,325 (32.58) 1.66 CR1-9 7,573 (33.68) 1.61 CR2-9 8,324 (37.02) 1.46 Mean 2.03 Standard Deviation 0.49 Coefficient of Variation 0.24

*PIC = 12,190 lb (54.22 kN)

112 Table 4.4. Continued

* Specimen Experimental Load (Pexp), lbs (kN) PIC /Pexp ES-5 5,580 (24.82) 2.18 ER-5 6,092 (27.10) 2.00 ES1-9 4,144 (18.43) 2.94 ES2-9 4,610 (20.51) 2.64 ER-9 6,939 (30.87) 1.76 Mean 2.30 Standard Deviation 0.48 Coefficient of Variation 0.21

*PIC = 12,190lb (54.22 kN)

* Specimen Experimental Load (Pexp), lbs (kN) PIC /Pexp NS-5 5,011 (22.29) 2.43 NR-5 6,357 (28.28) 1.92 NS-9 9,240 (41.10) 1.32 NR-9 8,594 (38.22) 1.42 Mean 1.77 Standard Deviation 0.51 Coefficient of Variation 0.29

*PIC = 12,190 lb (54.22 kN)

From Table 4-4, it is apparent that the theoretical model found by Willis et al. (2009) overpredicts the debonding resistance found in this experimental study regardless of environmental exposures or construction pattern. The materials used by Willis et al. (2009) were similar to those used in this research, which are presented in Section 3.4 of this thesis.

The brick units in the Willis et al. (2009) study had a lateral modulus of rupture of 551 psi

(3.80 MPa). These bricks were used in masonry compressive tests, which resulted in a compressive masonry strength of 2,321 psi (16 MPa) and elastic modulus of 513,289 psi

113 (3,539 MPa). The bond wrench test resulted in a flexural tensile strength of masonry of 88 psi (0.61 MPa). Therefore, the material properties were not the reason for the large difference in the prediction versus the experimental IC debonding resistance.

The generic model does not incorporate the type of epoxy or mortar strength and it is possible that the material properties of each alter the behavior of the FRP-reinforced specimen. Furthermore, the amount and type of cores that may be present in the brick is not accounted for in Equation 4-1. In the study completed by Willis et al. (2009), 10-core bricks were used whereas in this study, three larger cores were present in the brick. Upon comparing the solid pavers in the Petersen et al. (2009) study to the ten-core specimens of the

Willis et al. study (2009), it was found that prisms with cores had a lesser debonding resistance. The location of the cores could also affect the behavior of the specimen. It is suggested that to prevent splitting of the brick, the FRP strip should not come within too close of a distance to the cores (Petersen et al., 2009). However, the FRP strip’s location to adjacent cores is not included in the Willis Masonry Model though it has been suggested that this will decrease the debonding resistance.

The large difference in debonding resistance between the predicted and experimental values can be attributed to the specimens in this study splitting prematurely through the back of the masonry prism. Figure 4-28 presents the load-slip curves for a control specimen from this study and a specimen constructed by Willis et al. (2009). Although the curves presented are not for the exact same specimen, the drastic difference in the observed response is illustrated. In the study by Willis et al. (2009), it was observed that typically the slip increased after the peak load was reached as seen in Figure 4-28. However, while the slip

114 increased, the peak load remained constant, which is typical for most reported pull-tests. In this study, it was found that there is a sudden failure immediately following the peak load.

Immediately following debonding, it was determined that the specimens split through the entire thickness of the prism. Because the prism was splitting through, the FRP was no longer resisting any load as it pulled out of the specimen. For almost all specimens, it was observed that cracking present on the back face of the prism resulted in the immediate loss of strength post-peak.

Slip, mm 0 0.5 1 1.5 2 2.5 3 3.5 10,000 9,000 40 8,000 35 Willis et al. (2009) 7,000 (Typical Behavior) 30 6,000 25 5,000

CS1-9 20 Load, lbs Load, 4,000 kN Load, 15 3,000 2,000 10 1,000 5 0 0 0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14 Slip, in

Figure 4-28. Comparison of Load-Slip Curves for Willis et al. (2009) Study versus Specimen in Current Study

115 4.4.2 Comparison to Petersen et al. (2009) Tests

Petersen et al. (2009) discussed the differences between their study and the study performed by Willis et al. (2009). It was cited that key differences between the debonding resistances in the two research programs could be attributed to the construction patterns of the specimens, the FRP width, the epoxy used, and the lateral modulus of rupture for the bricks. Furthermore, Petersen et al. (2009) used thicker FRP strips that were stronger than those used by Willis et al. (2009) and solid bricks instead of 10-core specimens. Table 4-5 presents the experimental results of the pull-tests similar to the construction types, i.e. stacked and running bond, used in this research. Petersen et al. (2009) used two 0.06 inches

(1.4 mm) thick CFRP strips that were 0.59 inches (15 mm) wide. The experimentally determined lateral modulus of rupture of the bricks was 518 psi (3.57 MPa) and the elastic modulus of the CFRP was 30,026 ksi (207,021 MPa). Using these values in the Willis

Masonry Model in Equation 4-1, the predicted IC debonding load is 17,127 lbs (76.18 kN).

The specimens denoted as S1-A is the stack-bonded specimen where S1-C represents the running bond specimen. In all but one case, the specimens tested by Petersen et al. (2009) are generally within ±10% of the predicted IC debonding load from the Willis Masonry

Model and the mean ratio is 1.02 with a standard deviation of 0.12 and a coefficient of variation of 0.12.

116 Table 4-5. Theoretical versus Experimental Test Results, Petersen et al. (2009) Specimens

Specimen Peak Load, lbs (kN) PIC/Pexp S1-A-NG-1 18,760 (83.45) 0.91 S1-A-NG-2 15,981 (71.09) 1.07 S1-A-SG 18,317 (81.48) 0.94 S1-C-NG-1 14,360 (63.88) 1.19 S1-C-NG-2 15,603 (69.41) 1.10 S1-C-SG 18,996 (84.50) 0.90

Petersen et al. (2009) observed that the specimens had cracking visible along the top course of brick and also through the thickness of the prisms. Furthermore, some specimens also exhibited cracking along the back of the prism in line with the FRP strip. Running bonded specimens S1-C-NG-2 and S1-C-SG also split through the prism and the cracking had followed the mortar head joints. Petersen et al. (2009) deduced that this type of cracking would reduce the bond resistance of an FRP-reinforced structure. However, we see from the comparison in Table 4-5 that even with cracking through the specimen, the experimental failure load was well predicted by the Willis et al. (2009) model. However, splitting in the

Petersen et al. (2009) study occurred towards the end of the test, therefore, splitting did not cause a reduction in debonding resistance.

Figure 4-29 displays the load-slip response of prisms tested by Petersen et al. (2009)

Isolating prisms S1-A and S1-C, the load following the peak, decreases as the slip increases.

Comparatively, the load-slip behavior for prisms tested in this study is similar in that failure occurs immediately following the peak load. This response could be accredited to the splitting of specimens in both studies. Due to the higher debonding resistance, the slip at the

117 peak load in the Petersen et al. study is approximately 1.25 mm, which is higher than the slip at the peak loads for prisms tested in this research. The initial stiffness recorded for both studies is very comparable.

Slip, mm 0 0.5 1 1.5 2 20,000 18,000 Petersen et al. 80 Specimen S1-C-SG 16,000 70 14,000 Petersen et al. 60 Specimen S1-A-SG 12,000 50 10,000 CS1-9

40 Load, lbs Load, 8,000 kN Load, 30 6,000 4,000 CR1-9 20 2,000 10 0 0 0.00 0.02 0.04 0.06 0.08 Slip, in

Figure 4-29. Load-Slip Behavior, Petersen et al. (2009) Specimens vs. Current Study Specimens

Both Petersen et al. (2009) and Willis et al. (2009) used the same commercially available adhesive. Provided manufacturer’s data show a compressive strength of 8,702 (60

MPa) and flexural strength of 4,351 psi (30 MPa). Additional properties such as the elastic modulus, tensile strength, and Poisson’s ratio were also experimentally determined by Willis et al. (2009) (D. Hale, personal communication, February 19, 2010). The adhesive utilized in

118 the Willis et al. (2009) study had an elastic modulus of 972 ksi (6.7 GPa), tensile strength of

2,012 psi (13.9 MPa) and Poisson’s ratio of 0.312. Comparatively, the compressive strength of the epoxy used in this research was 4,088 psi (28 MPa) and the tensile strength was 2,423 psi (17 MPa). The difference in epoxy strength is not accounted for in the model presented by Willis et al. (2009) and therefore could also contribute to the difference between the predicted IC debonding load and the experimental debonding load. Furthermore, the porosity of the brick and viscosity of the epoxy could alter the debonding resistance. It has been found that the porosity of the brick directly affects the bond between the masonry and the

FRP (Roko, Boothby, & Bakis, 1999). The bricks used in the study by Willis et al. (2009) and Petersen et al. (2009) have different material properties than those used in this study.

Hence, the change in porosity could also contribute to the lesser experimental debonding load however, this would not alter the fundamental failure mode of the prisms such as splitting or shear failure through the adhesive.

4.4.3 Comparison to ACI 440.7R-10

ACI 440.7R-10 places a limit on the strain that can be developed in the FRP of FRP strengthened masonry structures. The strain that can be developed is limited by crushing of the masonry or by debonding of the FRP. The strain at which FRP debonds is limited by a factor, κm, which is the experimentally-determined bond reduction coefficient. For NSM

FRP systems, κm is equal to 0.35. The ultimate strain of the FRP is reduced by this factor to achieve an effective strain to be used in the design of FRP-strengthened masonry structures.

In this thesis, the ultimate strain determined from FRP coupon tests was 1.48% therefore the effective strain according to ACI 440.7R-10 is 0.00518 in/in (0.00518 mm/mm). The

119 experimentally determined elastic modulus from the FRP coupon testing was 26,943,713 psi

(186 GPa) and the FRP strips were each 0.079 inches (2 mm) thick and 0.5 inches (12.7 mm) wide. Therefore, the load at which debonding is predicted to occur is determined by multiplying the strain by the axial rigidity (EA) of the FRP. Therefore, the resulting debonding resistance is 5,495 lbs (24.4 kN). The experimentally determined resistance from the pull-tests was 6,534 lbs (29.1 kN) for the stacked specimens and 6,142 lbs (27.3 kN) for the running bonded specimens, which are 19% and 12% higher, respectively, than that predicted by ACI 440.7R-10. The ACI prediction may possibly be lower as it is based on experimental data with NSM FRP systems that may not be the most efficient, such as glass

FRP bars in horizontal bed joints. Therefore, it is expected to be inherently conservative as part of a design guideline.

4.5 Conclusions

From the pull-tests performed in this study, some general conclusions can be drawn about the behavior of the FRP-reinforced specimens exposed to various environmental conditioning. Isolating the specimens constructed in a running bond clearly shows the increase in bond resistance between series. Each running bonded specimen from the three conditioning environments increased in bond strength by the following test period. This was not true for the specimens constructed in a stack-bond. Conversely, Petersen et al. (2009) found stack-bond specimens had debonding resistances that were 8% greater than their stack- bonded counterparts. Willis et al. (2009) found similar debonding resistances for both types of specimens. In this study, although the FRP strips were located closer to vertical mortar

120 joints in the running bond pattern, the FRP strip was located away from the cores. Therefore, there was more resistance to splitting as this failure mechanism dominated over any influence from the vertical mortar joints.

Although specimens from the control and natural environments increased in strength over time, the specimens exposed to extreme conditioning decreased in strength. However, the unanticipated increase in bond resistance for specimens subjected to natural conditioning can be attributed to the prism not splitting through thus allowing it to reach a higher load from the different debonding mechanism. This research has shown that the natural weathering of the FRP-reinforced masonry prisms altered the failure mode of the specimen.

Shear failure through the adhesive would typically produce a lesser debonding resistance.

However, splitting of the control and extreme conditioned specimens resulted in premature failures that were at debonding resistances less than failure through the adhesive.

The natural conditioning environment is the most similar to the exposures that masonry structures will see during its service life. Therefore, it is important to prevent the deterioration of the bond as this may cause failure at lower loads than anticipated. It is also important to prevent splitting of the brick wall as this too could cause premature failure in a masonry structure. Additionally, it was found that in most instances, the specimens constructed in a running bond pattern, which is the typical construction pattern for most masonry structures, exhibited a debonding load near or exceeding the bond strength of the stack-bonded specimens. However, this is related to the running bond specimen’s resistance to splitting, as the FRP strips are located further away from the cores. Within a larger wall,

121 some lateral restraint against splitting will be provided by the surrounding bricks; therefore, the vertical mortar joints will have a greater influence on the debonding resistance.

Upon comparison of the test results from this study to the IC debonding resistance predicted by the Willis et al. model, it was clear that the debonding resistances seen in this study were much lower than predicted. The porosity of the bricks could produce the differing bond strength as moisture propagation and epoxy seepage behave differently in different bricks. Furthermore, the material properties of the epoxy and mortar are unaccounted for in the model. However, splitting of the prisms in this study is considered to be the major factor in the reduction of debonding resistance. In the following chapter, the effects of some of these variables will be presented through the use of numerical modeling of pull-tests completed in this study as well as the Petersen et al. (2009) and Willis et al. (2009) studies.

122 CHAPTER 5 – NUMERICAL ANALYSIS

5.1 Introduction

The following chapter presents the applied element modeling of pull-test specimens tested in this research. Using the program Extreme Loading for Structures® developed by

Applied Science International, the debonding behavior of FRP-reinforced masonry prisms can be modeled accurately. Extreme Loading for Structures® (ELS) uses the Applied

Element Method (AEM) to model all stages of loading from crack initiation to debonding.

Previous work using ELS has been for assessing vulnerability of structures and risk mitigations following blasts, impacts, or earthquake events. ELS also allows users to simulate structural demolitions (Extreme Loading, 2009).

5.2 Applied Element Modeling

During seismic and other “extreme” loading events, structures are prone to not only partial damage but also complete collapse. Collapse of masonry structures can cause injury and death to the occupants and cause the collapse of nearby structures (Meguro & Tagel-Din,

2000). One-third of the world’s population live in unreinforced masonry structures although the construction and the materials used in these types of buildings vary greatly. To ensure the safety of the occupants, it is often necessary to strengthen these structures. However, due to the wide variety of masonry material and structures around the world, it is not feasible from an economic or time standpoint to test each possible combination of material and strengthening system that could be used. Therefore, numerical simulations can be utilized to

123 model and observe the behavior of many different types of masonry buildings (Mayorca &

Meguro, 2003).

Numerical analysis of structures is typically divided into two subcategories of models based on continuum material equations or discrete elements. The first category describes methods similar to the Finite Element Method (FEM) in which analysis of a structure is accurate up until collapse (Meguro & Tagel-Din, 2000). Cracks in this method are defined by the “Smeared Crack” approach in which elements do not separate at a crack.

Furthermore, the crack does not have a distinct shape or size inside each element. The

“Discrete Crack” method is where a crack location must be predicted and identified in the model prior to performing the numerical simulation. Therefore, one must be experienced in the cracking behavior of a structure prior to running the numerical analysis to properly predefine the crack location. However, in cases where cracks can be unpredictable, this method may not be suitable and it may force failure to occur in a certain way depending on the crack prediction (Extreme Loading, 2006).

Discrete elements are used in methods such as the Rigid Body and Spring Model and the Extended Distinct Element Method, which are more proficient at simulating crack propagation. However, these methods rely on the element shape and size to accurately simulate cracking. The Rigid Body and Spring Model is similar to FEM in that it cannot accurately simulate complete collapse of a structure whereas the Extended Distinct Element

Method can. Unlike the aforementioned numerical analysis methods, the Applied Element

Method (AEM) is accurate in simulating cracking behavior from initial loading to ultimate

124 collapse of a structure. Computing time with AEM is kept low and the programming required is relatively simple to implement (Meguro & Tagel-Din, 2000).

AEM divides a structure virtually into small elements as seen in Figure 5-1 (A). The two elements highlighted in Figure 5-1 (A) are enlarged in Figure 5-1 (B) showing an array of springs connecting the elements, each with a normal and two shear springs. These springs are distributed over the entire surface of an element, therefore connecting the entire face of the element to adjacent elements. The displacement of connecting springs relates to the stresses and deformations within the area highlighted. Furthermore, partial element connectivity is allowed by the failure of some of the connecting springs along the surface.

This is not possible in FEM unless the nodes that connect the elements disconnect. AEM is more capable at modeling through progressive collapse of a structure than FEM (Extreme

Loading, 2006).

Each element in AEM has six degrees of freedom, three rotational, and three translational that determine the rigid body motion. On the contrary, typical 8-node elements in FEM have 16 degrees of freedom (Extreme Loading, 2006). However, due to the nature of the springs, even though the shape of the element remains rigid during loading, the elements’ stresses and deformations are calculated based on the spring deformations (Meguro

& Tagel-Din, 2000).

125 (A) (B)

Figure 5-1. AEM Modeling (A) Element Generation (B) Spring Distribution of each Pair of Springs (Meguro & Tagel-Din, 2000)

The normal and shear stiffness of the springs are calculated by:

Equation 5-1

Equation 5-2

in which E is the Elastic Modulus of the material and G is the Shear Modulus of the material.

The shaded area in Figure 5-1 (B) is defined by a height of d, which is the separation between springs. The length of the area is a, which is also the width of the element. In the stiffness calculation, the term T describes the thickness of the element (Meguro & Tagel-Din,

2000). In ELS, if adjacent elements are of different material properties or sizes, then the

Elastic Modulus is set to the minimum Elastic modulus of the elements and the area is the average between both elements. Furthermore, the distance a is determined from the average element size (Extreme Loading, 2006). Each element stiffness matrix is found by applying unit displacements to each degree of freedom direction. Load is applied in AEM by the equation:

126 Equation 5-3 in which [KG] represents the global stiffness matrix, [Δ] is the displacement vector, and [F] represents the applied load vector, which is determined prior to analysis in a load control setting. On the other hand, the load is found in displacement control cases by moving degrees of freedom by unit displacements. AEM then performs a transformation of the global stiffness matrix into a vector that is one-half the global stiffness matrix after eliminating the zero terms. Therefore, running an analysis on only the non-zero terms takes less computation time than other methods (Meguro & Tagel-Din, 2000). Then, the stresses and strains for each spring are calculated based on spring deformations for each loading step and are checked against the failure criteria (Tagel-Din & Meguro, 2000). Similar to concrete, springs subjected to tension in the brick and mortar will have zero stiffness following reaching the cracking point which is when the principal stress reaches the tensile strength (Extreme Loading, 2006). Therefore, the stiffness matrix is then reassembled after each loading step (Tagel-Din & Meguro, 2000).

Running a numerical simulation requires optimization of the number of connecting springs and size of the elements in order to reduce computational time to a reasonable level while maintaining accuracy. Upon increasing the number of connecting springs between elements, cracking is more accurately modeled. Furthermore, increasing the number of elements increases accuracy yet it also increases computation time. Smaller elements will also ensure accurate displacements. However, if an analysis does not require accuracy of the shear stresses involved in the model, then small elements are not necessary. For walls and

127 deep beams, it is recommended to utilize smaller sized elements to capture the shear behavior

(Meguro & Tagel-Din, 2000).

The shape of the element is also an important factor in numerical analysis and it is recommended to use a cube-shape element in most AEM models. Furthermore, it is best to progressively change the size of an element so that very large elements are not directly next to small elements. Smaller elements should also be concentrated around areas with high stresses to capture the change in stresses. However, transition elements are not required in

AEM as they are in FEM. Furthermore, the elements should be in the same plane so springs will connect between them. However, elements do not have to be in line with each other in this plane, as faces only have to be partially connected. In FEM, this would require more nodes, thus increasing computation time (Extreme Loading, 2006).

5.3 Modeling Parameters in Extreme Loading for Structures®

The parameters described in the following section are required as inputs for each material used in the structural model. Young’s Modulus, E, describes the relationship between the longitudinal stress and strain in the elastic range of the material. In order to obtain accurate displacements, it is imperative to input a correct value for the Young’s

Modulus. The Shear Modulus, G, is also calculated in the elastic range of the material and describes the relationship between the shear stress and shear strain. It is determined by:

Equation 5-4

Where ν is the Poisson’s ratio. The Tensile or Compressive Strength of the material is when tensile or compressive failure occurs, respectively. The Ultimate Strength/Yield Stress ratio

128 is the value of the ratio between the ultimate strength and the yield stress of a material. This is only used for models that include reinforcing bars. The Separation Strain is the value at which the connecting springs will separate between elements. Typically, values range from

0.1 to 0.2 with 0.1 being set as the default and the value used in this study. For comparison, a value of 0.01 was used for the model of a stack-bonded specimen used in this study. The peak load was increased by less than 0.25% therefore the default value of 0.1 was used for the remaining models. Following separation, elements will act as rigid bodies that are free to contact other parts of the structure. The friction coefficient is activated after separation between elements occurs and is applied when these free elements make contact with each other.

The Specific Weight of a material is calculated by ELS as gravity multiplied by the specimen’s mass. Self-weight of materials will not apply unless the specific weight is set to a value not equal to zero. Additional features of ELS that are primarily used for collapse and demolition scenarios and were not used in this study include the Normal and Shear Contact

Stiffness Factors which represent the normal and shear stiffness of the connecting springs that are located between elements. The Normal Spring Stiffness Factor has a default value of

0.01 whereas the Shear Spring Stiffness Factor default is 0.001 because it is typically smaller than the normal force. Both factors are necessary in models where elements will break apart and collide with other elements. The contact stiffness will determine the amount of force transferred to another element at collision. The Contact Spring Unloading Stiffness Factor defines how much energy is lost when elements contact one another. When it is set to 1, no energy is lost however when it is set to 10, 90% of the element’s energy is lost and the

129 element does not rebound as significantly upon collision. The Post-Yield Stiffness Ratio, applicable for steel materials, is the ratio between the post-yield and pre-yield modulus and is necessary for calculating the post-yield modulus.

After modeling the geometry and boundary conditions of the structure, the loading scenario is ready to be implemented. For the pull-tests simulated in this study, static displacement control applies a load to specific elements. The displacement load is then applied in increments to the selected elements up to a predetermined ultimate displacement.

Increasing the number of increments will create a more accurate solution. Unlike load control, the solver will continue following post-peak behavior with displacement control.

The peak load can then be determined from the output. However, displacement control cannot be used when loads are distributed or if loads are applied to multiple locations on the specimen (Extreme Loading, 2006).

5.4 Numerical Analysis

5.4.1 Introduction

Six different pull-test specimens were utilized to determine the suitability of using a numerical analysis program such as ELS as a tool to model pull-tests. The chosen specimens were selected from published articles that provided pull-test data for NSM FRP-reinforced masonry prisms. Thus, the IC debonding load and some load-slip data was available to validate the ELS model against experimental results.

130 5.4.2 Current Study’s Specimen

Both the stacked and running bond specimens from this study were modeled in ELS.

Both specimens were of the dimensions as seen in Figure 5-2 and had the same material properties as discussed in Chapter 3 of this thesis. The prisms, as modeled in ELS, are shown in Figure 5-3.

2" (51 mm) 2 1/4" (57 mm)

Ø1.5" (38 mm) 3 1/2" (89 mm)

1 3/4" (44 mm) 2" (51 mm) 7 5/8" (194 mm)

Figure 5-2. Current Study's Brick Dimensions (Actual)

Figure 5-3. Pull-Test Specimens

131 Material properties used for the model are provided in Table 5-1. The compressive strengths for both the brick and mortar were experimentally determined and results were presented in Chapter 3. The tensile strength of the brick was calculated from the lateral modulus of rupture, which was experimentally determined to be 551 psi (3.8 MPa) as tested by the set-up described in Chapter 3 of this thesis. The lateral modulus of rupture corresponds to the flexural tensile strength of the material. As per the discussion by Petersen

(2009), to relate the flexural tensile strength to the tensile strength, the lateral modulus of rupture is reduced by a factor of 1.5. This factor is deduced from dividing the flexural tensile strength from Equation 5-5 by the tensile strength given by Equation 5-6 to result in Equation

5-7. Therefore, the tensile strength of the brick is estimated to be 367 psi (2.53 MPa).

Equation 5-5

Equation 5-6

Equation 5-7

The tensile strength for the mortar was assumed to be one-tenth of the compressive strength of the mortar that was experimentally determined as described in Chapter 3. The elastic modulus of the mortar is a function of its compressive strength. Compressive tests performed on mortar cubes with compressive strengths between 450 psi (3.1 MPa) and 2,988 psi (20.6 MPa) presented a strong correlation between the elastic modulus and 200 times the compressive strength (Kaushik, Rai, & Jain, 2007). Therefore, the elastic modulus was estimated to be 490,000 psi (3,378 MPa).

132 The elastic modulus for the brick was based on the elastic modulus found in the prism compression tests, which encompassed five brick courses and four mortar joints. The average elastic modulus for the control specimens tested as discussed in Chapter 3 of this thesis was found to be 1,101,746 psi (7,596 MPa). The Uniform Building Code relates the elastic modulus of the prism to the elastic moduli of the mortar and brick by Equation 5-8

(Bakhteri, Makhtar, & Sambasivam, 2004):

Equation 5-8

Where:

Equation 5-9

Equation 5-10

The factor γt in Equation 5-9 represents the aspect ratio of the thickness of the mortar joint, tj, to the depth of the brick units, tb and γm of Equation 5-10 is the ratio of the elastic modulus of the mortar, Ej, to the elastic modulus of the brick, Eb. The thickness of the mortar joint is

3/8” (9.5 mm) and the depth of the brick is 2 ¼” (57 mm). The elastic modulus of the brick was then calculated as 1,391,228 psi (9,592 MPa). The shear modulus of the brick was calculated based on the assumption that the Poisson’s ratio was 0.2 for each material as assumed in Petersen (2009).

Material properties such as the elastic modulus and ultimate tensile strength of the

FRP were also discussed in Chapter 3. The shear modulus was calculated based on the assumption that the Poisson’s ratio for the material was 0.3 as assumed by Petersen (2009)

133 while the compressive strength was assumed to be one-tenth of the tensile strength of the

FRP. The compressive strength of the FRP is not necessary as the FRP is subjected strictly to a tensile load in the model. The epoxy properties were provided by the manufacturer with the exception that the shear modulus was based on a Poisson’s ratio of 0.32, which was similar to the value experimentally determined through testing completed by Willis (D. Hale, personal communication, February 19, 2010).

Table 5-1. Material Properties

Property Brick Mortar CFRP Epoxy Elastic Modulus, ksi (GPa) 1,391 (9.6) 490 (3.4) 26,944 (162) 225 (6.7) Shear Modulus, ksi (GPa) 580 (4.0) 204 (1.4) 10,363 (71) 85 (2.6) Compressive Strength, psi (MPa) 10,849 (75) 2,450 (17) 31,602 (275) 4,088 (60) Tensile Strength, psi (MPa) 367 (2.53) 245 (1.7) 316,024 (2,754) 2,423 (14)

Loading was applied to the end element of the FRP strip in 200 time steps to a final displacement of 0.15 inches (3.81 mm). The pull-test modeled is for the control specimens that were not subjected to environmental conditioning. Figure 5-4 illustrates crack propagation (highlighted as blue) through the stack-bonded pull-test specimen. The cracking on the bottom of the prism is a byproduct of ELS forcing the user to choose a material property for the ground surface. In this model, the ground was chosen as mortar and had no bearing on the results of the numerical analysis. The mortar being the weakest material cracks first. The simulation follows closely the behavior observed experimentally as cracking began in the first brick course and then propagated towards the bottom of the prism.

134

Figure 5-4. Crack Propagation in Stack-Bonded Specimen

135

Cracking through 1st Course, Cracking through 2nd Course, P = 7,011 lbs (31.2 kN) P = 7,501 lbs (33.4 kN)

Cracking through 3rd Course, Cracking through 4th Course, P = 8,226 lbs (36.6 kN) P = 8,286 lbs (36.9 kN)

136

At Ultimate Load, At Final Condition, P = 8,486 lbs (37.7 kN) P = 0.22 lbs (0.0 kN)

Figure 5-5 presents the load-slip curves for the ELS model and the control stack- bonded specimens from the final series of experimental testing. It was observed experimentally that the control prisms increased in debonding resistance with age.

Therefore, the control prisms from the final series were selected for comparison with the ELS model, as these prisms would represent the condition of brick masonry walls in service. The ultimate debonding resistance observed through the numerical analysis was 8,486 lbs (37.7 kN) and the slip at this load was 0.021 inches (0.53 mm). Comparing the ultimate load and associated slip in Table 5-2, it becomes obvious that ELS overpredicts the debonding resistance but provides too low of an estimate for the slip at the peak load. However, the peak load is only overpredicted by approximately 15%, which is very acceptable given the material and construction variability of masonry being so high. Typically, a coefficient of

137 variation up to 30% is considered to be within acceptable limits for tests completed on masonry due to the heterogeneous nature of the material (Daniels et al., 2009). The data provided by ELS is considered acceptable as the elastic moduli of the brick and mortar as well as the tensile strength of the mortar was estimated. Additionally, the splitting that was experimentally observed was not as intense in the numerical simulation. Cracking did propagate across the top of the specimen as seen in Figure 5-6 although it did not reach the back of the prism. This could also account for the higher debonding resistance seen in the numerical model than in the experimental data as seen in Table 5-2. Furthermore, due to the non-homogenous nature of masonry prisms, some variability should be expected such as in the case of the initial stiffness observed for specimens CS1-9 and CS2-9. The stiffness of

CS1-9 follows very closely to the stiffness determined through the numerical analysis however, failure was immediate which was quite different than specimen CS2-9.

138 Slip, mm 0 1 2 3 4 5 6 9,000 40 ELS Model - Stacked Specimen 8,000 CS1-9 35 7,000 30 6,000 25 5,000 20

4,000 CS2-9 Load, lbs Load, 15 kN Load, 3,000 2,000 10 1,000 5 0 0 0.00 0.05 0.10 0.15 0.20 0.25 Slip, in

Figure 5-5. Comparison of Numerical to Experimental Load Curves – Stack-Bonded Specimens

Table 5-2. Comparison of Numerical to Experimental Data – Stack Bonded Specimens

Specimen Pmax, lbs (kN) Slip At Pmax, in (mm) PELS/Pexp SlipELS/Slipexp ELS 8,486 (37.7) 0.021 (0.52) - - CS1-9 7,440 (33.1) 0.024 (0.61) 1.14 0.87 CS2-9 7,325 (32.6) 0.045 (1.14) 1.16 0.45

139

Extent of Cracking

Figure 5-6. Splitting of Top Brick

Similar to the stack-bonded specimen, cracking began in the top course of brick of the running bonded specimen. However, cracking also propagated in the bottom courses sooner than what was experimentally observed. As the load increased, cracking was then observed to propagate from the second to fourth course as experimentally observed. Figure 5-7 shows the complete cracking pattern of a running bond specimen subjected to the pull-test.

140

Figure 5-7. Crack Propagation in Running Bonded Specimens

141

Cracking through 1st Course, Cracking through Subsequent Courses, P = 6,011 lbs (26.7 kN) P = 8,515 lbs (37.9 kN)

Cracking through 2nd Course, At Ultimate Load, P = 9,037 lbs (40.2 kN) P = 9,169 lbs (40.8 kN)

142

Post-Peak, At Secondary Peak, P = 8,162 lbs (36.3 kN) P = 8,503 lbs (37.8 kN)

Post Secondary Peak, At Final Condition, P = 6,674 lbs (29.7 kN) P = 0.09 lbs (0.0 kN)

143 The load-slip curves for the ELS model and the control running bonded specimens from the final series of experimental testing are presented in Figure 5-8. The ultimate debonding resistance observed through the numerical analysis was 9,169 lbs (40.8 kN) and the slip at this load was 0.0175 inches (0.45 mm). The initial stiffness for both experimental curves follows very closely to the model in Figure 5-8. Comparing the ultimate load and associated slip in Table 5-3, again the ELS numerical simulation overpredicts the debonding resistance by 21% and 10%. However, the associated slip at the maximum load was 20% greater and 12% less than theoretical for specimens CR1-9 and CR2-9, respectively.

Material properties of masonry can have acceptable coefficients of variation up to 30% therefore the predicted values by ELS are very reasonable. Again, splitting was not as extreme as in the experimental program which could also account for the higher debonding resistance. Further, the running bond capacity exceeds the stack-bonded prism capacity by

8%. Conversely, Petersen et al. (2009) found that the running bond specimens were 8% weaker than the stack-bond specimens in the study. Willis et al. (2009) however observed very little difference in debonding resistance for the running bond specimens.

144 Slip, mm 0 1 2 3 4 5 6 10,000 ELS Model - Running Bonded Specimen 9,000 40 CR2-9 8,000 35 7,000 30 6,000 25 5,000

20 Load, lbs Load, 4,000 kN Load, CR1-9 15 3,000 2,000 10 1,000 5 0 0 0.00 0.05 0.10 0.15 0.20 0.25 Slip, in

Figure 5-8. Comparison of Numerical Analysis to Experimental Data – Running Bonded Specimens

Table 5-3. Comparison of Numerical to Experimental Data – Running Bonded Specimens

Specimen Pmax, lbs (kN) Slip At Pmax, in (mm) PELS/Pexp SlipELS/Slipexp ELS 9169 (40.8) 0.0175 (0.445) - - CR1-9 7573 (33.7) 0.0146 (0.371) 1.21 1.20 CR2-9 8324 (37.0) 0.0199 (0.505) 1.10 0.88

145 5.4.3 Numerical Analysis of Willis et al. (2009) Specimens

In the Willis et al. study (2009) discussed in Chapter 2 of this thesis, pull-test specimens were constructed using ten-core bricks and a CFRP strip that was 0.047 inches

(1.2 mm) in thickness. Similar specimens were also constructed by Daniels et al. (2009) but in this study, three CFRP strips were adhered together to create a 0.145 inch (3.6 mm) thick strip. In both of these studies, nominal brick dimensions were 9.06 x 4.33 x 2.99 inches (230 x 110 x 76 mm) as seen in Figure 5-9, and the mortar joints were approximately 0.39 inches

(10 mm) thick.

Ø0.98" (25 mm) 0.79" (20 mm) 2.99" (76 mm)

0.79" (20 mm) 4.33" (110 mm) 1.65" (42 mm) 9.06" (230 mm)

Figure 5-9. Nominal Brick Dimensions in Willis et al. (2009) and Daniels et al. (2009) Specimens

The prism, comprised of five bricks and four mortar joints, totaled 16.5 inches (420 mm) in height. The ELS model of the stack-bonded pull-test specimen was constructed with a 0.394 inch (10 mm) wide CFRP strip inserted into the specimen. The Willis et al. (2009) running bond specimen used a 0.59 inch (15 mm) wide CFRP strip. Both ELS models are presented in Figure 5-10.

146

Figure 5-10. Daniels et al. (2009) and Willis et al. (2009) Pull-Test Specimens

Material properties used for the stacked prism model are provided in Table 5-4.

Daniels et al. (2009) performed material property tests on the brick and mortar. The lateral modulus of rupture for the brick was experimentally determined to be 434 psi (2.99 MPa).

As previously discussed, the lateral modulus of rupture is related to the tensile strength of the brick by a factor of 1.5. Thus, the resulting tensile strength of the brick was 289 psi (1.99

MPa). The bond wrench tests of the mortar provided a flexural tensile strength of 56.6 psi

(0.39 MPa) however, the flexural tensile strength is typically 1.5 times greater than the tensile strength of the mortar as discussed by Petersen (2009). Therefore, the tensile strength of the mortar is calculated as 37.7 psi (0.26 MPa) after dividing the bond wrench strength by

147 the factor of 1.5. The compressive strengths for both the brick and the mortar were assumed to be ten times that of the tensile strength. Compression tests on brick prisms were conducted to determine the elastic modulus of the brick and the mortar. The shear modulus was based on the assumption made by Petersen (2009) that Poisson’s ratio was 0.2 for both materials.

Although strain gauges were utilized in measuring the elastic modulus of the CFRP in the Daniels et al. study (2009), coupon tests were completed on CFRP strips in the Willis et al. study (2009). Both studies used CFRP with an elastic modulus of 165 GPa and rupture strain of 0.014 according to manufacturer’s data. However, the illis et al. (2009) study experimentally determined the rupture strain to be 0.017 and the elastic modulus to be 162

GPa. Therefore, the FRP properties experimentally determined by Willis et al. (2009) were used. The shear modulus was based on the assumption that the Poisson’s Ratio of the material was 0.3. The tensile strength was calculated based on the material being perfectly linear-elastic. The compressive strength was then assumed to be one-tenth of the tensile strength of the FRP as it was not experimentally determined. The epoxy properties were provided by the manufacturer, for the MBrace Laminate Adhesive. However, the elastic modulus, Poisson’s ratio, and tensile strength of the epoxy were also experimentally determined. The shear modulus was based on the experimentally determined Poisson’s ratio of 0.312 (D. Hale, personal communication, February 19, 2010). These properties for the

FRP and epoxy system were used for both the stacked and running bonded prisms.

148 Table 5-4. Daniels et al. (2009) Material Properties

Property Brick Mortar CFRP Epoxy Elastic Modulus, ksi (GPa) 1,931 (13) 393 (2.7) 23,496 (162) 971 (6.7) Shear Modulus, ksi (GPa) 805 (5.5) 164 (1.1) 9,037 (62) 370 (2.6) Compressive Strength, psi (MPa) 2,890 (20) 377 (2.6) 39,943 (275) 8,702 (60) Tensile Strength, psi (MPa) 289 (2.0) 37.7 (0.26) 399,434 (2,754) 2,012 (14)

Daniels et al. (2009) observed IC debonding as the failure mode for all tested specimens. Cracking began in top brick course and propagated towards the bottom of the prism. Significant cracking was observed in the top courses of bricks with cracking becoming parallel to the FRP strip in the bottom bricks. As seen in Figure 5-11, cracking begins in the top course of brick and propagates towards the bottom as also observed experimentally in this study and the Daniels et al. (2009) study. Furthermore, as the load increases, more cracking is seen in the top two courses of brick which corresponds to the significant amount of cracking in top of the prism observed by Daniels et al. (2009).

149

Figure 5-11. Crack Propagation in Daniels et al. (2009) Stack-Bonded Specimen

150

Initial Cracking at Mortar Joints, Cracking through 1st Course, P = 3,075 lbs (13.7 kN) P = 7,870 lbs (35.0 kN)

Cracking through 2nd Course, Cracking through 3rd Course, P = 9,576 lbs (42.6 kN) P = 10,002 lbs (44.5 kN)

151

Cracking through 4th Course, Cracking through 5th Course, P = 10,277 lbs (45.7 kN) P = 12,013 lbs (53.4 kN)

At 1st Peak, At Ultimate Load, P = 12,967 lbs (57.7 kN) P = 13,090 lbs (58.2 kN)

152

Post-Peak, At Final Condition, P = 10,878 lbs (48.4 kN) P = 164 lbs (0.73 kN)

The load versus slip is plotted in Figure 5-12. The ultimate debonding resistance observed through the numerical analysis was 13,090 lbs (58.2 kN) and the slip at this load was 0.0951 inches (2.4 mm). In this case, the experimental debonding resistance was actually greater than that which was determined through the numerical analysis as seen in

Table 5-5. However, for specimens PT1 and PT2, the predicted slip was less than what was experimentally observed. On average, the percent difference between the predicted and experimental IC debonding load was 10%, which is acceptable considering acceptable coefficient of variations of masonry tests are up to 30%. However, the average difference between the theoretical and experimental slip was 22%. The stiffness of the ELS model and experimental pull-tests also vary in that the model initially softens prior to stiffening. Upon further analysis of the ELS output, cracking is visible at the brick to mortar interface at the

153 bottom of each individual brick course due to the extremely weak tensile strength of the mortar. These cracks initially open prior to cracking propagating in the top brick course.

Slip, mm 0 1 2 3 4 5 6 16,000 70 PT1 14,000 PT3 60 PT2 12,000 50 10,000 40 8,000

30 Load, lbs Load, 6,000 ELS Model kN Load, 20 4,000

2,000 10

0 0 0.00 0.05 0.10 0.15 0.20 0.25 Slip, in

Figure 5-12. Comparison of Numerical Analysis to Experimental Data – Daniels et al. (2009) Stack-Bonded Specimens

154 Table 5-5. Comparison of Numerical to Experimental Data – Daniels et al. (2009) Stack-Bonded Specimens

Specimen Pmax, lbs (kN) Slip At Pmax, in (mm) PELS/Pexp SlipELS/Slipexp ELS 13,090 (58.2) 0.095 (2.4) - - PT1 15,758 (70.1) 0.118 (3.0) 0.83 0.80 PT2 13,713 (61.0) 0.118 (3.0) 0.95 0.80 PT3 14,320 (63.7) 0.075 (1.9) 0.91 1.26

Running bonded specimens were not tested by Daniels et al. (2009). Therefore, comparisons were made using specimens from Willis et al. (2009) although estimations had to be made for the material properties that were not provided. The lateral modulus of rupture for bricks used in the running bond specimens was 551 psi (3.80 MPa) which is reduced by a factor of 1.5 to obtain a tensile strength of 368 psi (2.53 MPa). The compressive strength was assumed to be ten times greater than the tensile strength of the brick. The bond wrench test provided a flexural tensile strength of the mortar of 88 psi (0.61 MPa) which is also reduced by a factor of 1.5 to obtain a tensile strength of 59 psi (0.41 MPa). Again, the compressive strength is assumed to be ten times greater than the tensile strength of the mortar. However, the elastic moduli for the brick and mortar were not provided.

Furthermore, using Equation 5-8 and that the elastic modulus of the mortar is 200 times greater than the compressive strength resulted in an elastic modulus of mortar and brick that were lower than those presented in Table 5-4.

Typically, as the strength increases, as seen in Table 5-6 for the mortar and brick, the elastic modulus increases. It was assumed that the same bricks were used for both the Willis et al. (2009) and the Daniels et al. (2009) study. Therefore, the elastic moduli for both the

155 mortar and brick were proportionally increased from the compressive strength of the material and elastic modulus found by Daniels et al. (2009).

Table 5-6. Willis et al. (2009) Material Properties

Property Brick Mortar CFRP Epoxy Elastic Modulus, ksi (GPa) 2,292 (16) 615 (4.2) 23,496 (162) 971 (6.7) Shear Modulus, ksi (GPa) 955 (6.6) 256 (1.8) 9,037 (62) 370 (2.6) Compressive Strength, psi (MPa) 3,430 (24) 590 (4.1) 39,943 (275) 8,702 (60) Tensile Strength, psi (MPa) 343 (2.4) 59 (0.41) 399,434 (2,754) 2,012 (14)

Similar to observations made by Daniels et al. (2009), Willis et al. (2009) also observed cracking that propagated from the loaded end of the specimen to the bottom of the prism in the masonry substrate. However, rupture of the FRP in the area above the prism was also observed. This failure mode was mirrored in the numerical model presented in

Figure 5-13. Cracking propagated from the loaded end towards the bottom of the prism.

However, eventually the rupture strain of the FRP was reached thereby preventing complete debonding of the FRP.

156

Figure 5-13. Crack Propagation in Willis et al. (2009) Specimen

157

Cracking through 1st Course, Cracking through 2nd Course, P = 6,326 lbs (28.1 kN) P = 9,489 lbs (42.2 kN)

At 1st Peak, Cracking in All Courses, P = 9,836 lbs (43.8 kN) P = 11,070 lbs (49.2 kN)

158

Rupture of FRP, At Final Condition, P = 11,199 lbs (49.8 kN) P = 2.32 lbs (0.01 kN)

The load capacities for running bond specimens, HO1.5-4-15-NSG and HO1.5-4-15-

1/4, are depicted in Figure 5-14 as load-slip data for these specimens were not provided.

Specimen HO1.5-4-15-NSG failed by rupture at a load of 11,397 lbs (50.7 kN) whereas

HO1.5-4-15-1/4 failed by IC debonding at a load of 9,891 lbs (44.0 kN). From the numerical analysis, the ultimate load at which rupture occurred was 11,200 lbs (50.2 kN) and the slip at this load was 0.0359 inches (0.91 mm). It is significant that the ELS model closely mirrors the behavior of the specimen that has no strain gauges attached to the FRP strip. Specimen

HO1.5-4-15-NSG reaches a higher load that results in rupture of the FRP. The specimen with strain gauges did not reach such a high load due to the loss of the bonded area due to strain gauges. Therefore, it is expected that the numerical analysis would follow most closely the behavior for specimens without strain gauges as these are not modeled. From

159 Table 5-7, it can be seen that specimen HO1.5-4-15-NSG is within 2% of the peak load that

ELS predicts however, due to the rupture failure of the FRP, this is in reality a function of the rupture property of the material.

Slip, mm 0 0.5 1 1.5 2 2.5 3 14,000 60 HO1.5-4-15-NSG HO1.5-4-15-1/4 12,000 50 10,000 40 8,000 30

6,000

Load, lbs Load, Load, kN Load, ELS Model 20 4,000

2,000 10

0 0 0.00 0.01 0.02 0.03 0.04 0.05 Slip, in

Figure 5-14. Comparison of Numerical Analysis to Experimental Data – Willis et al. (2009) Running Bonded Specimens

160 Table 5-7. Comparison of Numerical to Experimental Data – Running Bonded Specimens

Specimen Pmax, lbs (kN) Slip At Pmax, in (mm) PELS/Pexp SlipELS/Slipexp ELS 11,200 (50.2) 0.0359 (0.91) - - HO1.5-4-15-NSG 11,397 (50.7) N/A 0.98 N/A HO1.5-4-15-1/4 9,891 (44.0) N/A 1.13 N/A

5.4.4 Numerical Analysis of Petersen et al. (2009) Specimens

The third model is based on the pull-test specimens tested by Petersen et al. (2009) as discussed in Chapter 2 of this thesis. Specimens consisted of four solid pavers and FRP strips that were 0.11 inches (2.8 mm) in thickness and 0.59 inches (15 mm) in width. Each brick was of nominal dimensions 9.06 x 4.33 x 2.99 inches (230 x 110 x 76 mm) as seen in

Figure 5-15, and the mortar joints were approximately 0.39 inches (10 mm) thick. The pull- test specimen as modeled in ELS is depicted in Figure 5-16.

2.99" (76 mm)

4.33" (110 mm)

9.06" (230 mm)

Figure 5-15. Brick Dimensions Used in Petersen et al. (2009) Study

161

Figure 5-16. Petersen et al. (2009) Pull-Test Specimens

Material properties used for the model are provided in Table 5-8. The tensile strength of the brick was calculated from the modulus of rupture, tested in the same manner as the set- up described in Chapter 3 of this thesis. The lateral modulus of rupture provided from

Petersen et al. (2009) was 518 psi (3.57 MPa). The elastic modulus for the brick unit was determined from compression tests on seven brick high prisms similar to those tested in

Chapter 3 of this thesis. The elastic moduli of the mortar and the masonry prism were experimentally determined. The elastic modulus of the brick unit was then calculated based on averaging the elastic moduli of the mortar and the prism and considering strain compatibilities between the prism and the mortar. Poisson’s ratio was set as 0.2 for calculation of the shear modulus as assumed by Petersen (2009). The compressive strength

162 of the brick unit was provided by the manufacturer. The elastic modulus for the mortar was experimentally determined through the prism compressive tests. Again, the Poisson’s ratio was assumed 0.2 for calculation of the shear modulus by Petersen (2009). The bond strength used in prisms for which the mortar’s elastic modulus was determined was 177 psi (1.22

MPa). The bond strength is related to the tensile strength of the mortar by a factor of 1.5; therefore, the tensile strength of the mortar was taken as 118 psi (0.81 MPa). It was assumed that the mortar compressive strength was ten times that of its tensile strength.

The elastic modulus of the CFRP was based on data resulting directly from the pull- test specimens while the shear modulus was calculated based on the assumption made by

Petersen (2009) that the Poisson’s ratio was 0.3. A rupture strain of .012, based on manufacturer’s data, and the knowledge that the material is perfectly linear-elastic was used to calculate the ultimate tensile strength of the CFRP. The compressive strength was then assumed to be one-tenth of the tensile strength. Petersen (2009) did not experimentally determine material properties for the epoxy. Manufacturer’s data was provided for the flexural strength and compressive strength of the epoxy, which was 4,351 psi (30 MPa) and

8,702 psi (60 MPa), respectively. The same commercially available epoxy was used by both

Petersen et al. (2009) and Willis et al. (2009). Therefore, the epoxy material properties utilized for the Willis et al. (2009) model were also used for the Petersen et al. (2009) model.

163 Table 5-8. Petersen (2009) Material Properties

Property Brick Mortar CFRP Epoxy Elastic Modulus, ksi (GPa) 4,002 (28) 737 (5.1) 30,026 (207) 971 (6.7) Shear Modulus, ksi (GPa) 1,667 (11) 307 (2.1) 11,548 (80) 370 (2.6) Compressive Strength, psi (MPa) 5,076 (35) 1,180 (8.1) 36,027 (248) 8,702 (60) Tensile Strength, psi (MPa) 345 (2.4) 118 (0.81) 360,274 (2484) 2,012 (14)

Figure 5-17 depicts the cracking that initiated in the top course along the brick elements located next to the FRP strip. Cracking then propagates into the subsequent courses along the FRP strip as the load increases. Cracking then continues along all brick courses post- peak until no further load is applied to the FRP strip. It is also important to note that cracking propagates towards the back of the prism as was observed in this study. Significant cracking along the front face and splitting occurred following the peak load, which was also observed in the Petersen et al. (2009) study.

164

Figure 5-17. Cracking Behavior of Petersen et al. (2009) Specimen

165

Cracking through 1st Course, Cracking through 2nd Course, P = 9,678 lbs (43.1 kN) P = 11,233 lbs (50.0 kN)

Cracking through 3rd Course, Cracking through 4th Course, P = 12,611 lbs (56.1 kN) P = 14,312 lbs (63.7 kN)

166

Splitting of Bottom Courses, At Peak Load, P = 16,315 lbs (72.6 kN) P = 20,823 (92.6 kN)

Post-Peak, At Final Condition, P = 16,466 lbs (73.2 kN) P = 124.2 lbs (0.55 kN)

167 Some material properties were estimated for this model, as they were not experimentally determined by Petersen (2009). Therefore, some variability between the peak load and slip at the peak load should be expected. As seen in Figure 5-18, the load-slip curves are similar although the ELS overpredicts the peak load at a smaller slip. The load- slip data was not provided for specimens without strain gauges. Table 5-9 presents the comparison between the experimental and numerical data. ELS models the debonding resistance of the Petersen specimens on average, within 18% of the experimental data.

Slip, mm 0 1 2 3 4 5 6

20,000 90 80 ELS Model 15,000 70 60 S1-A-SG 50

10,000 Load, lbs Load, 40 kN Load, 30 5,000 20 10 0 0 0.00 0.05 0.10 0.15 0.20 0.25 Slip, in

Figure 5-18. Load-Slip Curves – Petersen (2009) Stack-Bonded Specimen

168 Table 5-9. Comparison of Numerical to Experimental Data – Stack Bonded Specimens

Specimen Pmax, lbs (kN) Slip At Pmax, in (mm) PELS/Pexp SlipELS/Slipexp ELS 20,823 (92.6) 0.039 (0.99) - - S1-A-NG-1 18,760 (83.45) N/A 1.11 N/A S1-A-NG-2 15,981 (71.09) N/A 1.30 N/A S1-A-SG 18,317 (81.5) 0.051 (1.3) 1.14 0.76

Figure 5-19 depicts the cracking that initiated in the top course along the brick elements located next to the FRP strip. Cracking then propagates into the subsequent courses along the FRP strip as the load increases. Cracking then continues along all brick courses post-peak until no further load is applied to the FRP strip. It is also important to note that cracking propagates towards the back of the prism as was observed in this study. Significant cracking along the front face and splitting occurred following the peak load, which was also observed in the Petersen (2009) study.

169

Figure 5-19. Cracking Behavior of Petersen (2009) Running Bond Specimen

170

Cracking through 1st Course, Cracking through 2nd Course, P = 9,613 lbs (42.6 kN) P = 11,292 lbs (50.2 kN)

Cracking through 3rd Course, Cracking through 4th Course, P = 12,454 lbs (55.4 kN) P = 13,853 lbs (61.6 kN)

171

At Peak Load, At Final Condition, P = 14,469 lbs (64.4 kN) P = 7,867 (35.0 kN)

As seen in Figure 5-20, the initial stiffness of the load-slip curve is similar. Specimen

S1-C-SG is an anomaly in that it had a significantly higher debonding resistance than specimens S1-C-NG-1 and S1-C-NG-2 as shown in Table 5-10. Petersen et al. (2009) did not provide load-slip data for these specimens; therefore, only the load capacities are provided.

172 Slip, mm 0 2 4 6 20,000 S1-C-SG 18,000 80 S1-C-NG-2 16,000 70

14,000 60 12,000 S1-C-NG-1 50 10,000

ELS Model 40 Load, lbs Load, 8,000 kN Load, 30 6,000 4,000 20 2,000 10 0 0 0.00 0.05 0.10 0.15 0.20 0.25

Slip, in Figure 5-20. Load-Slip Curves – Petersen (2009) Running Bonded Specimens

Table 5-10. Comparison of Numerical to Experimental Data – Running Bonded Specimens

Specimen Pmax, lbs (kN) Slip At Pmax, in (mm) PELS/Pexp SlipELS/Slipexp ELS 14,469 (64.4) 0.042 (1.07) - - S1-C-NG-1 14,365 (63.9) N/A 1.01 N/A S1-C-NG-2 15,601 (69.4) N/A 0.93 N/A S1-C-SG 18,996 (84.5) 0.053 (1.35) 0.76 0.79

173 5.5 Sensitivity Analysis

As was discussed in Chapter 4 of this thesis, the Willis Masonry Model overpredicts the load at which IC debonding occurs for pull-tests tested in this study. Many variables could account for the reduction in strength seen by prisms in this study such as the material properties of the epoxy or mortar both of which are not accounted for in the Willis Masonry

Model. As observed, the brick geometry and/or material properties can lead to premature failure by splitting through the prism. For comparison, material properties of the mortar and epoxy were altered to observe the difference in the response.

Figure 5-21 presents the load-slip curve for the stack-bonded specimen tested in this study. The strength and stiffness of the mortar were decreased by 5, 10, 20, 30, 40, and 50% to observe changes in the response. The load, Po, in Table 5-11 describes the peak load found in the ELS model of the stack-bonded specimen from this study. Load, Paltered, is the debonding resistance found when the material properties of the stack-bonded specimen were changed. Decreasing the strength of the mortar decreases the debonding resistance of the

FRP-reinforced prism by twice the percentage that the mortar was decreased. For example, decreasing the mortar strength and stiffness by 5% resulted in a 10% reduction in the debonding resistance. However, this trend reaches an asymptote where an additional decrease in the mortar strength no longer results in a significant decrease in the debonding resistance.

The literature review suggests that epoxy properties such as the stiffness and strength can decrease up to 50% when temperatures near the glass transition temperature (Borchert &

Zilch, 2005). The initial stiffness changed significantly with the change in the material

174 properties as seen in Figure 5-22. When the strength and stiffness of the epoxy decreased by

10, 25, and 50%, the debonding resistance decreased by 11, 13, and 10% respectively as shown in Table 5-12.

The Willis Masonry Model overpredicted the IC debonding load of the pull-test specimens by 34% to 175%. Therefore, the use of a different epoxy or mortar was not primarily responsible for the vast difference in the theoretical and experimental debonding resistances. Thus, it can be inferred that splitting through of the prisms was primarily responsible for reducing the debonding resistance of the masonry prisms and causing premature failure.

175 Slip, mm 0 0.5 1 1.5 2 2.5 3 3.5 9,000 40

8,000 5% 35 10% 7,000 30 20% 6,000 30% 25 5,000 Current Study 20

4,000 Load, lbs Load, 15 kN Load, 3,000 40% 2,000 10 1,000 50% 5 0 0 0.00 0.02 0.04 0.06 0.08 0.10 0.12 Slip, in

Figure 5-21. Load-Slip Curves with Decrease in Mortar Strength and Stiffness

Table 5-11. Comparison of Stack-Bonded Specimens through Sensitivity Analysis of Mortar

Specimen Pmax, lbs (kN) Slip At Pmax, in (mm) Po/Paltered Slipo/Slipaltered Current Study 8,486 (37.7) 0.021 (0.521) - - 5% Decrease 7746 (34.7) 0.017 (0.432) 1.10 1.19 10% Decrease 7089 (31.8) 0.015 (0.381) 1.20 1.33 20% Decrease 7154 (32.1) 0.016 (0.406) 1.19 1.25 30% Decrease 6801 (30.5) 0.016 (0.406) 1.25 1.31 40% Decrease 6330 (28.4) 0.015 (0.381) 1.34 1.41 50% Decrease 6336 (28.4) 0.015 (0.381) 1.34 1.34

176 Slip, mm 0 0.5 1 1.5 2 2.5 3 3.5 9,000 35 10% 8,000 30 7,000 25 6,000 25% 5,000 20 50%

4,000 15

Load, lbs Load, Load, kN Load, 3,000 Current 10 Study 2,000 5 1,000 0 0 0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14 Slip, in

Figure 5-22. Load-Slip Curves with Decrease in Epoxy Strength and Stiffness

Table 5-12. Comparison of Stack-Bonded Specimens through Sensitivity Analysis of Epoxy

Specimen Pmax, lbs (kN) Slip At Pmax, in (mm) Po/Paltered Slipo/Slipaltered Current Study 8,486 (37.7) 0.021 (0.521) - - 10% Decrease 7621 (34.1) 0.018 (0.457) 1.11 1.17 25% Decrease 7497 (33.6) 0.025 (0.635) 1.13 0.81 50% Decrease 7693 (34.5) 0.030 (0.762) 1.10 0.67

177 5.6 Conclusions

Extreme Loading® for Structures has been proven an effective tool at predicting the IC debonding load of masonry pull-tests. Although some material properties were estimated as they were not provided, the predictions for the debonding resistance are within 20% of the experimental loads, which is reasonable considering the variability of masonry properties.

However, ELS overpredicts the debonding resistance for a majority of the specimens. Not only can this be caused by assuming the properties for materials but also it can stem from the idealization of the masonry prism. Due to the non-homogenous nature of a masonry prism, there is often a high variability between specimens. Therefore, caution should be taken when using ELS as a design tool for FRP-reinforced masonry structures as they will most likely reach lesser debonding resistances than modeled. The same caution should be used when using a FEM modeling tool. However, with the AEM, a bond-slip model was not necessary to include in the model and ELS still provided good results. Thus, the need for experimentally determining the bond-slip model prior to running a numerical analysis is eliminated. Just by using the material properties, the debonding resistance of the FRP- reinforced masonry prism can be found with reasonable accuracy.

As previously discussed, splitting was thought to be the main cause for the reduction in debonding resistance compared to the Willis Masonry Model. Through the use of ELS, it was found that changing the epoxy or mortar properties did not have a significant impact on the debonding resistance as to create the large gap between the theoretical and experimental

IC debonding loads. Therefore, the splitting through of the prisms observed in this study that

178 was not seen in the Willis et al. (2009) study was primarily responsible for the difference in the theoretical and experimental debonding resistances.

The geometry of the brick units used in the pull-tests influence the splitting behavior of the prisms. Comparing the geometric properties of the bricks such as the section modulus can help predict the susceptibility of the prism to splitting. Of the bricks used in this study and by Willis et al. (2009) and Petersen (2009), it was found that the section modulus for the solid bricks, which was 9.4 in3 (154 cm3) is greatest. Brick units in the Willis et al. (2009) study resulted in a section modulus of 7.0 in3 (115 cm3) that was 25% smaller than the solid units used by Petersen et al. The smallest section modulus of 4.2 in3 (69 cm3) was associated with the brick units used in this research as it was 55% smaller than the section modulus of the solid brick units. Decreasing the section modulus of the cross-section will result in higher tensile stresses for the brick unit due to dilation of the debonding crack, increasing the likelihood of splitting cracks. Petersen et al. (2009) suggested that the splitting of the prism was also due to cracking induced by the grooves cut for the FRP insertion. Because the cracking was in line with the FRP, splitting was more likely to occur when the groove was also in line with the cores as there was less resistance to crack propagation. Petersen et al.

(2009) suggests that the FRP depth be kept to a minimum, particularly in walls with cored brick as these have a greater potential to split.

Currently, the Willis Masonry Model accounts for the tensile strength of the brick yet it does not take into consideration the material properties of the mortar and adhesive system used. The sensitivity analysis has shown that these properties do affect the resulting debonding resistance. Therefore, their inclusion in the Willis Masonry Model is necessary as

179 using a weaker mortar or epoxy in the field could adversely affect the strength or behavior of the FRP-reinforced structure.

180 CHAPTER 6 – CONCLUSIONS & RECOMMENDATIONS

6.1 Conclusions

The objective of this research was to determine the variation in debonding behavior of

FRP-reinforced masonry prisms following environmental conditioning. It was found that the environmental conditioning had a significant impact on the failure mechanism of the FRP- reinforced prism in a relatively short period of exposure time. The following chapter will present the major conclusions of this research as well as discuss future studies that expand on the findings illustrated in this thesis.

6.1.1 Splitting of Control Specimens

During the pull-tests, it was observed that the control prisms split through the entire width of the prism as seen in Figure 6-1. As cracking propagated down the front face of the prism, the resulting rough surfaces of the brick moving against each other caused lateral dilation of the prism, thus inducing splitting through the specimen. Further, the large area of the cores and proximity to the NSM FRP strip encouraged splitting across the top brick as it became easier for the prism to split due to the minimum thickness of the brick causing higher stress concentrations. The resulting experimental IC debonding loads were on average 86% lower than the IC debonding load prediction given by the Willis Masonry Model due to the splitting of the prism as this can be considered a premature failure. This was further proven through the use of ELS as changing the material properties of the epoxy and mortar, both of which are not accounted for in the Willis Masonry Model, did not significantly alter the

181 debonding resistance of the FRP-reinforced specimen. Therefore, the splitting was the cause for the lesser debonding resistances than predicted by Willis et al. (2009).

Figure 6-1. Splitting Through of Prisms

From observations made during the experimental program, there is now a need to determine all real-life conditions where splitting may occur. In these conditions, the structure may fail prematurely as the masonry is also splitting as the FRP debonds from the structure. The typical cracking pattern for an URM wall is illustrated in Figure 6-2. With a similar cracking pattern and vertical reinforcement on the tension side of the wall, a situation may present itself where the bricks split. When vertical FRP reinforcement crosses a crack as highlighted in Figure 6-2, a situation similar to a pull-test can occur in which the FRP is loaded in tension. As IC debonding occurs, the widening of the debonding crack may cause splitting of the brick. However, in full-size structures, lateral support will be offered by the

182 surrounding wall, which may delay or prevent splitting. Further, analysis must be done on the geometric properties of the brick units used in the wall, as the section modulus will affect the onset of splitting. Continued research on this topic will ensure that premature failure due to splitting of the brick does not occur.

NSM FRP Strip

Figure 6-2. Cracking Pattern of URM Wall (Wylie, 2009)

6.1.2 Naturally Conditioned and Shear Failure in Adhesive

Following conditioning of specimens subjected to daily variations in temperature and precipitation, it was observed that there was a significant impact on these prisms in a relatively short period of exposure time. Due to excessive temperatures on the roof where the prisms were stored, the glass transition temperature of the epoxy was most likely reached as the texture of the epoxy changed to being rubber-like. Therefore, a different failure mechanism was observed instead of the splitting observed for the control specimens. Shear

183 failure through the adhesive was observed as the material properties of the epoxy had changed. The naturally conditioned specimens exhibited very little cracking prior to failure.

Because the debonding failure plane remained smooth, there was no dilation of the brick prism because of cracking. Consequently, splitting of the prism was not observed.

Typically, a shear failure through the adhesive layer would be weaker than the debonding resistance associated with IC debonding as the literature review has shown that environmental conditioning weakens the FRP system. However, the debonding resistance observed for prisms that failed by IC debonding was compromised by the splitting through of the prism. Therefore, this IC debonding load was lower than expected as proven by the comparison with the Willis Masonry Model. Had these prisms not split, the debonding resistance would be much higher than that associated with the naturally conditioned specimens. However, the naturally conditioned specimens were also well below the predicted IC debonding load from the Willis Masonry Model demonstrating that this too was an undesirable failure mode.

6.1.3 Numerical Simulation

Through the use of the computer program Extreme Loading® for Structures that utilizes the Applied Element Method, it was determined that pull-test specimens of all types of construction or material properties can be easily modeled. As exhibited through the Willis et al. (2009), Petersen et al. (2009), and this study’s models, ELS can accurately model the load-slip behavior of a variety of pull-test specimens without the need for a bond-slip model.

As more FRP products become commercially available, the use of a modeling program will be necessary in determining the most efficient and safest combination of

184 materials for the structural design. Through the use of ELS, it was discovered that the debonding resistance of an FRP strengthening system is sensitive to the epoxy and mortar material properties. In an effort to avoid spending time and money on performing pull-tests on specimens with varying epoxies, ELS can be used to quickly observe the differences in the debonding loads. Identifying the expected debonding resistance for a selected combination of materials will be beneficial in designing or analyzing structures reinforced with FRP. It will also enable the designer to select from a vast array of brick types that may be locally available as the location and amount of cores will also affect the debonding resistance.

6.2 Recommendations and Future Work

As the use of FRP grows in the civil engineering field, it will be necessary to fully understand the long-term behavior of an FRP strengthening system. To do so, further research must be completed in the durability aspects of the system. Subjecting the pull-test specimens to a more regimented conditioning treatment would supplement the results found in this study by testing the extreme limits of the system. Parameters to study include additional freeze/thaw or wet/dry cycles to specifically isolate the causes for the shear failure through the adhesive in the naturally conditioned specimens. Furthermore, a test period of longer than nine months should be used to fully understand the deterioration of the bond in the FRP strengthening system.

Pull-tests performed on larger prisms could possibly prevent the splitting through of the specimens observed in this study. Instead of using a stacked or running bond specimen, a

185 larger wallette could be tested. In practice, the support from the surrounding bricks in the wall may prevent the splitting seen in the pull-tests in this study. Additionally, minimum

FRP widths should be utilized, particularly in cored brick, to minimize the groove depths that could induce cracking. Further, the same stacked and running bond specimens could be tested with the addition of adding compression to the sides similar to the Petersen et al.

(2009) study in an effort to eliminate splitting.

ELS can also be used to model FRP-reinforced walls in an effort to match the behavior observed in published experimental programs. Being able to use a modeling program to predict the behavior of an FRP-reinforced masonry wall may help advance the use of FRP in the industry for design work. Further, it was observed through the use of ELS that the material properties of epoxy and mortar altered the debonding resistance of the strengthened prism. It will be necessary to incorporate these material properties in debonding load prediction models, as they are not currently included in the Willis Masonry Model.

Additionally, the effects of the brick geometry should be included in future work as splitting was seen in this study and in the solid pavers tested by Petersen et al. (2009) but not in the ten-core specimens tested by Willis et al. (2009).

6.3 Summary

Previously, there has been very little research published on the use of pull-tests to determine the changes in the debonding behavior of an FRP-reinforced masonry prism.

Through this study, the results have shown that there has been a significant change in the debonding failure of a naturally conditioned specimen due to exposure to moisture and high

186 temperatures. Discomforting is that these specimens were subjected to the open air for a relatively short period compared to the service life of a full-scale structure yet the changes were considerable. A longer conditioning treatment is necessary to fully understand the long-term changes on the FRP-reinforced structure. Although this study observed a larger debonding resistance for the naturally conditioned specimens than the controls, this would not be the case in a full-scale structure. The premature failure of the control specimens by splitting of the prism would be prevented by the lateral restraint provided by the surround bricks. Therefore, the shear failure through the adhesive observed in this study would be the weakest failure mechanism. Without a full understanding of the durability of the system, many FRP-reinforced structures may fail prematurely following a short period of weathering after being put into service.

The properties for the brick, epoxy, or FRP will vary depending on the manufacturer providing the materials. Therefore, a numerical analysis program will become increasingly important as it will be necessary to confirm the debonding resistance of the design with the selected materials prior to installation. It has been well researched that NSM FRP provides a significant strength gain for URM walls yet more research must be completed to ensure the system resists weathering in the field. Currently, ACI 440.7R-10, “Guide for the Design and

Construction of Externally Bonded Fiber-Reinforced Polymer Systems for Strengthening

Unreinforced Masonry Structures” requires a reduction factor of 0.85 to be used in design work for structures placed in harsh environments. However, upon comparison with the

Willis Masonry Model, the predicted debonding resistance was on average, 77% greater than the experimental. Therefore, a reduction of only 15% as ACI 440.7R-10 suggests, may not

187 be sufficiently conservative. It is recommended to complete a more thorough investigation into the long-term behavior of an FRP-reinforced masonry system particularly through the use of a pull-test, as this topic is still relatively unstudied.

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